ASM Principles of Soldering

Principles of

Soldering

Giles Humpston

David M. Jacobson

Materials Park, Ohio 44073-0002www.asminternational.org

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Library of Congress Cataloging-in-Publication DataHumpston, Giles.

Principles of soldering / Giles Humpston, David M. Jacobson.p. cm.

Includes bibliographical references and index.ISBN 0-87170-792-6

1. Solder and soldering. 2. Brazing. I. Jacobson, David M. II. Title.

TS610.H84 2004671.5’6—dc22 2003058379

SAN: 204-7586

ASM International®Materials Park, OH 44073-0002


Printed in the United States of America



Contents

Preface .............................................................................................................................................. vii

About the Authors............................................................................................................................ ix

History ................................................................................................................................................ x

Chapter 1: Introduction................................................................................................................... 11.1 Joining Methods .................................................................................................................... 1

1.1.1 Mechanical Fastening ................................................................................................. 11.1.2 Adhesive Bonding ...................................................................................................... 21.1.3 Soldering and Brazing ................................................................................................ 31.1.4 Welding ....................................................................................................................... 41.1.5 Solid-State Joining ..................................................................................................... 41.1.6 Comparison between Solders and Brazes ................................................................. 51.1.7 Pressure Welding and Diffusion Bonding ................................................................. 8

1.1.7.1 Pressure Welding .................................................................................................. 91.1.7.2 Diffusion Bonding ................................................................................................ 9

1.2 Key Parameters of Soldering ............................................................................................. 121.2.1 Surface Energy and Surface Tension ....................................................................... 121.2.2 Wetting and Contact Angle ...................................................................................... 131.2.3 Fluid Flow ................................................................................................................ 181.2.4 Filler Spreading Characteristics ............................................................................... 191.2.5 Surface Roughness of Components ......................................................................... 221.2.6 Dissolution of Parent Materials and Intermetallic Growth ..................................... 241.2.7 Significance of the Joint Gap ................................................................................... 251.2.8 The Strength of Metals ............................................................................................ 27

1.3 The Design and Application of Soldering Processes ........................................................ 281.3.1 Functional Requirements and Design Criteria ........................................................ 28

1.3.1.1 Metallurgical Stability ........................................................................................ 291.3.1.2 Mechanical Integrity ........................................................................................... 291.3.1.3 Environmental Durability ................................................................................... 291.3.1.4 Electrical and Thermal Conductivity ................................................................. 30

1.3.2 Processing Aspects ................................................................................................... 301.3.2.1 Jigging of the Components ................................................................................ 301.3.2.2 Form of the Filler Metal .................................................................................... 311.3.2.3 Heating Methods ................................................................................................ 331.3.2.4 Temperature Measurement ................................................................................. 341.3.2.5 Joining Atmosphere ............................................................................................ 351.3.2.6 Coatings Applied to Surfaces of Components .................................................. 371.3.2.7 Cleaning Treatments ........................................................................................... 371.3.2.8 Heat Treatments Prior to Joining ....................................................................... 371.3.2.9 Heating Cycle of the Joining Operation ............................................................ 381.3.2.10 Postjoining Treatments ....................................................................................... 391.3.2.11 Postjoining Cleaning .......................................................................................... 391.3.2.12 Statistical Process Control ................................................................................. 42

1.3.3 Health, Safety, and Environmental Aspects of Soldering ....................................... 42Chapter 1: Appendices ..................................................................................................................43

A1.1 Solid-State Joining with Gold, Indium, and Solder Constituents ........................... 43A1.2 Relationship among Spread Ratio, Spread Factor, and Contact Angle of

Droplets ................................................................................................................. 44

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Chapter 2: Solders and Their Metallurgy................................................................................... 492.1 Survey of Solder Alloy Systems ........................................................................................ 51

2.1.1 Lead-Tin Solders ...................................................................................................... 562.1.2 Other Tin-Base Solders ............................................................................................ 582.1.3 Zinc-Bearing Solders ................................................................................................ 602.1.4 Gold-Bearing Solders ............................................................................................... 642.1.5 High-Lead Solders .................................................................................................... 722.1.6 Indium Solders ......................................................................................................... 73

2.2 Effect of Metallic Impurities .............................................................................................. 752.3 Application of Phase Diagrams to Soldering .................................................................... 77

2.3.1 Examples Drawn from Binary Alloy Systems ........................................................ 792.3.2 Examples Drawn from Ternary Alloy Systems ....................................................... 832.3.3 Complexities Presented by Higher-Order and Nonmetallic

Systems ................................................................................................................. 922.4 Depressing the Melting Point of Solders by Eutectic Alloying ........................................ 93

2.4.1 Liquid Alloys Based on Gallium ............................................................................. 932.4.2 Cadmium-Base Solders ............................................................................................ 932.4.3 General Features ....................................................................................................... 932.4.4 Implications for Lead-Free Solders ......................................................................... 95

Chapter 2: Appendices ..................................................................................................................96A2.1 Conversion between Weight and Atomic Fraction of

Constituents of Alloys .......................................................................................... 96A2.2 Theoretical Modeling of Eutectic Alloying ............................................................. 97

Chapter 3: The Joining Environment........................................................................................ 1033.1 Joining Atmospheres ......................................................................................................... 103

3.1.1 Atmospheres and Reduction of Oxide Films ........................................................ 1053.1.2 Thermodynamic Aspects of Oxide Reduction ....................................................... 1063.1.3 Practical Application of the Ellingham Diagram .................................................. 107

3.1.3.1 Soldering in Inert Atmospheres and Vacuum .................................................. 1073.1.3.2 Soldering in Reducing Atmospheres ............................................................... 1093.1.3.3 Alternative Atmospheres for Oxide Reduction ................................................ 111

3.1.4 Forming Gas as an Atmosphere for Soldering ...................................................... 1113.2 Chemical Fluxes for Soldering ......................................................................................... 111

3.2.1 Fluxes for Tin-Base Solders ................................................................................... 1163.2.1.1 Soldering Fluxes That Require Cleaning ......................................................... 1163.2.1.2 No-Clean Soldering Fluxes .............................................................................. 1183.2.1.3 Measure of Cleaning Effectiveness: The Surface Insulation

Resistance (SIR) Test ................................................................................... 1193.2.2 Fluxes for “Unsolderable” Metals ......................................................................... 120

3.2.2.1 Aluminum Soldering Fluxes ............................................................................ 1213.2.2.2 Stainless Steel Soldering Fluxes ...................................................................... 1223.2.2.3 Magnesium Soldering Flux .............................................................................. 122

3.2.3 High-Temperature Fluxes ....................................................................................... 1223.3 Fluxless Soldering ............................................................................................................ 123

3.3.1 Oxide Formation and Removal .............................................................................. 1243.3.2 Self-Dissolution of Solder Oxides ......................................................................... 1253.3.3 Reduction of Solder Oxides by Hydrogen ............................................................ 1263.3.4 Reduction of Solder Oxides by Atomic Hydrogen ............................................... 1273.3.5 Mechanical Removal of Oxides (Ultrasonic Soldering) ....................................... 1283.3.6 Reactive Gas Atmospheres for Reduction of Oxides ........................................... 1303.3.7 Surface Conditioning Processes ............................................................................. 1313.3.8 Fluxless Soldering Processes Considerations ........................................................ 132

3.3.8.1 Solderable Component Surfaces ...................................................................... 1333.3.8.2 Preform Geometry ............................................................................................ 133

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3.3.8.3 Mechanically Enhanced Solder Flow .............................................................. 1343.3.8.4 Metallurgically Enhanced Solder Flow ........................................................... 134

3.3.9 Example of a Fluxless Soldering Process Using In-48SnSolder .................................................................................................................. 135

3.3.10 Fluxless Soldering of Aluminum ........................................................................... 136Chapter 3: Appendix ...................................................................................................................137

A3.1 Thermodynamic Equilibrium and the Boundary Conditions forSpontaneous Chemical Reaction ........................................................................ 137

Chapter 4: The Role of Materials in Defining Process Constraints ...................................... 1454.1 Metallurgical Constraints and Solutions .......................................................................... 147

4.1.1 Wetting of Metals by Solders ................................................................................ 1474.1.2 Wetting of Nonmetals by Solders .......................................................................... 149

4.1.2.1 Solderable Coatings on Nonmetals .................................................................. 1494.1.2.2 Active Solders .................................................................................................. 152

4.1.3 Erosion of Parent Materials ................................................................................... 1534.1.4 Phase Formation ..................................................................................................... 1544.1.5 Filler-Metal Partitioning ......................................................................................... 155

4.2 Mechanical Constraints and Solutions ............................................................................. 1574.2.1 Controlled Expansion Materials ............................................................................. 159

4.2.1.1 Iron-Nickel Alloys ............................................................................................ 1604.2.1.2 Copper-Molybdenum and Copper-Tungsten Alloys ........................................ 1614.2.1.3 Copper-Surface Laminates ............................................................................... 1624.2.1.4 Composite Materials ......................................................................................... 163

4.2.2 Interlayers ............................................................................................................... 1644.2.3 Compliant Structures .............................................................................................. 1654.2.4 The Role of Fillets ................................................................................................. 167

4.3 Constraints Imposed by the Components and Solutions ................................................. 1684.3.1 Joint Area ................................................................................................................ 169

4.3.1.1 Trapped Gas ...................................................................................................... 1694.3.1.2 Solidification Shrinkage ................................................................................... 173

4.3.2 Void-Free Soldering ............................................................................................... 1734.3.3 Joints to Strong Materials ...................................................................................... 175

4.3.3.1 Joint Design to Minimize Concentration of Stresses ...................................... 1754.3.3.2 Strengthened Solders to Enhance Joint Strength ............................................ 178

4.3.4 Thick- and Thin-Joint Gap Soldering .................................................................... 178Chapter 4: Appendices ................................................................................................................180

A4.1 A Brief Survey of the Main Metallization Techniques ......................................... 180A4.2 Critique of Void-Free Soldering Standards ........................................................... 183A4.3 Dryness and Hermeticity of Sealed Enclosures .................................................... 184

Chapter 5: Advances in Soldering Technology......................................................................... 1895.1 Lead-Free Solders ............................................................................................................. 189

5.1.1 The Drive for Lead-Free Soldering ....................................................................... 1905.1.2 Compatibility with Lead-Tin Solder ...................................................................... 1915.1.3 Alternatives to Lead-Tin Solder ............................................................................ 1915.1.4 Silver-Copper-Tin Ternary Phase Equilibria ......................................................... 1935.1.5 Metallurgical, Physical, and Chemical Properties of

Lead-Free Solders ............................................................................................... 1935.1.5.1 Surface Tension ................................................................................................ 1935.1.5.2 Other Physical Properties ................................................................................. 1945.1.5.3 Mechanical Properties ...................................................................................... 1945.1.5.4 Corrosion Resistance ........................................................................................ 1955.1.5.5 Susceptibility to Tin Pest and Tin Whiskers ................................................... 195

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5.1.6 Process Window for Lead-Free Solders ................................................................ 1965.1.7 Wetting and Spreading Characteristics of Lead-Free

Solders ................................................................................................................ 1975.1.8 High-Melting-Point Lead-Free Solders ................................................................. 197

5.2 Flip-Chip Interconnection ................................................................................................. 1995.2.1 The Flip-Chip Process ............................................................................................ 1995.2.2 Characteristics of Flip-Chip Technology ............................................................... 2025.2.3 Underfill .................................................................................................................. 2035.2.4 Inspection ................................................................................................................ 2035.2.5 Rework .................................................................................................................... 2045.2.6 Self-Alignment of Flip-Chip Structures ................................................................ 2045.2.7 Surface Topography ................................................................................................ 2065.2.8 Step-Soldered Flip-Chip Interconnects .................................................................. 206

5.3 Solderability Test Methods and Calibration Standards ................................................... 2075.3.1 Assessment of Wetting ........................................................................................... 2075.3.2 Assessment of Spreading ....................................................................................... 2105.3.3 Solderability Calibration Standards ....................................................................... 212

5.4 Amalgams as Solders ....................................................................................................... 2145.4.1 Amalgams Based on Mercury ............................................................................... 2155.4.2 Amalgams Based on Gallium ................................................................................ 2165.4.3 Amalgams Based on Indium .................................................................................. 217

5.5 Strengthening of Solders .................................................................................................. 2175.5.1 Grain Refinement ................................................................................................... 2185.5.2 Oxide-Dispersion-Strengthened Solders ................................................................ 2185.5.3 Composite Solders .................................................................................................. 219

5.6 Reinforced Solders (Solder Composites) ......................................................................... 2225.7 Mechanical Properties and Numerical Modeling of Joints ............................................. 223

5.7.1 Measurement of Mechanical Properties ................................................................ 2235.7.2 Numerical Modeling of Joints ............................................................................... 224

5.7.2.1 Dimensional Stability of Soldered Joints ........................................................ 2245.7.2.2 Prediction of Joint Lifetime ............................................................................. 226

5.8 Solders Doped with Rare Earth Elements ....................................................................... 2275.8.1 Effect of Rare Earth Additions on Solder Properties ............................................ 2275.8.2 Implications for Soldering Technology ................................................................. 229

5.9 Diffusion Soldering ........................................................................................................... 2305.9.1 Process Principles ................................................................................................... 2305.9.2 Diffusion Soldering of Silver ................................................................................. 2315.9.3 Diffusion Soldering of Gold .................................................................................. 2335.9.4 Diffusion Soldering of Copper ............................................................................... 2345.9.5 Practical Aspects ..................................................................................................... 2345.9.6 Modeling of Diffusion-Soldering Processes .......................................................... 235

5.10 Advances in Joint Characterization Techniques .............................................................. 2355.10.1 Ultrasonic Inspection (Scanning Acoustic Microscopy) ....................................... 2355.10.2 X-Radiography ....................................................................................................... 2365.10.3 Optical Inspection ................................................................................................... 237

Abbreviations and Symbols.......................................................................................................... 243

Index................................................................................................................................................ 245

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Preface

Since the first edition of Principles of Soldering and Brazing, published in 1993, the authors havereceived valuable feedback from readers representing a wide range of technical interests. This hasprompted the decision to expand the text and organize it into two companion books, one coveringsoldering and the other brazing. This first book primarily aims at providing information aboutsoldering in a form that is hopefully readily accessible and as easy to assimilate as possible. Priorityis given to the fundamental principles that underlie this field of technology rather than recipes formaking joints. The largely artificial distinctions between soldering and brazing are preservedbecause, despite their many commonalities, it has been found that practicing engineers are eitherconcerned with soldering or brazing and seldom are involved with both simultaneously. The plannedcompanion book, Principles of Brazing, addresses this complementary need. A large proportion ofthe literature on soldering and brazing may be charged with being heavy on description and light oncritical analysis. We have endeavored to redress the balance, while striving to avoid being undulysimplistic or overly mathematical in our approach. Admittedly we may not always have succeeded inthis aim.

As in Principles of Soldering and Brazing, we have striven to maintain the focus on the fundamentalaspects of soldering and have deliberately avoided entering into specific joining technologies in detail.At the same time, we recognize that the range and extent of the knowledge base of metal joining isnot immediately obvious, and it requires a fairly deep understanding of materials. To cite a singleexample, nichrome (an alloy of nickel and chromium), which is a perfectly satisfactory and widelyused metallization for soldering, is rendered useless if the solder contains bismuth. If there is an evidentbias towards electronic and photonics applications, this reflects the recent professional orientation ofthe authors. Some topics are inevitably not accorded due consideration, although it is hoped thatsufficient references are provided to enable the reader to pursue these further.

No attempt has been made to gather a comprehensive list of published papers. Those that areincluded have been selected because they are useful basic texts, cover important subject matter, orrelate to exemplary pieces of work, whether in respect of methodology, technique, or other noteworthyfeatures. It was felt that if the value of the book depended on its bibliography, it would rapidly becomedated. The advent of computer search facilities and databases of scientific journal and conferenceabstracts should enable the reader who wishes to find references on a specific topic to obtain furtherinformation without too much difficulty. The search term “lead-free solder” will yield an astounding25,000� publications in the public domain, virtually none of which are more than 10 years old.

The reader should note that all compositions given in this book are expressed in weight percentagein accordance with the standard industrial practice. These have, for the most part, been rounded tothe nearest integer. The ratio of elements in intermetallic compounds, again by convention, refers tothe atomic weight of the respective constituents. The general convention used for specifying alloycompositions is that adopted by the alloy phase diagram community, namely in the alphabetical orderof the elements, by chemical symbol. We have not been entirely rigorous in this regard as it issometimes helpful to group alloys by the dominant constituents. Minor additions to bulk compositionsare given in order of concentration; for example, Pb-62Sn-0.5Lu-0.02Ce.

Specific references are given with each chapter. For those wishing to read more generally onparticular topics, the authors would recommend the texts listed as Selected References at the end ofthis preface.

Many phase diagrams are subject to ongoing research, resulting in continued improvement in theaccuracy and detail of the information. The most recent version of a diagram may be identified byconsulting the latest cumulative index of phase diagrams, published in the Cumulative Index of theperiodical Journal of Phase Equilibria (ASM International). This will refer to the source of thethermodynamically assessed diagram of interest. The reader is advised that the four compendia ofbinary phase diagrams published in the 1960s, ’70s and ’80s (colloquially referred to as Hansen,Elliott, and Shunk) are now known to contain many errors and omissions.

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Information on new developments in soldering and brazing is scattered throughout a wide rangeof periodicals, as reflected in the sources cited in the references appended to the individual chapters.To keep abreast of the literature, the authors have found especially useful the following abstractpublications: Metals Abstracts and Science Abstracts. Technical libraries can provide automatedsearches against specified key words as a monthly service.

We wish to thank our many colleagues and ex-colleagues for their helpful advice and encourage-ment, particularly James Vincent, for insights into lead free soldering.

Giles HumpstonDavid M. Jacobson

SELECTED REFERENCES

Soldering

• Brandon, D. G., and Kaplan, W. D., 1997. Joining Processes: An Introduction, John Wiley & Sons• Frear, D.R., Jones, W.B., and Kinsman, K.R., 1990. Solder Mechanics: A State of the Art As-

sessment, TMS• Hwang, J.S., 1996. Modern Solder Technology for Competitive Electronics Manufacturing,

McGraw-Hill• Hwang, J.S., 2001. Environment Friendly Electronics: Lead-Free Technology, Electrochemical

Publications• International Organization for Standardization (IOS), 1990. Welding, Brazing, and Soldering

Processes: Vocabulary, (ISO/DIS 857-2), ISO (currently under revision)• Klein Wassink, R.J., 1989. Soldering in Electronics, 2nd ed., Electrochemical Publications• Liebermann, E,, 1988. Modern Soldering and Brazing Techniques, Business News• Manko, H.H., 2002. Solders and Soldering, 4th ed., McGraw-Hill• Nicholas, M.G., 1998. Joining Processes: Introduction to Brazing and Diffusion Bonding, Kluwer

Academic• Strauss, R., 1998. SMT Soldering Handbook, 2nd ed., Butterworth-Heinemann• Thwaites, C.J., 1983. Capillary Joining: Brazing and Soft–Soldering, Books Demand UMI• Woodgate, R.W., 1996. The Handbook of Machine Soldering: SMT and TH, John Wiley & Sons

Alloy Constitution

• John, V.B., 1974. Understanding Phase Diagrams, Macmillan• Prince, A., 1966. Alloy Phase Equilibria, Elsevier• West, D.R.F., 1982. Ternary Equilibrium Diagrams, Chapman and Hall

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About the Authors

Giles Humpston took a first degree in metallurgy at Brunel Uni-versity in 1982, followed by a Ph.D. on the constitution of solder al-loys in 1985. He has since been employed by several leading indus-trial companies, where he has been involved with determining alloyphase diagrams and developing processes and procedures for pro-ducing precise and high-integrity soldered, brazed, and diffusion-bonded joints to a wide variety of metallic and nonmetallic materials.His expertise extends to fine-pitch flip-chip, new materials develop-ment, and packaging and interconnection for electronics, radio fre-quency, and optical products. He is the cited inventor on more than75 patents, the author of more than 60 papers, and recipient of sixinternational awards for his work on soldering and brazing.

Dr. Humpston is a licensed amateur radio enthusiast and haspublished several articles and reviews on electronics, radio, andcomputing. His other interests include exploring vertical-axis wind

turbines, building power inverters, flying radio controlled gliders, wine making, and growing bonsai.He lives with his wife, Jacqueline, and their three children in a small village in Buckinghamshire,England and San José (Silicon Valley), California.

David M. Jacobson graduated in physics from the Universityof Sussex in 1967 and obtained his doctorate in materials sci-ence there in 1972. Between 1972 and 1975 he lectured inmaterials engineering at the Ben Gurion University, Beer-Sheva, Israel, returning as Visiting Senior Lecturer in 1979-1980. Having gained experience in brazing development withJohnson Matthey Ltd., he extended his range of expertise tosoldering at the Hirst Research Centre, GEC-Marconi Ltd.,which he joined in 1980. Currently, he holds the position ofsenior research associate at the Centre for Rapid Design andManufacture, Buckinghamshire Chilterns University College inHigh Wycombe. He is the author of more than 80 scientific andtechnical publications in materials science and technology andmore than a dozen patents. He has been awarded three presti-gious awards for his work on brazing.

Dr. Jacobson’s principal outside interests are archaeology and architectural history, focusing on theNear East in the Graeco-Roman period. He has published extensively in these fields on subjects thatextend to the numismatics and early metallurgy of that region. He recently completed a Ph.D. thesison Herodian architecture at King’s College, London, and teaches part-time in this subject area atUniversity College, London. Dr. Jacobson is married with two grown-up children and lives inWembley, England, close to the internationally famous football stadium.

Giles Humpston and David Jacobson are the coauthors of the book Principles of Soldering andBrazing, which was published by ASM International in 1993, with more than 4000 copies sold.

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History

Origins of Solders and Soldering

The word solder derives from the Old French, soudure, which in turn stems from the Latin solidare,which means to fasten together. Its earliest use in a completely English context as a noun meaning“a fusible metallic alloy used for uniting less fusible metal surfaces or parts” dates to about 1350.It is interesting to note that in 19th century English, just as in modern French, the “l” would havebeen omitted and the word pronounced “sod-der,” a form that still persists in the United States ofAmerica today.

Although the origin of solders and soldering is lost to antiquity, it is possible to speculate on howthe invention arose. Lead was first obtained as a by-product of silver production. Silver extractionfrom ores involved cupellation of lead, and the base metal was then recovered from the litharge[Tylecote 1976]. The softness and malleability of lead were clearly recognized, and there existexamples of lead being used as a setting agent to fix posts in the ground and lock morticed stones.It was observed that in this instance the lead filler could give a stronger joint than a simple frictiongrip. Lead was used by the Mesopotamians (3000 B.C.) to join pieces of copper together, althoughperhaps more by luck than design since pure lead does not wet copper at all readily. The Romans areknown to have produced lead separately from silver, taking advantage of the fact that this metal canbe easily extracted from its sulfide ore, galena, simply by roasting the mineral in air [Tylecote 1976].

The earliest examples of tin are Egyptian and date from 2000 B.C. What might be construed as amanufactured solder alloy has been found in King Tutankhamun’s tomb (1350 B.C.), although thereis some debate among scholars about the deliberateness of the metallurgy of this joint.

Solders comprising alloys of lead and tin were almost certainly used during the Iron Age [Tylecote1962]. By the Roman Imperial period there is evidence, both from literary sources and from survivingartifacts, that lead-tin solders were in regular use. Pliny the Elder (1st century A.D.) speaks of tertiarum,an alloy of two parts of (black) lead and one part of white lead (tin) being used for joining metal pipes[Pliny, Natural History xxxiv 161 (Rackham 1952)]. Pliny also remarks that the price of this alloyis 20 denarii per pound. With 25 denarii (silver pieces weighing approximately 4 gm, or 0.14 oz, each)to 1 gold aureaus of close to 8 gm (0.28 oz), the price of Roman solder works out at $70 per kilogram,assuming that gold has maintained its purchasing power since Pliny’s day. The current price for thesame alloy (Pb-33Sn) is lower by an order of magnitude, which indicates how much more precioussolder was in antiquity.

An analysis of soldered joints in Roman artifacts has shown that both tin-rich and lead-rich alloyswere used. The solder in a force-pump from Roman Silchester contains lead to tin in a weight ratioof close to 3 to 1, which is similar to the composition of plumbers’ solder [Tylecote 1962]. Elsewhere,solders containing mainly tin (80 to 100% Sn), have been encountered in finds from 4th and 5th centurysites in Britain [Lang and Hughes 1991].

Soldering, unlike many Roman crafts, either did not die out during the Dark Ages or enjoyed anearly revival. The soldering iron, not mentioned at all in Classical times, was well known and inwidespread use by the early Middle Ages. Soldering was used for joining the lead strips in stainedglass windows, with the oldest complete examples being the Five Prophets windows in AugsburgCathedral that date from the late 11th century. From 1700 onwards it is clear that soldering was wellestablished with the appearance of “tinsmiths” and “white-iron men” as trades. Newcomen’s discoveryof the effectiveness of the internally condensing steam engine in 1708 is attributed to the faulty repair,by soldering, of a blowhole in the cast bronze cylinder. This permitted a spray of external condenserwater into the cylinder and the development of the internal condenser; a design that was not supersededuntil Watt developed the separate condenser nearly 70 years later.

Modern soldering practice dates to the early 20th century when improved extraction techniques,which enabled exotic metals to be available at affordable cost, coupled with the appearance of alloyphase diagrams, gave rise to the diversity of alloys now available.

x



REFERENCES

• Tylecote, R.F., 1976. A History of Metallurgy, The Metals Society• Tylecote, R.F., 1962. Metallurgy in Archaeology: A Prehistory of Metallurgy in the

British Isles, Edward Arnold• Rackham, H., 1952. Natural History, Vol 10, Cambridge, MA, Translation of Pliny

1. Historia Naturalis, Vol 34 (No. 161)• Lang, J. and Hughes, M.J., 1991. “Joining Techniques in Aspects of Early Metal-

lurgy,” British Museum Occasional Papers, No. 17, British Museum, p 169–177

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CHAPTER 1

Introduction

1.1 Joining Methods

SOLDERINGAND BRAZING represent oneof several types of methods for joining solidmaterials. These methods may be classified as:

• Mechanical fastening• Adhesive bonding• Soldering and brazing• Welding• Solid-state joining

Other methods, such as glass/metal sealing, elec-trostatic welding, and so forth, are dealt withelsewhere [Bever 1986].

Schematics of these joiningmethods are givenin Fig. 1.1. These different methods have a num-ber of features in common but also certain sig-nificant differences. For example, soldering andbrazing are the only joining methods that canproduce smooth and rounded fillets at the pe-riphery of the joints. The joining methods arelisted in the first paragraph of this chapter in theorder in which they lead to fusion of the jointsurfaces and tend toward a “seamless” joint.

Because soldering and brazing lie in themiddleof this sequence, they share several features withthe other methods. For example, soldered andbrazed joints can be endowed with the advan-tageousmechanical properties ofwelded and dif-fusion-bonded joints; at the same time they canbe readily disassembled, without detriment to thecomponents, like mechanically fastened joints.These featuresmake soldering and brazing highlyversatile.

The principal characteristics of the variousjoining methods are summarized in the para-graphs that follow.

1.1.1 Mechanical Fastening

Mechanical fastening involves the clampingtogether of components without fusing the joint

surfaces. This method often, but not always, re-lies on the use of clamping members such asscrews and rivets. In crimping, the componentsare keyed together by mechanical deformation.

Characteristic features of mechanical fasten-ing include:

• Aheating cycle is generally not applied to thecomponents being joined. A notable excep-tion is riveting, where the rivets used forclamping are heated immediately prior to thefastening operation. On subsequent coolingthe rivets shrink, causing the components tobe clamped tightly together.

• The reliance on local stressing to effect join-ing requires thickening or some other meansof reinforcement of the components in thejoint region. This places a severe restrictionon the joint geometries that may be used andimposes a weight penalty on the assembly.Another constraint on permissible joint con-figurations is the need for access to insert theclamping member.

• Themethod usually requires specialmechani-cal preparation, such as drilling holes, ma-chining screw threads, or perhaps chamfer-ing of abutting surfaces, in the case ofcomponents to be crimped.

• The choice of suitable joint configurations ishighly dependent on service conditions—forexample, whether or not leak tightness is re-quired. Joints may be designed to accommo-date thermal expansion mismatch betweenthe components in the assembly. In the ex-treme case, joints can bemade to permit com-plete freedom of movement in the plane per-pendicular to the clamping member, asapplied to the fishplates used to couple trainrails.

• The electrical and thermal conductance acrossthe joint is a function of the effective area thatis in contact. This depends on many other

Principles of Soldering Giles Humpston, David M. Jacobson, p1-47 DOI:10.1361/prso2004p001

Copyright © 2004 ASM International® All rights reserved. www.asminternational.org

parameters, such as the clamping force andthe materials used, and in service the con-ductance is unlikely to be constant.

1.1.2 Adhesive BondingAdhesive bonding involves the use of a poly-

meric material, often containing various addi-tives, to “stick” the components together. Theprocess involves a chemical reaction, which maysimply comprise exposure of the adhesive to air,leading to the formation of a hydrogen-type bond

between the cured adhesive and the respectivecomponents. The original interfaces of the jointare preserved in this type of bonding process.

Characteristic features of adhesive bonding in-clude:

• It is inherently a low-stress joining methodbecause it is carried out at relatively low tem-peratures andmost adhesives have high com-pliance.

• A diverse range of methods are available forcuring adhesives.

Fig. 1.1 Principal methods for joining engineering materials

2 / Principles of Soldering

• The geometry of the components tends not tobe critical.

• Constraints apply to the geometry of the ac-tual joint; in particular large areas and verynarrow gaps are necessary to ensure me-chanical integrity.

• Joints tend to be weak when subject to forcesthat cause peeling. For this reason, adhesivejoints are frequently used in combinationwithmechanical fastening—for example, in air-frame assembly.

• Joint integrity tends to be sensitive to the at-mosphere of the service environment and tothe state of cleanliness of the mating surface.

• The service temperature range of adhesivelybonded joints is usually limited, as is theircompatibility with solvents.

• Joints usually possess poor electrical and ther-mal conductivity, although by loading the or-ganic adhesive with metal particles moderateconductance can be achieved that approachesthat of some solder alloys.

Polymer chemistry is a rapidly evolving sci-ence. As a result, some very advanced adhesiveshave appeared on the market in the last few yearswith properties highly tailored for particular func-tion in the electronics industry. These includethermally conductive adhesives, electrically con-

ductive adhesives, and anisotropically conduc-tive adhesives. Table 1.1 provides an indicationof the product range of thermally conductive ad-hesives available from one manufacturer. Theseadvanced materials are being augmented by thedevelopment of polymers with high imperme-ability to moisture and low thermal expansioncoefficients for use as electronic packaging ma-terials. Other polymers have been developed thatfunction as both flux and underfill material forflip-chip applications. More exciting advanceswill no doubt continue to become available.

1.1.3 Soldering and Brazing

Soldering and brazing involve using a moltenfiller metal to wet the mating surfaces of a joint,with or without the aid of a fluxing agent, leadingto the formation of metallurgical bonds betweenthe filler and the respective components. In theseprocesses, the original surfaces of the compo-nents are “eroded” by virtue of the reaction oc-curring between the molten filler metal and thesolid components, but the extent of this “ero-sion” is usually at the microscopic level (<100μm, or 4000 μin.). Joining processes of this type,by convention, are defined as soldering if the

Table 1.1 Selection of commercially available conductive adhesives, used in place of solder forsome applications

4030SD 4030LD 4030Hk 4030SR 4130HT 5030P 6030Hk(a)

Paste Properties

Viscosity at 10 rpm,kcP(b)

30–45 25–40 25–45 35–50 30–40 30–40 30(typical)

25 °C (77 °F) shelf life,months

6 6 6 6 6 6 6

Paste density, g/cm3 3.17 3.42 4.35 3.07 3.33 3.7 3.8–4.5

Processed Properties

Die shear(c)(d),kgf/cm2 (psi)

56 (800) 56 (800) 42 (600) 98 (1400) 105 (1500) 105 (1500) 105 (1500)

Bulk resistivity(d),μ� • cm

40 40 13 160 40 25 6–10

Thermal conductivity,W/m • K

15 15 35–40 5 20 20–25 30–60

Young’s modulus,GPa (ksi)

0.9 (125) 0.9 (125) 0.9 (125) . . . 1.8 (250) 2.5 (350) . . .

Thermal expansion,10–6/°C (10–6/°F)

15 (28) 15 (28) . . . . . . 17 (30) 13 (23) . . .

Rework temperature(e),°C (°F)

100 (212) 100 (212) 100 (212) 150 (300) 150 (300) 200–250(390–480)

(f)

Designed product use Highlyconductiveadhesive

Large areadeviceattach

Very highthermal andelectricalconductivityadhesive(g)

Solderreplacementfor SMT

Higher-temperatureapplications

Withstandswire bondingat 250 °C(480 °F)

Extremelyhighthermal/electricalconductivity

Note in particular the cited thermal conductivity values are comparable to those of many solders. (a)Anovel thermoset development material capable of room-temperaturestorage. (b) kcP is 1000 centipoise (1 Pa • s). (c) 6.48 mm (0.255 in.) die, ceramic. (d) 175 °C/15 min profile. (e) 70 kPa (10 psi) force. (f) Rework as for typical epoxies.(g) Maximum die size 7 � 7 mm (0.28 � 0.28 in.). Source: Multicore Solders Ltd.

Chapter 1: Introduction / 3

filler melts below 450 °C (840 °F) and as brazingif it melts above this temperature.

Characteristic features of soldering and braz-ing include:

• All brazing operations and most solderingoperations involve heating the filler and jointsurfaces above ambient temperature.

• In most cases, the service temperature of theassemblymust be lower than themelting tem-perature of the filler metal.

• It is not always necessary to clean the sur-faces of components prior to the joining op-eration because fluxes are available that arecapable of removingmost oxides and organicfilms. However, there are penalties associ-ated with the use of fluxes—for example, theresidues that they leave behind, which areoften corrosive and can be difficult to re-move.

• The appropriate joint and component geom-etries are governed by the filler/componentmaterial combination and by service require-ments (need for hermeticity, stress loading,positional tolerances, etc.). Complex geom-etries and combinations of thick and thin sec-tions can usually be soldered or brazed to-gether.

• Intricate assemblies can be produced withlow distortion, high fatigue resistance, andgood resistance to thermal shock.

• Joints tend to be strong if well filled, unlessembrittling phases are produced by reactionbetween the filler metal and the components.

• Soldered and brazed joints can be endowedwith physical and chemical properties thatapproximately match and, in some cases,even exceed those of the components, butusually have limited elevated-temperatureservice and stability.

• Fillets are formed under favorable condi-tions. These can act as stress reducers at theedges of joints that benefit the overall me-chanical properties of the joined assembly.

Soldering and brazing can be applied to a widevariety of materials, including metals, ceramics,plastics, and composite materials. For many ma-terials, and plastics in particular, it is necessaryto apply a surface metallization prior to joining.

1.1.4 WeldingWelding involves the fusion of the joint sur-

faces by controlled melting through heat beingspecifically directed toward the joint. Commonly

used heating sources are plasma arcs, electronbeams, lasers, and electrical current through thecomponents and across the joints (electrical re-sistance). Filler metals may be used to supple-ment the fusion process for components of simi-lar composition, as for example when the jointgap is wide and of variable width. In that situ-ation, the filler is often chosen to have a mar-ginally lower melting point than the componentsin order to help ensure that it completely melts.

Characteristic features of welding include:

• Welding invariably involves a heating cycle,which tends to be rapid, and a very widevariety of welding processes are available.

• Welding cannot be used to join metals tononmetals or materials of greatly differingmelting points. There are exceptions to this,but these are generally limited to precise com-binations of materials and highly specificwelding methods.

• Joint geometries are limited by the require-ment that all joint surfaces are accessible tothe concentrated heat source.

• Welded joints may approach the physical in-tegrity of the components, but are often in-ferior in their mechanical properties, particu-larly fatigue resistance. This is due to stressconcentrations produced by the high thermalgradients developed during joining and therelatively rough surface texture of welds.

• The heating cycle usually affects the micro-structure and hence the properties of the com-ponents over a macroscopic region aroundthe joint, called the heat-affected zone (HAZ).The HAZ is often influential in determiningthe properties of welded joints.

• Welding tends to distort the components inthe region of the HAZ. This is associatedwith the thermal gradients developed throughthe use of a concentrated heat source to fusethe joint surfaces.

1.1.5 Solid-State Joining

The term solid-state joining covers a verywiderange of joining processes. The two extremes arepressure welding and diffusion bonding. Pres-sure welding, at its simplest, involves physicallydeforming two abutting, faying surfaces to dis-rupt any intervening surface films and enabledirect metal-to-metal contact. Diffusion bondingin its purest form merely requires placing twofaying surfaces in contact and heating the as-sembly until the voids at the interface have been


removed by diffusion. Further details of processparameters for diffusion bonding of gold andindium are given in Appendix A1.1. Pressurewelding generally works better if the compo-nents are heated (e.g., friction welding), and dif-fusion bonding is usually greatly accelerated bythe application of pressure or mechanical agita-tion (e.g., thermosonic ball and wedge bonding,see Fig. 1.2) to force a greater area of the fayingsurfaces into contact.

Solid-state joining constitutes a subject in itsown right, quite separate from soldering and braz-ing, which rely on liquid-state metal joining.However with the development of the diffusion-soldering and diffusion-brazing processes, whichare a hybrid of the two, some consideration ofsolid-state joining, in particular diffusion bond-ing, is merited. Pressure welding is sometimesused to prepare filler metals in various geom-etries and to tack preforms in position. Occa-sionally, conventional filler metals are used toproduce a pressure weld or diffusion bond be-tween dissimilar metals in a solid-state joiningprocess. For this reason, further information onboth pressure welding and diffusion bonding isincluded in section 1.1.7 in this chapter.

Characteristic features of solid-state joining:

• This method generally involves heating thejoint to a temperature below themelting pointof the components.

• Pressure welding is often a much faster pro-cess than soldering or brazing (<1 s), the

extreme being explosive welding, while dif-fusion bonding is much slower (>10 min).

• The joints have no fillets.• The service temperature of joined assemblies

can be higher than the joining temperatureand tend toward themelting point of the com-ponents.

• Solid-state joining is limited in application tospecific combinations of materials that pro-vide specific combinations of mechanical ordiffusion characteristics.

• Of all the joining methods, they are the leasttolerant of poor mating of the joint surfaces.

• Joint surfaces need to be scrupulously cleanbecause solid-state joining is a fluxless pro-cess.

• The properties of solid-state joints can ap-proach those of the parent materials.

Further details on pressure welding are given insection 1.1.7.1 and on diffusion bonding in sec-tion 1.1.7.2 in this chapter.

1.1.6 Comparison betweenSolders and Brazes

In many respects it is fruitful to consider sol-ders together with brazes. This integrated treat-ment can be justified on metallurgical grounds.These two classes of filler cannot be demarcatedby a temperature boundary as is habitually done:conventionally, solders are defined as filler met-als with melting points below 450 °C (840 °F)and brazes as having melting points above thisvalue. This distinction has a historical origin: theearliest solders were based on alloys of tin, whilebrazes were based on copper-zinc alloys (see“History of Soldering” in this volume and “His-tory of Brazing” in the planned companion vol-ume Principles of Brazing for a brief historicalbackground of solders and brazes, respectively).

The type of metallurgical reaction between afiller and parent metal is sometimes used to dif-ferentiate soldering from brazing. Solders usu-ally react to form intermetallic phases, that is,compounds of the constituent elements that havedifferent atomic arrangements from the elementsin solid form. By contrast, most brazes formsolid solutions, which are mixtures of the con-stituents on an atomic scale. However, this dis-tinction does not have universal validity. For ex-ample, Ag-Cu-P brazes react with steels to formthe interfacial phase of Fe3P in a similar mannerto the reaction of tin-base solders with iron andsteels to form FeSn2. On the other hand, solid

Fig. 1.2 An electronic module in which the semiconductordies have been interconnected using fine wire at-

tached by thermocompression bonding


solutions form between silver-lead solders andcopper just as they do between the common sil-ver-base brazes and copper. Also, there existbrazes for aluminum that melt below 450 °C(840 °F).

Soldering and brazing involve essentially thesame bonding mechanism: that is, reaction withthe parent material, usually alloying, to formme-tallic bonds at the interface. In both situations,good wetting promotes the formation of filletsthat serve to enhance the strength of the joints.Similar processing conditions are required, andthe physical properties are comparable, providedthe same homologous temperature (the tempera-ture at which the properties are measured as afraction of the melting temperature expressed indegrees Kelvin) is used for the comparison.

The perpetuation of the distinction of soldersfrom brazes on the basis of the 450 °C (840 °F)boundary has arisen from the significant gap thatexists between the melting points of availablesolder alloys, the highest being Au-3Si, whichmelts at 363 °C (685 °F), and the lowest tem-perature standard braze, the Al-4Cu-10Si alloy,which melts at 524 °C (975 °F) but, being anoneutectic alloy, is fully liquid only above 585°C (1085 °F). Eutectic alloys are defined inChap-ter 2, section 2.3; for the present, it shall sufficeto state that eutectic alloys are akin to pure met-als in melting and freezing at the same tempera-

ture. The temperature ranges of the principal sol-der and braze alloy families are shown in Fig. 1.3and 1.4.

For most purposes, the temperature gap be-tween solders and brazes is substantially widerthan 160 °C (290 °F). This is because the gold-base solders are very expensive and are largelylimited in use to the high added-value manufac-turing of the electronics industry. Removing thehigh-gold-content alloys from consideration, thehighest-melting-point solders are the lead-richalloys, which melt at about 300 °C (570 °F). Thelowest-melting-point brazes that are used com-mercially in significant quantities are the rea-sonably ductile aluminum-silicon alloys, whichmelt at 577 °C (1070 °F). A selection of eutecticalloys with melting points in the temperatureinterval 300 to 550 °C (570 to 1020 °F) that atsome time or other have been promoted as sol-ders and brazes are listed in Table 1.2. They are,without exception, brittle and often contain oneor more volatile constituents, notably magne-sium, cadmium, or zinc. Some multicomponentalloys that have been developed and are designedto fill the temperature gap are described in Chap-ter 2. However, none of them are readily avail-able from commercial sources.

The dearth of filler metals with melting pointsin the range 300 to 550 °C (570 to 1020 °F) isnot necessarily a handicap; techniques are avail-

Fig. 1.3 Principal solder alloy families and their melting ranges


able for making joints using molten filler metalwith effective melting points in this temperatureinterval. Transient-liquid-phase diffusion bond-ing is one such example and is discussed inChap-ter 5, section 5.9.

From the “maps” of solders and brazes in Fig.1.3 and 1.4, it might appear that there are manymore solders than brazes. In fact, the contrary istrue. The alloys that are specifically indicated in

these figures are the mostly eutectic composi-tions or those characterized byminimummeltingranges. Most commercially used solders are in-cluded because these are almost all of eutecticcomposition. However, whole families of brazeshave been omitted because there is no eutectic inthe alloy system; instead they exhibit completeintersolubility. Examples are the copper-nickel,silver-gold, silver-palladium, and silver-gold-

Fig. 1.4 Principal braze alloy families and their melting ranges

Table 1.2 Selected eutectic alloys that are offered as high-melting point solders and low-meltingpoint brazes

Melting pointSolder composition(a),wt% °C °F Problems

5Ag-95Cd 340 644 Toxic fumes, volatile75Au-25Sb 356 673 High cost, brittle88Au-12Ge 361 682 High cost, brittle97Au-3Si 363 685 High cost, brittle6Al-94Zn 381 718 Volatile, brittle48Al-52Ge 424 795 High dross, brittle36Al-37Mg 450 842 Volatile, brittle75Pb-25Pd 454 849 Poor fatigue resistance, brittle56Ag-44Sb 485 905 Volatile, brittle58Au-42In 495 923 High cost, brittle68Al-27Cu-5Si 524 975 Difficult to clean, brittle23Ag-53Cd-24Cu 525 977 Toxic fumes, volatile, brittle24Cu-76Sb 526 979 Volatile, brittle62Cd-38Cu 549 1020 Toxic fumes, volatile, brittle

(a) All compositions given are in weight percent.


palladium alloys. Alloys in such systems meltover a temperature range that varies with thecomposition.

The higher process temperatures needed tomake a brazed joint have important conse-quences because more thermal activation energyis present. These are:

• More extensive metallurgical reaction be-tween the filler metal and the substrate. Sol-ders typically do not dissolvemore than a fewmicrons of the component surfaces, whereasbrazes often dissolve tens of microns. Largerchanges in the composition of the filler metaltherefore occur during brazing, which in turnsignificantly affects the fluidity of and wet-ting by the molten filler as well as the prop-erties of the joint.

• Greater reactivity with the atmosphere sur-rounding the workpiece.All other factors be-ing equal, brazes are less tolerant of oxidiz-ing atmospheres than solders, but, for thesame reasons, are also better suited to clean-ing by reducing atmospheres. When jointsare made in air with the aid of a flux, thegreater reactivity of brazes means that ahigher proportion of flux to filler metal isgenerally required. In consequence, flux-cored solders are adequate for use in air, whilebrazing rods intended for use in ambient at-mosphere must be provided with a thick ex-ternal coating of flux. Fluxes are discussed inChapter 3, section 3.2.

Most, but not all, soldering and brazing pro-cesses are performed at small excess tempera-tures above the melting point of the filler metal,commonly referred to as the “superheat.” Muchhigher process temperatures are occasionallyused where it is desirable to exploit thermal ac-tivation. For example, tin-containing solders canwet and join nonmetallized ceramics providedthe solder incorporates an active ingredient, suchas titanium, and the alloy is heated above about900 °C (1650 °F) [Kapoor and Eagar 1989]. Al-though the freezing point of the solder is un-changed at about 250 °C (480 °F), such “acti-vated” solders have several of the characteristicsof brazes at the process temperature. Further in-formation on these alloys can be found in Chap-ter 4, section 4.1.2.2. On the other hand, thealuminum-germanium eutectic alloy melting at424 °C (795 °F) behaves like a typical braze onaluminum and copper surfaces, although by con-ventional definition, it is classed as a solder.

Several general features distinguish the ma-jority of solders from common brazes, namely:

• Most commercial solders are of eutectic com-position because there is usually a need tominimize the processing temperature whilemaintaining reasonable fluidity of the moltenfiller. Also, solders are intrinsically soft andmust be conferred with optimal mechanicalproperties; generally these are achieved byhaving a fine-grained microstructure, whichis a characteristic feature of a true eutecticalloy.

• Most brazes, by comparison, possess mutualsolid solubility between their constituents andare therefore offered with a wide range ofcompositions and melting ranges. The lowdegree of intersolubility and the propensity toform intermetallic compounds possessed bysolder alloys are related to their constituentelements having a noncubic crystal symme-try.

• Solders find application at temperatures at afraction between 50 and 90% of their meltingpoint in degrees Kelvin, under strain levelsthat often exceed 10%. At these relativelyhigh temperatures, the alloys are not metal-lurgically stable and the joint microstructuretends to change with time. Brazes tend to beused at temperatures that are relatively muchlower and usually below half their meltingpoint in degrees Kelvin.

These points are discussed in further detail inChapters 2 and 3, and reference should also bemade to the planned companion volume Prin-ciples of Brazing. Notwithstanding the differ-ences, solders and brazes operate on similar prin-ciples, and hence the frequent use of thecollective term “filler” throughout this book hassome justification.

1.1.7 Pressure Welding andDiffusion Bonding

Solid-state joining methods are not new, andexamples of gold-base artifacts fabricated usingpressure welds have been dated to 1000 B.C.[Tylecote 1968], while a cup and chalice deco-rated by diffusion bonding have been dated to3200 B.C. [Tylecote 1967]. Although more re-cent interest in welding has been almost totallydominated by fusion welding processes, bothpressure welding and diffusion bonding continueto satisfy niche applications because of theunique combination of process and joint param-


eters they offer. Some solid-state joining proce-dures are a combination of pressure welding anddiffusion bonding, as is evident from the funda-mental characteristics of each.

1.1.7.1 Pressure Welding

Pressure welding utilizes pressure to rupturesurface films at the joint interface and also toextrude virgin parent metal between islands ofsurface contamination so that metallic bondingcan take place. Thus, the process is character-ized by high pressures, applied for short peri-ods of time, on metals that may be either coldor hot. By necessity, bulk plastic deformationof the metals will occur. Possibly the mostcommon examples of pressure welding that arepertinent to solders and brazes are butt weld-ing to join lengths of wires, roll-bonding, andindentation welding.

In pressure welding, it is generally acceptedthat bond formation is controlled by the extent ofdeformation of the faying surfaces. The term“threshold deformation” is used extensively inthe literature on this subject and is defined as theminimum deformation needed to achieve anybonding, although the strength of a bond at thislevel of deformation is generally much less thanthat of the parent metal (see Fig. 1.5).

The bonding process can be described as fourconsecutive stages:

1. Removal of surface contamination andbreakup of brittle surface layers, in particu-lar oxides. This is frequently accomplished

by mechanically abrading the surface im-mediately prior to bonding. Adsorbed wateris believed to be the main surface contami-nant and responsible for preventing bondingif the deformation is less than 8%. Typically,40% deformation is required to affect asound joint when bonding metals in atmo-spheres other than vacuum.

2. Establishment of physical contact betweenregions of uncontaminated metal as virginmetal extrudes between gaps in the rupturedsurface films

3. Activation of contacting atoms to form ametallic bond. The contact area determinesthe extent of bonding.

4. Subsequent atom rearrangement as a con-sequence of postweld heat treatment and/orstress relaxation

Pressurewelding is particularly effectivewhenjoining dissimilar metals. For good weldability,the softer metal should have the more brittle andstronger oxidefilmandviceversa.Thehardoxidelayer can then promote and assist in the breakupof the surface layers on the harder metal, but isitself easily ruptured by the yielding metal sup-porting it. For example, the oxide on aluminumfulfills the requirements of strength and brittle-ness compared to the oxides ofmost othermetals,while the metal is relatively soft, and thereforepressurewelding of aluminum to othermetals oc-cursatlowerdeformationsthanwhenautogenous-ly welded. Also, the different deformation char-acteristics of dissimilar metals may result ininterfacial movement that will enhance bondingcomparedtoautogenouswelding.Theuseofpres-sure welding to fabricate ductile preforms ofbrittle alloys using partitioned constituents is dis-cussed further in Chapter 4, section 4.1.5.

1.1.7.2 Diffusion Bonding

Diffusion bonding relies on a combination oftemperature and time to remove voids from thefree interfaces between two abutting metal parts.Fundamentally, the process is defined as one inwhich no plastic deformation of the componentsbeing joined takes place, although it is usual toapply some pressure to ensure that the nominallyflat faying surfaces are indeed in intimate con-tact. Typical process conditions are durations ofup to several hours at temperatures that may beas high as two-thirds of the melting point of theleast thermally stablemetal in the bonded couple.The use of long times at relatively high tem-

Fig. 1.5 The strength of pressure-welded joints as a functionof the deformation induced during the bonding pro-

cess. Below the threshold deformation level, no joining occurs.With increasing deformation the joint strength also increases even-tually up to that of the parent materials. Note that the joiningprocess modifies the properties of the parent material as it willwork-harden when mechanically deformed.


peratures necessitates some form of atmospherecontrol to preserve surface cleanliness. Soft(roughing) vacuum and controlled atmospheresare equally suitable.

Since diffusion processes are the main mecha-nisms for bonding, with no means for the physi-cal displacement of any intervening nonmetallicsurface films, there are two major considerationsin diffusion bonding. The first is the necessity toensure that these films do not constitute a barrierto atom migration. Secondly, in bimetallic sys-tems the formation of intermetallic compoundsand porosity arising from inequality of diffusionrate by different species (Kirkendall porosity)must be controlled. Table 1.3 presents some ofthe better-known direct diffusion-bonding com-binations of metals and metalloids.

Diffusion bonding does not take place by onedominant mechanism, but is a consequence ofone or more possible mechanisms that often op-erate in parallel. Each mechanism results in ma-terial (or void) transport so that the surface en-ergy associatedwith the interface is progressivelyreduced as joining proceeds. Some possiblemechanisms include:

• Plastic yielding of surface asperities• Creep of the surface asperities• Surface and volume diffusion altering the

shape of the voids• Grain boundary and volume diffusion from

the bond interface to reduce the void vol-ume

A detailed theoretical treatment of solid-statediffusion bonding is provided byHill andWallach[1989]. In practice, the extent of bonding and therate at which it is achieved is governed both bymaterial properties (such as surface, grain bound-ary, and volume diffusion coefficients, creep andyield strength, etc.) and process parameters, ofwhich the four main variables are:

• Pressure. Adequate pressure is required toachieve contact on an atomic scale by local-ized deformation of asperities on the nomi-nally flat surfaces being joined and also toallow creep mechanisms to contribute tobonding.

• Temperature. Thermal energy promotes fasterbonding since plastic deformation, creep, andall diffusion mechanisms are temperature de-pendent. Typically, temperatures around 0.7Tm are used, where Tm is the absolute meltingtemperature of the lowest melting point com-ponent. Heating rates are not critical.

• Time. Creep and diffusion mechanisms arealso strongly time dependent, and there mustbe a sufficient interval afforded to allow forvoid closure by material transfer. As the tem-perature increases, the time required for bond-ing decreases.

• Surface condition. The height and fre-quency of surface asperities defining the jointwill control the extent of initial surface con-tact and thus influence the bonding rate. Gen-erally, flatter and more highly polished sur-faces are easiest to bond. The removal ofsurface contamination and thick oxides is es-sential prior to bonding since these will eitherpersist at the joint line or must be removed bysolution in the parent material as bondingproceeds. It therefore takes higher relativetemperatures and pressures to bond alumi-num-base alloys than copper-base alloys.Paradoxically, tungsten, titanium, and tanta-lum do not exhibit diffusion-barrier prob-lems, despite the fact that their oxides or car-bides are very stable. Certain titanium alloysare of particular industrial interest since theycan be diffusion bonded and superplasticallyshaped in one processing operation, madepossible because above about 900 °C (1650°F), titanium can dissolve oxygen into its vol-ume as fast as a surface scale can form.

Diffusion-bonded joints normally exhibit 80to 100% of parent material strength. One per-ceived problem with diffusion bonding is thethermal cycle time, particularly compared withfusion welding. However, a complex weldingjob might take several hours to prepare and jigthe components, in which case diffusion bondingmight offer advantages.Unlikemostwelding pro-cesses, the process-time curve for diffusion bond-ing is almost flat in relation to the job size, be-cause the process time is essentially independentof joint area provided adequate compressivestress is applied.

Interlayers are often used to diffusionbonddis-similar metals. For example, silver foil is used inbonding steel to titanium and nickel foil is oftenused to bond high-carbon steel to itself and othermaterials.Agoldflashapplied toaprecleanedsur-face permits diffusion bonding of nickel and cop-per components.An obvious extension to this ap-proach is making use of interlayers that melt,thereby increasing diffusion rates, helping to fillthe joint gap and disrupt surface films; this is thecharacteristic feature of diffusion-soldering anddiffusion-brazing processes. Diffusion solderingisdiscussed inChapter5, section5.9anddiffusion


Tabl

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3M

atrix

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dm

etallo

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nbe

diffus

ion

bond

ed

Xindicatespotentially

workablecombinations

Ag

Al

Au

Be

Cr

Co

Cu

Fe

Nb

Mg

Mo

Ni

Pd

Pt

TaTi

UV

W

Stainl

ess

stee

l

Cas

t

iron

Car

bide

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phite

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Ada

pted

from

[Fea

ture

1976

]


brazing in the planned companion volume Prin-ciples of Brazing.

1.2 Key Parameters of Soldering

The quality of soldered joints depends stronglyon the combination of filler and component ma-terials, including surface coatings that may beapplied to the components, and also on the pro-cessing conditions that are used. It is preciselyfor this reason that a sound understanding of themetallurgical changes accompanying the se-quence of events that occur in making solderedjoints is so vital for developing reliable joiningprocesses.

Soldering technology has generally evolvedin an empirical manner, largely by trial and er-ror. Theoretical principles have helped to fur-nish insights, guidelines, and qualitative expla-nations for this technology, but have rarelyprovided reliable data for use in the design ofjoining processes. The basic difficulty is thatthe real situation is highly complex, as it bringsinto play a large number of variables, some ofwhich may not be easy to recognize. Amongthe relevant factors are the condition of thesolid surfaces (i.e., the nature of any oxides orother coatings, surface roughness, etc.) and thetemperature gradients that develop during thejoining operation, as well as the metallurgicalreactions involving the filler and parent mate-rials and also the chemical reactions withfluxes, where these are used.

Another key aspect of joiningwith fillers is themanner and extent of flow of the molten fillerinto the joint. These are influenced by:

• Dimensions of the joint• Spread characteristics of the filler metal• Surface condition of the components

The limitation of theory in accounting for ob-served behavior is well illustrated by the clas-sical model of wetting and spreading. Thismodelnevertheless does provide useful concepts andinsights. It is given a detailed treatment by Har-kins [1952] and is not repeated here. For thepurposes of the present discussion, it suffices tooutline the main features of this model.

In the classical model of wetting, the surfaceof the solid is taken to be invariant as a liquiddroplet spreads over it. That is to say, the reactionbetween the liquid and the solid componentsacross their common interface is considered to

be negligible. It is also assumed that the com-position and other characteristics of the solid andliquid components, likewise, do not change withtime. This assumption is not generally valid, asis shown in this section.

Modeling of self-alignment and other wettingand spreading processes can be done with a pro-gram entitled “Surface Evolver.” A search of theWorld Wide Web using this as the keywordshould identify a site from which the softwarecan be obtained.

1.2.1 Surface Energy andSurface Tension

The concepts of surface energy and surfacetension are briefly reviewed in this section. Fig-ure 1.6 provides a simplified representation ofthe atomic structure of a solid close to one of itsfree surfaces. The atom at positionA, in the bulkof the solid, has a balanced array of neighboringatoms, whereas atom B at the surface of the solidis lacking in neighbors above it, apart from theoccasional vapor molecule and, therefore, it hassome unsaturated bonds.

The potential energy of atoms at the free sur-face, such as B, is higher than the energy ofatoms within the bulk of the solid, such as A, bythe energy of the unsaturated bonds. The aggre-gate of this excess energy that is possessed byatoms in the vicinity of the free surface consti-tutes the surface energy of the solid. In a similarmanner, a liquid also possesses a surface energy,which is directly manifested in the tendency todraw up into drops. If small, the droplets are

Fig. 1.6 Simplified diagram of surface energies. Atom B, atthe surface, has unsaturated bonds and thus a higher

energy than atom A. This difference in energy is the origin ofsurface energy �SV.


perfect spheres. Because a sphere has the small-est surface-to-volume ratio of any shape, it isclear that the surface energy of a liquid is greaterthan its volume energy. In the classical model,when a liquid spreads over a surface, the volumeremains constant, because evaporation and re-action with the substrate are excluded. There-fore, only surface-energy changes must be con-sidered.

A surface of a liquid acts like an elastic skincovering the volume; in other words, the sur-face is in a state of tension. The tensile force(F), known as surface tension (�), is defined asthe force acting at right angles to a line of unitlength (L) drawn in the surface. The relation-ship between surface tension and surface en-ergy under specific conditions can be seen asfollows.

Consider a liquid film of length L and widthW. Apply a force F at a barrier AB, as shown inFig. 1.7, parallel to one surface of the film, so asto extend the liquid film a distance x.

The increase in area of the film is x • L.The work done in obtaining this increase is the

mathematical product of the force applied timesthe distance moved, or F • x.

The work done by the liquid film in opposingthe increase in area, under isothermal conditions(i.e., constant temperature), is 2 • � • x • L, where� is the surface tension force acting on eachsurface at the prescribed temperature.

At a fixed temperature (under isothermal con-ditions):

Fx � 2�xL

Rearranging, F/L � 2� or F/L � � for eachsurface.

Thus, surface energy is equivalent to surfacetension under isothermal conditions. In the mod-ern metric or International System of Units (SI),the unit of surface energy is joule per squaremeter (J/m2) and that of surface tension is new-ton per meter (N/m). Because these parametersare properties of an interface (e.g., between liq-uid and air), surface energy and tension must bedefined with reference to the appropriate pair ofmaterials that meet at the interface, and the testconditions, such as temperature and atmosphere,also must be specified.

1.2.2 Wetting and Contact AngleAccording to the classical model of wetting,

the liquid will spread over a solid surface untilthe three surface tensions—between the liquiddroplet and the solid substrate, the liquid dropletand the atmosphere, and the substrate and theatmosphere—are in balance as shown in Fig. 1.8.

According to the balance of forces:

�SL � �SV � �LV cos � (Eq 1.1)

where �SL is the surface tension between thesolid and liquid, �LV is the surface tension be-tween the liquid and vapor, �SV is the surfacetension between solid and vapor, and � is thecontact angle of the liquid droplet on the solidsurface.

Equation 1.1, known as the wetting orYoung’sequation, shows that � < 90° corresponds to thecondition�SV>�SL.This imbalanceinsurfaceten-sion (i.e., surface energy) provides the drivingforce for the spreadingof liquidover the solid sur-face and diminution of the unwetted surface area.

The contact angle � provides a measure of thequality of wetting. Thus, if 90° < � < 180°, somewetting is said to occur, but a liquid droplet willnot spread on the surface with which it is incontact. On the other hand, if � < 90°, a liquiddroplet will wet the substrate and also spread

Fig. 1.7 Diagram used to explain the relationship betweensurface energy and surface tension

Fig. 1.8 Surface tension forces acting when a liquid dropletwets a solid surface, according to the classicalmodel


over an area defined by the contact angle �.Clearly, the area of spreading will increase withdecreasing contact angle. For further details ofthe interrelationship between these two param-eters, refer to the Appendix A1.2.

Rewriting Eq 1.1 in terms of cos �:

cos� ��SV � �SL

�LV

Thus, wetting is improved by decreasing � ascos� increases; that is, as � approaches 0. Cos �may be maximized by:

• Increasing �SV• Decreasing �SL• Decreasing �LV

The term �SV can be maximized for a givensolid by cleaning the surfaces. The presence ofadsorbed material, such as water vapor, dust, andother nonmetallic surface films on a metal sur-face, markedly reduces �SV and correspondinglyraises the contact angle �. Therefore, it is im-portant in soldering operations that joint surfacesare clean andmetallic—hence the need for fluxesor protective atmospheres to achieve and thensustain this condition.

The term �SL is a constant at a fixed tempera-ture for a particular solid-liquid combination, ac-cording to the classical model of wetting. Thisparameter can be reduced by changing the com-position of the materials system, as can be seenfrom Fig. 1.9.

However, changing the composition is not usu-ally easy to achieve in practice because compo-nent materials are specified to fulfill certain otherfunctional requirements. Fortunately, �SL is

highly temperature dependent and usually de-creases rapidly with increasing temperature[Schwartz 1987, Table 1.4], thereby providing aready means of controlling spreading.

The term �LV is a constant at a fixed tempera-ture and pressure for a particular liquid-vaporcombination, but can be varied by altering thecomposition and pressure of the atmosphere. Al-though the composition of the atmosphere usedfor the joining operation is known to affect thecontact angle, in practice it is often easier topromote spreading by reducing the pressure ofthe atmosphere. This is one of the reasons for thepopularity of vacuum-based joining processes.

In general, the relative magnitudes of the sur-face energies are �SV > �SL > �LV. For waterwetting on mica, subjected to an atmosphere ofwater vapor, the following values have beenmea-sured [Tabor 1969]:

�SV � 0.183 N/m

�SL � 0.107 N/m

�LV � 0.073 N/m

Thus, cos � � (0.183 – 0.107)/0.073 � 1 (withinthe limits of experimental error) and the contactangle, � � 0°.

The surface energies of pure metals correlatequite well with their melting points. This is to beexpected because the refractoriness of metals re-flects the strength of the bonds between adjacentatoms in the lattice, and the asymmetry betweenthis and a free atom is responsible for surfaceenergy and tension.

Fig. 1.9 Wetting angle of lead-tin solder on copper at 10 ºCabove the melting point, 1 min after reflow using

rosin mildly activated (RMA) flux, as a function of lead concen-tration. Adapted from Liu and Tu [1998]

Table 1.4 Surface roughness (Ra) of cold-rolledcopper after sanding with wet silicon carbidepaper or polishing with a colloidal suspensionof alumina in water

AbrasiveNominal particle

size, μmRa obtained on

cold-rolled copper, μm

80 grit 200 2.2240 grit 63 0.95400 grit 23 0.511200 grit 5 0.23Polishing alumina 0.05 0.012

For comparison, copper surfaces on electronic component leads usually have anoriented Ra of approximately 0.1 μm (4 μin.).


It is possible to calculate the surface tensionof solder alloys from thermodynamic prin-ciples using data for the pure metals. It mostlyvaries as an essentially linear relationship be-tween the values for the two pure metals. Sol-ders that exhibit poor spreading even at largesuperheats above the liquidus temperature, suchas Sn-40In and Sn-65Bi, have similar surfacetension to other solders, but the wetting is con-trolled by the reaction kinetics at the solder/substrate interface, which are less favorable[Park et al. 2000].

The wetting equation (Eq 1.1) applies whenthe liquid is practically insoluble in the solidover which it spreads (i.e., the solubility is lessthan 0.1%). For binary metal systems wherethis condition is satisfied (e.g., tin-chromium),it has been shown that the wetting equation canbe reduced to:

cos � � 1 � k � TmS

TmL

�1 �where k is a constant equal to approximately 0.3,T m

S is the melting point of the solid metal, and T mL

is the melting point of the liquid metal. This ex-pressionhas beenverified experimentally [Eusta-thopoulos and Coudurier 1979]. Higher-ordermetal systems (ternary, quaternary, etc.) are con-siderablymorecomplex,and thewettingequationcannot be truncated to such a simple form.Amoresophisticated analysis of wetting that takes intoaccount the influence of certain microscopic fea-tures, including the influence of local defects andvan der Waals forces, is provided by de Gennes[1985].However, this is still a continuumanalysisand does not consider the local atomic environ-ment. Indeed, it has further been suggested thatYoung’s equation is only valid under certain spe-cial cases and there are some difficulties with thetheoretical definition of solid-surface tension[Xian 2000]. Further academic endeavor willhopefully resolve these issues.

So far, this chapter has idealized filler spreadover a single surface. In a joint there are alwaystwo facing surfaces. If both contact angles areless than 90°, the surface energies will give riseto a positive capillary force that will act to fill thejoint. For a pair of vertical parallel plates D mmapart and partly immersed in a liquid, the cap-illary force per mm length of joint is equal to2�LV cos �. Under this force, the liquid will rise

to an equilibrium height h where the capillaryforce balances the hydrostatic force (as shown inFig. 1.10), such that:

h �2�LV cos �

g D(Eq 1.2)

where is the density of the liquid and g is theacceleration due to gravity.

As might be expected, experimental assess-ment of capillary rise of solders reveals that cap-illary rise is less than predicted by theory, al-though the general principles of Eq 1.2 aresubstantiated. Meniscus rise is usually greatestfor solders that exhibited the lowest contact angleand surface tension and in the narrowest gaps.However, for many solders the correlation withgap width is weak. This has been studied andultimately was attributed to be due to a rapidincrease in voids in the joint as the gap width wasdecreased progressively (see Chapter 4, section4.3.1). Indeed, in practical terms capillary flowin narrow gaps is largely dictated by the efficacyof the flux/filler combination used [Vianco andRejent 1997].

The actual situation in soldering is much morecomplex than that represented by Eq 1.2 and theclassical wetting model. The irreversible natureof spreading and the time dependence of contactangle that is commonly observed are at variancewith this simple model. These and other depar-tures from the classical model occur because the

Fig. 1.10 Rise of a liquid between two parallel plates bycapillary force


joining process almost invariably involves a de-gree of chemical reaction between the fillermetaland the solid surface, which is neglected in theconventional model. This is clearly demon-strated in a study [Schwartz 1987] that showedthat the contact angle for various liquidmetals onfreshly cleaned beryllium generally decreasedwith time, over a timescale of several minutes,at a fixed temperature. Predictably, perhaps, itwas also found that the contact angle decreasedwith increasing temperature and the atmospherein which the test was conducted also made adifference.

Reactions between a filler metal and the sub-strate often result in dissolution of the surface ofthe substrate; this process usually leads to thegrowth of new phases. Frequently, these phasesare intermetallic compounds that either appeardistributed throughout the joint or form as layersadjacent to the surface of the solid substrate.

The energy of formation of an intermetalliclayer by reaction between a molten filler and asolid substrate has been calculated by Yost andRomig [1988] and Wang and Conrad [1995].The energy of formation considered is the ther-modynamic function known as the Gibbs freeenergy. This function and its properties arebriefly explained in the Appendix A3.1. In or-der to simplify the analysis, Yost and Romiglimited their consideration to the clean sur-faces of pure metals, wetted by liquids of el-emental metals, in the absence of fluxes, toform binary interfacial phases. It was demon-strated that the free energy of formation of in-termetallic phases by reaction of liquid anti-mony, cadmium, and tin with solid copper wasapproximately two orders of magnitude largerthan the energy release created by the surface-energy imbalance during the advance of aspreading solder droplet, which is exclusivelyconsidered in the classical model. Therefore, inthese cases, and probably more generally insoldering processes, the Gibbs free energychange that occurs on reaction by a filler withthe substrate is demonstrably the dominantdriving force for wetting. Empirical evidencefor this is provided, for example, by the factthat the measured contact angle of molten ger-manium on silicon carbide at 1430 °C (2600°F) is approximately 120°, whereas that ofmolten silicon on this ceramic at the same tem-perature is 38° [Li and Hausner 1991]. Thesubstantial difference in the two contact anglescannot be accounted for by the difference in

�LV in the wetting equation (Eq 1.1). It canonly be due to the greater intersolubility of sili-con with silicon carbide. This example clearlydemonstrates that the simple classical wettingequation cannot be relied on for a quantitativedescription of wetting, contact angle, or spread-ing. A more direct example is provided by Fig.1.11. In this simple system of copper-siliconbraze wetted onto graphite, the final contactangle is insensitive to alloy composition, butthe rate of attainment of equilibrium wetting isdirectly related to the concentration of silicon,which is the active ingredient in the braze.

Modifications have been proposed to incor-porate the Gibbs free energy change accompa-nying metallurgical reaction into the classicalwetting equation by adding additional terms. Inparticular, the following equation has been de-veloped for the contact angle in reactive wetting[Kritsalis, Coudurier, and Eustathopoulos 1991;Laurent, Chatain, and Eustathopoulos 1991]:

cos � � cos �0 ��SL � �SL

0

�LV

�Gr

�LV

(Eq 1.3)

where �SL is the solid-liquid interfacial energyafter reaction, �SL

0 is the interfacial energy beforereaction, �0 is the contact angle before reaction,and Gr is the Gibbs free energy of the reaction.Equation 1.3 is probably more of theoretical in-terest than practical value because its use pre-supposes knowledge not only of the Gibbs freeenergy of reaction, but also values of the beforeand after contact angle or interfacial energy.

The effect of metallurgical interaction be-tween filler and the component material in pro-

Fig. 1.11 Contact angle of copper-silicon brazes of differentcomposition on vitreous carbon substrates dem-

onstrating the effect of driving force of alloying on wetting rateand the dependence of the equilibrium wetting angle on thereaction product. Adapted from Landry, Rado, and Eustathopou-los [1996]


moting wetting is exploited in active filler met-als: the addition of a small fraction of a reactivemetal such as titanium, hafnium, or zirconium tofillers enables them to wet and spread over ce-ramic materials. In this instance, wetting of andreaction with the ceramic are inextricably linked.Activated filler alloys are discussed further inChapter 4, section 4.1.2.2 and the planned com-panion volume Principles of Brazing.

Although a low contact angle is used as anindex for judging the quality of wetting, there aresituations where higher contact angles are pre-ferred. This is illustrated in Fig. 1.12, whichshows two joints, one between two componentsurfaces of unequal area and the other betweencomponent surfaces that entirely correspond. Inthe first case, a low contact angle serves to forma gentle concave fillet, which enhances the me-chanical properties of the joint. In the other con-figuration, a low contact angle encourages theformation of a neck in the joint, which can be asource of weakness. A contact angle close to 90°will eliminate this problem.

It is usually presumed that a substrate surfaceis perfectly wettable, or at least will be when theflux has had sufficient time to perform its “clean-ing” action. Sometimes, however, the majorityof the surface area will be wettable and the re-mainder covered with an array of nonwettablepatches. This can be due to inadequate surfacepreparation or an incorrect choice of flux or pro-cess conditions, or it is simply an inherent featureof the substratematerial.An example of the latteris some of the new generation of metal-matrixcomposites that comprise a finely dispersed mix-ture of a metal and powder of refractory com-pounds such as Be/BeO and Au/TiN. The neteffect of these nonwettable patches is to causelocal impediments to wetting and spreading andan increase in the effective contact angle, asshown in Fig. 1.13.

A further point to be aware of in connectionwith wetting is that a situation can arise wherethe molten filler is physically prevented fromachieving its equilibrium contact angle, as, forexample, when a solder droplet is confined to a

Fig. 1.12 Effect of contact angle on fillet formation and joint filling. Low contact angles tend to be preferred when external filletscan form. In other geometries, higher contact angles result in lower stress concentrations.


small metallized area. This is commonly en-countered on electronic circuit boards, where sol-der droplets are constrained to individual metalpads. In this situation, the pad is often too narrowto accommodate the spherical metal cap thatwould form if this restriction did not apply. Theenforced wetting angle imposes a pressure on thedroplet that is often adequate to cause the solderto flow along the length of the conductor thatleads away from the pad and, more seriously, tolift and flow under the solder resist that sur-rounds the pad. An analysis of the pressure aris-ing from a nonequilibrium contact angle, usingthe classical wetting model, is given by KleinWassink [1989].

1.2.3 Fluid FlowThewetting equation determines the degree of

wetting for a given liquid-solid combination, butwill not provide information on the rate of wet-ting. Knowledge of the contact angle(s) enablesthe surface energy to be determined and hencethe force that acts to fill the joint gap with liquid.The liquid will flow into the joint under this forceat a rate that is governed by its viscosity. Simplefluid-flow theory assumes that:

• There is no interaction between the liquid andthe solid surfaces with which it is in contact.

• All surfaces are smooth and perfectly clean.• Flow is laminar, not turbulent.

For a detailed treatment of this subject, the readeris referred to a paper by Milner [1958]. Thischapter merely quotes the expression (given asEq 8 in Milner’s paper) for the volume rate of

liquid flow, dV/dt, between a pair of horizontalparallel plates, length l, separated a distance D,under a pressure P per unit area transverse to theplates. The viscosity of the liquid is �.

dV

dt�

PD3

12�l

It is assumed that the liquid front will advanceat a rate (dl/dt) equal to themean velocity of flow,that is:

dl

dt� � 1

D� � dV

dt� �

PD2

12�l

From the wetting equation (Eq 1.1), under iso-thermal conditions the change in surface energyas a unit area of a surface becomes wetted by theliquid is:

�SV � �SL � �LV cos �

Therefore, the change in surface energywhen thepair of parallel plates becomes wetted is:

2l (�SV � �SL ) � 2l �LV cos �

It follows that the force acting on the liquid tocause it to wet the plates is:

F �2l �LV cos �

l

Fig. 1.13 Effect of nonwettable surface features on the contact angle of solder on copper. Data of Yost, Hosking, and Frear [1993]augmented by the authors. Lead-tin solder wetted onto a copper surface containing embedded nonwettable particles

10–20 μm (400–800 μin.) in diameter (RMA flux, 180 °C, or 356 °F).


so that the pressure is:

P �2�LV cos �

D

and the velocity of flow of the liquid into thespace between two parallel surfaces, of separa-tion D, according to this simple model is givenby:

dl

dt�

�LVD cos �

6�l(Eq 1.4)

Equation 1.4 shows that the rate of liquid flowincreases when:

• The liquid-vapor surface tension, �LV, in-creases.

• The joint gap, D, increases.• The contact angle � decreases.• Filler metal viscosity is low.

Andrade [1952] derived an empirical formularelating viscosity, when molten, to the molecularweight of metals (in SI units):

�m � 1.65 � 10�7 Tm0.5 A0.5

V 2/3

where �m is the viscosity at the melting point ofthe metal, Tm is the absolute melting point, V isthe molar volume, and A is the atomic weight ofthe metal. By assuming limited solubility be-tween the constituents in an alloy and applyingthe rule of mixtures, it is thereby possible toprovide an estimate of the theoretical viscosity ofsolder.

Rates of flow calculated from Eq 1.4 for mol-ten solders in joints 50 μm (2000 μin.) wide aretypically 0.3 to 0.7 m/s (1 to 2.3 ft/s). In otherwords, a joint 5 mm (0.2 in.) long will be filledin a time of the order of 0.01 s. This implies thatjoint filling by the molten solder occurs virtuallyinstantaneously and that transient effects asso-ciated with fluid flow can generally be neglectedin joining processes. De Gennes [1985] offers amore developed model of the dynamics of liquidspreading, in which the surface-energy drivingforce is opposed by viscous drag and surfaceirregularities. Joint filling times of the order of0.1 s are routinely measured on instruments used

for determining solderability. It should be notedthat, although the rate of filling is proportional tothe joint gap D, the driving force for filling, ac-cording to the classical model, is inversely pro-portional to D; that is, these two aspects of fillingact in opposition.

This simple model needs to be modified insituationswhere interfacial reaction occurswhileliquid spreading is proceeding. Models that havebeen tentatively proposed for this situation havebeen reviewed byMeier, Javernick, and Edwards[1999]. Currently, the lack of relevant data onreaction-rate kinetics, interfacial energy beforeand after reaction, and diffusion hampers a morecomplete understanding of spreading, and alsowetting, of molten fillers, especially where in-terfacial reaction with solid components is sig-nificant. However, much can be learned fromempirical observations, as shown in the follow-ing section.

1.2.4 Filler Spreading CharacteristicsMolten filler metals do not all have the same

spreading characteristics, although, with few ex-ceptions, the degree of spread over an “ideal”substrate increases as the temperature is raisedand the environment is made more reducing. Inthis context, an “ideal” substrate, suitable forreference purposes, needs to be defined. This isunderstood to possess a perfectly clean metalsurface that is highly wettable by the filler metalunder consideration, but with which it does notsignificantly alloy.Any alloying reactionswill behighly specific to the combination of materials inquestion, so that the substrate will lose its idealcharacteristics.

An example of a substrate that approximatesthe ideal, and that has been used by the authorsin comparative soldering assessments, com-prises a flat glass plate sputter coatedwith 0.1 μm(4 μin.) of chromium and overlaid with 0.1 μm(4 μin.) of gold. The chromium represents ametalthat is essentially insoluble in most solders, andthe gold layer provides this reactive metal withprotection against oxidation. The gold layer issufficiently thin to not significantly alter the com-position of a solder pellet as it spreads over thesubstrate [Humpston and Jacobson 1990].

Eutectic composition alloys are often re-garded as having the best spreading character-istics, and this is frequently one of the reasonscited for their selection in preference to hypo-eutectic and hypereutectic compositions. The su-perior spreading of alloys of eutectic composi-


tion in comparison with off-eutectic alloys of thesame system, which is often observed, can beexplained by the different melting characteristicsin the two cases. An alloy of eutectic composi-tion melts instantly. Spreading of the molten al-loy is then driven by interaction with the sub-strate [Ambrose, Nicholas, and Stoneham 1992].In the case of a noneutectic filler metal, melting,wetting, and spreading commences before thealloy is entirely molten and it will tend to besomewhat viscous. Under these conditions,movement of the filler will be relatively sluggish.By the time the alloy is completely molten, thefiller will have partly alloyed with the substrate,and the driving force for spreading will havebeen diminished. Eutectic composition alloysalso have lower viscosity than adjacent compo-sitions when completely molten; further detailsare given in Chapter 2, section 2.3.1.

Whether or not the filler alloy is of eutecticcomposition is of much less importance to thephenomenon of spreading than the compositionper se. The spreading of a filler metal dependsgreatly on the elemental constituents present andtheir relative proportions. The authors have com-pared the spreading characteristics, as a functionof excess temperature above the melting point(“superheat”), of all combinationsof theelementsbismuth, indium, lead, silver, and tinwhenusedaseutectic solders on “ideal” substrates [Humpstonand Jacobson 1990]. The results presented in Fig.1.14 show that the area of spreading increases atan accelerating rate as a function of the excesstemperature above themelting point of the solder.Furthermore, this study has demonstrated thatthere is a consistent ranking order for these ele-ments in their ability to promote spreading—namely, tin > lead > silver > indium > bismuth.This ranking order is maintained even for ternaryand quaternary solders and when applied to arange of substrates, in air, using mild fluxes.

Although high fluidity of a filler metal is a de-sirablepropertywhen it is required toflowinto thejoint gapof a heated assemblyby capillary action,it is not quite so important when the preferredmethod of applying the filler is to sandwich a thinfoil preform between the components, which arethen joined together in an appropriate heatingcycle.For this typeofconfiguration, ahighdegreeof spreading is detrimental to joint filling, as thefiller tends to flow out of the joint. Placement ofthe filler metal and its influence on joint filling isdiscussed in Chapter 4, section 4.3.1.1.

In a vacuum or neutral protective atmosphere,the spreading of a filler metal will tend to be in-

ferior to that obtained in air in the presence of achemicalflux.This is tobeexpected inviewof thelimited effectiveness of these environments: nei-ther a vacuumnor a protective atmosphere is usu-ally capable of removing oxides that form on thesurface of components or the filler while exposedto air before the joining operation. In both cases,the spreading is inferior to that achieved in thepresence of an active flux that can remove the sur-face oxide [Humpston and Jacobson 1991].

Detailed investigation reveals that even an os-tensibly simple parameter such as contact angleexhibits somewhat complex behavior. Figure1.15 shows the contact angle of Pb-60Sn solderon copper at 194 °C (381 °F), protected by achemically inert flux, as a function of wettingtime. The results indicate that there are at leastfour distinct stages of wetting and spreading.During the first 10 s of melting, the solder formsa spherical cap and there is subsequent rapidspreading with a corresponding decrease in thecontact angle. The contact angle then tempo-rarily stabilizes, as a dynamic balance is struckamong the growth rate of interfacial intermetal-lic compounds, the diffusion rate in the moltenpool of liquid, and the efficacy of the flux incleaning the substrate surface. This situation per-sists for about 500 s. Thereafter, a further re-duction in contact angle occurs. This is thoughtto be associated with a progressive change in thecomposition of the solder resulting in a suddenchange in the intermetallic compounds formed inthe halo at the edge of the solder pool and hencea change, in this case a further decrease, in thewetting angle. Finally, after many minutes, thecontact angle reaches a settled value as the solderpool becomes saturated with the substrate. Themolten liquid then commences isothermal freez-ing as the solidus temperature progressively in-creases, owing to the alloying with the substrate[Wang and Conrad 1995].

Some attempt has been made to undertake atheoretical analysis of the kinetics of spreading ofa molten metal over a wettable solid surface. Thecurrent theoreticalapproachconsiders thespread-ing of an inert sessile drop on a smooth and per-fectly wetted substrate as a balance between sur-face-energy drive and viscosity impedance [deGennes1985].However, comparisonof this theo-retical model with practical experience reveals anumber of flaws with the model, not the least isthat measured flow rates are in the region of fourorders of magnitude slower than predicted bytheory (see Fig. 1.16). These discrepancies aremostly due to the added metallurgical and


Fig. 1.14 Spread characteristics of binary solder alloys on an “ideal” substrate as a function of excess temperature above the meltingpoint. The substrate is a flat, microscope slide, sputter metallized with 0.1 μm (4 μin.) of chromium overlaid with 0.1 μm

(4 μin.) of gold. Spread ratio is defined in Appendix A1.2.

physical complexity in filler metal wetting andspreading as discussed earlier and are among thesimplifying assumptions of this model.

Nevertheless, the de Gennes model does pre-dict some interesting dependencies of spreading.First, the initial spreading ofmolten fillermetal isdescribed by the imbalance between Young’sforces and viscous damping.Thismodel also pre-dicts a relative insensitivityof spreading toexcesstemperature in filler/substrate combinations thatwet well. Continued research in this area mayachieve a complete mathematical description ofwetting and spreading by filler metals that takesinto account isothermal solidification and thephysical and chemical state of the surface [Am-brose, Nicholas, and Stoneham 1993].

1.2.5 SurfaceRoughness of Components

The roughness of joint surfaces can have asignificant effect on both the wetting and spread-

ing behavior of a filler. Surface roughness re-duces the effective contact angle �*, where �* isrelated to �, the contact angle for a perfectly flatsurface through the relation:

cos �* � r cos �

where:

r �Actual area of rough surface

Plan area

At the same time, by producing a networkof fine channels, the texturing may increase thecapillary force acting between the filler and thecomponent surfaces. Both phenomena will tendto aid spreading. A directionally oriented sur-face texture promotes preferential flow parallelto the channeling [Nicholas and Crispin 1986].

From surface-energy calculations it is possibleto show that if the instantaneous contact angle ofthe molten filler is less than the surface angle(i.e., the root angle of V-shaped valleys), thenprofuse wetting tends to occur along the valleys.This is a frequent observation and, indeed, rep-resents a problem when soldering to rough ma-chined surfaces in that the filler does not spreaduniformly in all directions.

Exactly the same situation pertains to mi-croscopically rough surfaces. An example of amicroscopically rough surface is thick electro-plated copper that has been deposited withoutbrighteners (surface leveling agents) or a steelsurface that has been etched so as to furrowout the grain boundaries. The resulting surfacecomprises roughly spherical nodules, whichtherefore have a continuous network of valleys

Fig. 1.16 Comparison between measured and predicted rates of spreading by molten solder. The large discrepancy arises becausethe models are based on fluid flow and do not take into account the metallurgical driving force for spreading.

Fig. 1.15 Contact angle of lead-tin solder on copper as afunction of wetting time, using an inert flux and

low superheat. There are four distinct stages of wetting, the lastbeing the equilibrium contact angle that is obtained using moretypical process parameters.


between the nodule peaks. Profilometer mea-surements indicate that the many valleys havecontact angles in the region of 15° and aretherefore capable of enhancing spreading. Thedifference here is that rather than a contact linethat undulates over a rough surface, the wet-ting front has a fractal, or lacelike character.The solder wets the channels between thepeaks and gives the wetted solder an extensivehalo ahead of the main molten pool [Yost,Michael, and Eisenmann 1995].

Where capillary enhancement of spreading isrequired, the surface texture should be as jaggedas possible. A surface prepared by grit blastingor abrading with silicon carbide impregnated pa-per is therefore preferable to a shot-peened sur-face. The reason for this is as follows. Sharpreentrant angles that exist on jagged surfaces co-incide with sudden changes in the crystallo-graphic orientation of the exposed parent metal.Theadhesionofnativeoxidesat thesemicrostruc-tural discontinuities will tend to be relativelyweak and provide sites at which the oxide layercan be more readily undermined or penetrated.

There is a limit to the roughness of surface thatcan be used to promote spreading by a moltenfiller. If the texturing is too deep, then capillarydams can be formed and these will impede thespreading of the filler metal [Funk and Udin1952].Another factor that shouldbeconsidered inconnectionwith texturing is the extent of alloyingbetween the filler and the parent material. For ex-

ample,whenlead-tineutecticsolder isused to joingold-coated components, there is a limit of about4%to theconcentrationofgold thatcanbeaccom-modatedbeforethesolder isembrittled(seeChap-ter 2, section 2.3.2). If the solder spreads over agold-coated surface, the critical thickness of thegold coatingwill be reduced by a factor related tothe surface roughness, r, and which must be con-sideredwhen calculating the total volume of goldto be applied corresponding to a given plan areaof spread. Table 1.4 indicates values of surfaceroughness thatcanbeobtainedbyabrasionofcop-per by various means.

Attempts have been made to improve jointfilling by introducing capillarity enhancers to thejoint gap. Such enhancers include finely dividedpowders and fine meshes that are wetted by thefiller but that are effectively inert. This type ofapproach has been explored by the authors andhas not been found to radically improve jointfilling. A volume fraction of powder that wascalculated to significantly increase capillaryforces had an adverse effect on the fluidity of thefiller metal. On the other hand, meshes providedstable traps for air and evolved gases in the joint.This led to the formation of an array of voidscorresponding to each aperture in the mesh, asrevealed by a high-resolution radiograph of ajoint with a No. 400 gauze that was solderedusing a foil of Ag-96Sn alloy (Fig. 1.17).

The importance of surface roughness is ex-emplified by a recent study of the fracture tough-

Fig. 1.17 Radiograph of a 50 mm (2 in.) diam component soldered using two 50 μm (2000 μin.) thick foils of Ag-96Sn solder, inhigh vacuum and incorporating a No. 400 gauze. (b) High-resolution radiograph that reveals the true nature of the joint

filling in (a), with a void present at the center of each aperture in the mesh. Magnification: 640�


ness of joints made to copper components usingAg-96Sn solder. It was found that the only sig-nificant variable was surface roughness. Jointtoughness was found to be largely independentof joint thickness and soldering time for limitedranges of these variables (60–200 s and 150–400μm, or 6–16 mils, respectively). The response tovariation in surface roughness is reproduced inFig. 1.18. The low fracture toughness at inter-mediate roughness occurs when adjacent sites ofintermetallic compound growth meet at unfa-vorable crystallographic orientations [Stroms-wold, Pratt, and Quesnel 1994].

1.2.6 Dissolution ofParent Materials andIntermetallic Growth

It is frequently observed that a filler metal willcontinue to spread beyond an initially wettedsurface area over an extended period of time(>10 s), which would not be expected from clas-sical fluid-flow theory. Clearly, classical expres-sions for fluid flow, exemplified by Eq 1.4, do notstrictly apply in such cases. Indeed, this type offlow can usually be associated with solid-liquidinterfacial reactions, which are neglected in themodel described inMilner’s paper [1958].Wherejoint filling is sluggish because of reactions oc-curring between the filler and the solid surface,increasing the temperature to reduce the viscos-ity of the molten filler is unlikely to enhance

filling, because the reactions that are occurringtransverse to the flow directions will accelerate[Tunca, Delamore, and Smith 1990].

Dissolution of the substrate and resultinggrowth of intermetallic compounds both followArrhenius-type rate relationships, represented by:

Rate � exp ��Q

kT�

where Q is an activation energy that character-izes the reaction taking place at temperature T (indegrees Kelvin) and k is the Boltzmann constant.The alternative of increasing the joint gap is notusually an option because this is likely to lead toa reduction in joint filling and/or joint strength,as discussed in Chapter 4, section 4.3. The so-lution then is to change the materials system;several means by which this can be achievedwithout changing the parent materials are de-scribed in Chapter 4, section 4.1.

Interfacial reactions are important, not only indetermining the flow characteristics of the fillerand its wetting behavior, but also the propertiesof the resulting joints. When a molten filler wetsthe parentmaterials, there is normally some inter-solubility between them. It is usually manifestedas dissolution of the surfaces of the parent ma-terials in the joint region and the formation ofnew phases at either interface between the parentmaterials and the molten filler or within the filleritself when it solidifies. The effects of dissolutionof the parent materials and compound formationon joints are discussed in detail in Chapter 2,section 2.3.

The rate of dissolution of a solid metal in amolten metal is described by Weeks and Gurin-sky [1958, p 106–161] and Tunca, Delamore,and Smith [1990]:

dC

dt�

K A (Cs � C )

V(Eq 1.5)

where C is the instantaneous concentration of thedissolved metal in the melt, Cs is the concen-tration limit of the dissolved metal in the melt atthe temperature of interest, t is the time, K is thedissolution rate constant, A is the wetted surfacearea, and V is the volume of the melt. This equa-tion is known as both the Nernst-Shchukarev and

Fig. 1.18 Effect of the surface roughness of copper substrateson the fracture toughness of jointsmadewith silver-

tin eutectic solder. It is worth noting that the joints under whichthe test joints were made are relatively extreme in terms of jointthickness (i.e., quantity of solder present) and soldering time, sothat the thickness of the copper-tin intermetallic layers formed islikely to be substantially thicker than encountered in normal sol-dering practice.


the Berthoud equation. In the integral form, Eq1.5 becomes:

C � Cs�1 � exp ��KAt

V �� (Eq 1.6)

assuming initial conditions of C � 0, t � 0.Equation 1.6 reflects the fact that, in general,

the concentration of dissolved metal in the mol-ten filler increases in an inverse exponentialman-ner with respect to time. That is, the dissolutionrate is initially very fast but then slows as theconcentration of the dissolved parent materialtends toward its saturation limit (i.e., equilib-rium), as shown in Fig. 1.19. Substituting mea-sured values into Eq 1.6 shows that, for a sol-dered joint of typical geometry, the equilibriumcondition is reached within seconds at the pro-cess temperature. Thus, it is possible to use anequilibrium phase diagram to predict the changein the composition of the filler metal that willoccur in typical joining operations and the as-sociated depth of erosion of the joint surfaces.Equilibrium phase diagrams and their use in sol-dering and brazing are considered more fully inChapter 2, section 2.3.

In some materials systems, the product of re-action between molten filler metal and the parentmaterials is a continuous layer of an intermetalliccompound over the joint interface. Once formed,

the rate of erosion greatly decreases because it isthen governed by the rate at which atoms of theparent material can diffuse through the solid in-termetallic compound. As a rough guide, solid-state diffusion processes are two orders of mag-nitude slower than solid-liquid reactions, and thuscontinued dissolution of the parent materials ef-fectively ceases, within the timescales of typicaljoining processes. Intermetallic growth will,however, continue throughout the life of the prod-uct, the practical implications of which are dis-cussed in Chapter 4, section 4.1.4.

1.2.7 Significance of the Joint Gap

The joint gap at the process temperature in-fluences both the joint filling and the mechanicalproperties of the resulting joint. The relationshipbetween joint dimensions and mechanical prop-erties is discussed in Chapter 4, section 4.3 andin the planned companion volume Principles ofBrazing. In summary, the thinner a joint is, thegreater its load-bearing capability tends to be,until a limiting condition is reached.

Contact angle, surface tension, and viscosityall reduce with increasing temperature, makinggood joint filling in narrow joints more readilyachievable as the joining temperature is raisedabove the melting point of the filler metal. Alower practical limit to the joint gap is imposedby three factors:

• The need to provide a path for vapors toescape. Flux vapors evolved within the jointand pockets of air must be allowed to escape,if the formation of voids through gas entrap-ment is to be prevented (see Chapter 4, sec-tion 4.3.1.1). At the same time, any reducingagent needs to gain access to all joint surfacesand be present in sufficient concentration towork effectively.

• Reaction with the components. The metal-lurgical reaction that occurs between a mol-ten filler metal and the surfaces of the com-ponents can take one of two forms.

a. The surface region of the work piece haslimited solubility in the molten filler. Thisis the preferred situation. The dissolutionof metal from the surface of the compo-nents can result in either compound for-mation at the interface, which may prevent

Fig. 1.19 The concentration of a solidmetal in a liquidmetalwetted by it changes in an inverse exponential

manner with respect to time and is limited by the saturationconcentration of the solid constituent in the liquid at that tem-perature.


further dissolution, or alloying with thefiller, which will change its compositionand hence its melting point.

On the whole, solders tend to form in-terfacial compounds with parent materials,while brazes usually exhibit more exten-sive alloying between the materials. Thiscan be partly explained by the fact thatmost solder alloys are based on elementswith crystal structures that differ from thoseof most common parent metals. Conse-quently, intermetallic compounds tend toform in preference to solid solutions.

A reaction that depresses the meltingpoint of the filler metal is desirable fornarrow joints, because its fluidity will beenhanced by such a reaction at a constanttemperature.Areaction that raises themelt-ing point of the filler metal will tend toincrease its viscosity and can cause thefiller to solidify at the process temperaturebefore it has filled the entire joint. Widerjoints mitigate this effect because the al-loying will tend to be diluted.

b. Dissolution of the filler in the parent metal.In this situation, the volume of filler willshrinkas the reactionprogresses; therefore,a larger volume of filler metal accommo-dated in awider joint gap is again preferredand for similar reasons. However, absorp-tionof thefiller isgenerallyundesirable,be-cause its constituents will tend to penetrateinto the parent materials preferentiallyalonggrainboundaries,generallytothedet-riment of the mechanical properties of theassembly and sometimes resulting in em-brittlement and/or hot shortness.

• Control of the joint gap. The width of thejoint gap must be predictable and stableduring the bonding cycle. The size of thegap will be influenced by the coefficients ofthermal expansion of the respective com-ponents, and allowances need to be madefor different expansivities of the matingcomponents. The expansivities of a repre-sentative range of engineering materials atroom temperature (25 °C, or 77 °F) arelisted in Table 1.5. Temperature gradientsalong the joint must be considered from thesame viewpoint. Variations of joint gapshould be avoided wherever possible, asthis can have a serious effect in impedingflow of the filler by capillary action.

An upper practical limit to the joint gap isdetermined by:

• Mechanical properties of the joint. As thegap is increased, themechanical properties ofthe joint declines progressively to those ofthe bulk filler metal, which in the case ofsolders are particularly poor in relation tomost structural materials. This aspect is dis-cussed further in Chapter 4, section 4.3.3.

• Joint filling requirements. As noted in sec-tion 1.2.2 in this chapter, the capillary forcedecreases as the joint gap increases, and thiswill place a practical upper limit on the jointgap.At the same time, a sufficient quantity offiller must be supplied to the joint to entirelyfill it. Hydrostatic forces will promote theflow of low-viscosity filler metals out of widegap joints.

The optimal balance of these factors isachieved when the joint gap is about 10 to 100μm (400 to 4000 μin.), depending on the type ofreaction that occurs between the filler and thecomponent. This is substantiated by theoretical

Table 1.5 Typical thermal expansivities ofcommon engineering materials at normalambient temperatureMaterial Linear expansivity, 10�6/K

Polymers

Polymers, rubbers 150–300Polymers, semicrystalline 100–200Polymers, amorphous 50–100

Metals

Zinc alloys 25–30Aluminum alloys 20–23Copper alloys 16–19Stainless steels 15–17Iron alloys 13–15Nickel alloys 12–15Cast irons 10–13Titanium alloys 8–10Tungsten/molybdenum alloys 4–7Low-expansion alloys (Fe-Ni-base) 1–5Graphite 7–9

Ceramics

Ceramics, glass 6–10Ceramics, oxide 4–8Ceramics, porcelain/clay 3–7Ceramics, nitride/carbide 2–6Diamond/silica/carbon fiber –1 to 1

The values given are representative of the most widely used materials, rather thanprovide absolute limits for the different classes listed. The thermal expansivitywill depend not only on elemental composition but also on microstructure andtemper. Composite materials can have expansivities that effectively range be-tween those of the constituents and depend on the relative proportions of thematrix and reinforcement phases. To convert to customary units of 10�6/°F,multiply given values by 0.55556.


calculations of capillary force and viscous dragof liquid flow; see Fig. 1.20.

Generally,when components rest freely ononeanother and the assembly is heated until the filleris molten, the joint will tend to self-regulate towidths around 50 μm (2mils). Indeed, it has beendemonstrated that for afixed combinationoffillermetal, component materials, and process condi-tionsthejointgapwill tendtoafixedvaluespecificto the combination. This value must be deter-mined by experiment. If there is insufficient fillermetal to fill this gap then the joint will containvoids, or if too much filler is applied the excesswill spill out [Bakulin, Shorshorov, and Shapiro1992].Where thinner orwider joints are required,it is necessary to insert spacers (such as wires) ofthe desired width between the components and,for thin joints, to apply pressure during the bond-ing cycle to overcome the hydrostatic forces thatwill act to levitate the upper component.

1.2.8 The Strength of Metals

The purpose of making soldered joints is usu-ally to form a metallic bond between compo-nents. A fundamental question, therefore, is how

strong is the interface between the parent mate-rial and the filler metal in an ideal situation?

The cohesive strength of metals results fromattractive forces between the constituent atoms.Normally, each atom will occupy a physical lo-cation where the net force on it is zero.When thesolid metal is strained by the application of anexternal load, the atoms move from their equi-librium positions, and an opposing stress is set upin the metallic crystal. The attractive force be-tween atoms that share the same electron cloudincreases with the distance between them up toa maximum and thereafter decreases abruptly,when failure occurs. A perfect metal lattice willfail at this point by cleavage across the crystal-lographic plane because this is the region wherethe interatomic forces are weakest.

To a first approximation, the interatomic forceper unit area varies with interatomic separation,x, according to a sine wave with wavelength �,as shown in Fig. 1.21. The interatomic force perunit areamay then be represented by a sinewave:

� o sin 2�x

�(Eq 1.7a)

where o is the maximum theoretical strength.The work done per unit area in completely sepa-rated neighboring planes of atoms, which are anequilibrium distance apart (i.e., a � �/4), is then:

�0

�/2

dx � �0

�/2

o sin 2�x

�dx �

� o

�

Thiswork corresponds to the total surface energyof the two new surfaces created in the fracture,that is, 2�SV, where �SV is the surface energy perunit area of the solid.

Accordingly:

o �2��SV

�(Eq 1.7b)

Within the elastic range of strain, Hooke’s lawapplies, that is:

� � E�x

a

Fig. 1.20 Calculated time for molten tin and copper to flowup a perfectly wetted capillary [Nicholas 1989]

Fig. 1.21 Variation of interatomic force, per unit area, withdistance


Differentiating Eq 1.7(a) gives:

d

dx� o

2�

��cos 2�

x

��

At zero strain, that is, x � 0:

�d

dx�

x�0

� o2�

a

Hence:

o �E�

2�a(Eq 1.8)

From Eq 1.7(b) and 1.8:

� �2��SV

o

� o2�a

E

so that:

o � �E�SV

a�1/2

The theoretical fracture stress is about o /10for metals, although in practice strengths of met-als tend to be only one-tenth of this value (i.e., o /100), owing to the presence of lattice defectsand other discontinuities.

Possibly somewhat surprisingly, solderedjoints subject to simple mechanical stress willoften fail in a brittle manner. The reasons for thisare elaborated in Chapter 4, section 4.3.3. Inbrittle materials, failure takes place by the ex-tension of cracks that either preexist in the struc-ture or nucleate at lattice imperfections. Thestress to cause fracture can be deduced from Eq1.9 by replacing the denominator with c, wherec is the crack length, thus:

b � (E�SV�c)1/2

Since c is very much larger than a, the mechani-cal strength of a brittle material is low relative to

its theoretical strength. Only in special materialssuch as carbon fiber are the two values remotelycomparable.

In ductile metals, application of stress resultsin the movement of dislocations and other de-fects through the lattice of individual grains. Theinterfaces between grains are another regionwhere physicalmaterial transport and plastic flowtakes place. Failure occurs when the rate of in-crease in strength of the material due to workhardening falls below the rate of decrease in theload-bearing cross section resulting from theplastic flow.

The preceding discussion pertained to bulkmaterials, that is, the components and the fillermetal, when considered in isolation. In reality,the joint interfaceswill often be a source of voids,microcracks, local interfacial mismatch stresses,and brittle intermetallic layers. These featurestend to be a common source of joint weakness,and they should be minimized through judiciouschoice of filler/parent material and joining con-ditions.

1.3 The Design and Application ofSoldering Processes

A soldered joint is usually required to satisfya specific set of requirements. Most frequently,it must achieve a certain mechanical strength,which it must retain to the highest service tem-perature in the intended application. The jointmust also endure a particular service environ-ment, which may be corrosive, and it may haveto provide good electrical and thermal conduc-tance. In addition, the joint must be capable ofbeing formed in a cost-effective manner withoutdetriment to other parts of the assembly.

The principal aspects that need to be ad-dressed can be divided:

• The functional requirements of the applica-tion and themeans of satisfying these throughappropriate structural design

• The achievement of the specified assemblythrough successful processing

Each of these stages is examined in this section.

1.3.1 Functional Requirements andDesign Criteria

All soldered joints used inmanufactured prod-ucts must remain solid in service and retain the


associated components in fixed positions whensubjected to stress. These requirements are usu-ally satisfied by suitable design of the geometryand the metallurgy of the joint, but there are alsoother aspects to consider. Not the least amongthese is the fact that solders, in particular, areoften operating under conditions that are rela-tively at least as severe as those encountered injet engines [Plumbridge 1995]. Factors that havea bearing on the functional integrity of solderedjoints are discussed below.

1.3.1.1 Metallurgical Stability

For a joint to remain solid, in most cases themelting point (solidus temperature) of the fillermetal needs to exceed the peak temperature thatthe component is ever likely to experience. Thereare exceptions to this rule, which are discussedin Chapter 5, section 5.9. Because the strength ofall metals decreases rapidly as the melting pointis approached, the peak operating temperatureshould not exceed about 70%of themelting pointof the filler, in degrees Kelvin, if the joint isrequired to sustain a load.

The phases that form on solidification in asoldered joint are frequently unstable at elevatedservice temperatures. Instability of phasespresent in the joint at the service temperaturemay be undesirable, because it can affect its in-tegrity. The effects of continued reactions be-tween the filler and the components also must beconsidered, as explained in Chapters 2 and 3.Because most solders are softer than the mate-rials that are commonly joined, the mechanicalproperties of a joint are generally limited by thoseof the filler metal. An exception arises when thejoints are very thin and constrained between par-ent materials of high modulus, as described inChapter 4, section 4.3.3.

1.3.1.2 Mechanical Integrity

The durability of engineering and consumerproducts often depends on joints maintainingtheirmechanical integrity for the duration of theirexpected service life. The mechanical integrityof a soldered joint depends on a number of fac-tors, including:

• The mechanical properties of the bulk fillermetal (Chapter 5, sections 5.5 and 5.6)

• The joint geometry—namely area, width, andshape (Chapter 4, section 4.3)

• Themechanical properties of any new phasesformed in the joint by reaction between thefiller and the components, either during thejoining operation or subsequently in service(i.e., there is an interdependence with the mi-crostructure) (Chapter 4, section 4.1).

• The number, size, shape, and distribution ofvoids within the joint (Chapter 4, section 4.3)

• The quality of fillets formed between the fillerand the surface of the components at the edgeof the joint (i.e., their radius of curvature andextentofcontinuity) (Chapter4, section4.2.4)

The mechanical properties of joints, taking intoaccount the influence of joint geometry, are ex-tensively reviewed elsewhere. The reader is re-ferred to Schwartz [1987], Frear, Jones, andKins-man [1990], Brandon and Kaplan [1997],Nicholas [1998], and Manko [2002].

1.3.1.3 Environmental Durability

Joints are normally expected to be robust inrelation to the service environment. This mostcommonly involves exposure to corrosive gases,including sulfur dioxide and other constituentsof a polluted atmosphere; to moisture, perhapsladen with salt; and to variable temperature. Thecorrosion and stress-corrosion characteristics ofthe joint are then relevant. Corrosion mecha-nisms are generally very complex and specific toa given combination of materials, chemical en-vironment, and joint geometry. Therefore, eachsituation should be determined empirically.

The temperature of a joint can be shifted wellbeyond a normal ambient range, especially inaerospace applications and in situations whereheat is generatedwithin the assembly itself. Then,thermal fatigue and other changes to the metal-lurgical condition of the joint, such as the growthof phases, can occur, and these invariably affectthe properties of the joint. In other words, thereis interdependence between environmental sta-bility and microstructural stability. An appropri-ate choice of the materials combination usedshould enable these changes to be constrainedwithin predictable and acceptable limits. In thisregard, solders tend to be used at service tem-peratures that are proportionately very close totheir melting points, with respect to the thermo-dynamic reference point of absolute zero tem-perature (–273 °C, or –459 °F). Hence, they aremetallurgically unstable and microstructuralchanges take place readily.


1.3.1.4 Electrical andThermal Conductivity

In certain applications, soldered joints are re-quired to provide electrical and/or thermal con-tact between components. Generally, thin, well-filled soldered joints amply satisfy thisrequirement. Only in a few extreme situationsare the thermal and electrical properties of suchjoints close to the allowed limits. A case in pointis high-power silicon device assemblies, wherethe joint between the silicon device and themetalbacking plate is required to conduct away >1W/mm2 of thermal power. Here, it is crucial toensure that the joints are kept thin (<30 μm, or1200 μin.) and essentially void-free (<5% byvolume) [Humpston et al. 1992].

1.3.2 Processing Aspects

An important aspect that must be consideredwhen designing a joint is the practicality of theprocess involved.Among the relevant issues are:

• Jigging of the components• The form of the filler metal• Heating method• Temperature measurement• Joining atmosphere• Coatings applied to surfaces of components

(as necessary)• Cleaning treatments• Heat treatments prior to joining• Heating cycle of the joining operation• Postjoining treatments

1.3.2.1 Jigging of the Components

The components being joined normally mustbe held in the required configuration until thefiller metal has solidified. Even if the compo-nents can be preplaced without a fixture, theuse of some form of jig is still frequently ben-eficial to ensure that the components are notdisturbed by capillary forces originating fromthe molten filler metal. On the other hand, it isalso possible to exploit the capillary forces tomake the assembly self-align during the join-ing process; this is widely practiced in the fab-rication of electronic circuits using lightweightsurface-mounted components, for example,style 0201 and flip-chip assembly (see Chapter5, section 5.2).

The jig can be used to fulfill more than a hold-ing function. For example, it can serve as a heat

spreader, as a heat sink, or as a heat source. A jigshould be constructed from a nonporous materialto prevent contamination of the atmosphere sur-rounding the workpiece. Moreover, the jig ma-terial should not be wettable by the molten fillermetal, in case they come into contact throughaccidental spillage. Materials such as brassshould be avoided as zinc readily volatilizes.Care should be taken in designing jigs so as notto stress the constrained components throughthermal expansion mismatch.

A note of caution needs to be sounded whenjoining to stressed components, as this can lead tobrittle failureof thecomponents throughamecha-nism known as liquid metal embrittlement. Thedegradation in themechanicalpropertiesof anen-gineering steel, when stressed in tension andwet-ted by Pb-4Sn solder, has been studied by Wat-kins, Johnson, andBreyer [1975]. In practice, thisfailure mechanism is seldom encountered.

Graphite is often favored as a jig material. Itis inexpensive, easy to machine, a good thermalconductor (making it an effective heat spreader),is an absorber of radiant heat, and is not wettedby the majority of molten filler metals. Graphitealso has the merit of “mopping-up” oxygen in anoxidizing atmosphere to formCOandCO2.How-ever, if the oxygen partial pressure is alreadylow, then its role as a reducing agent is negli-gible. Care should be taken to ensure that thegraphite used for jigs is of a dense grade so thatit will be mechanically robust and have a lowporosity to minimize outgassing. The desorptionof water vapor, in particular, frequently deter-mines the quality of a gas atmosphere in thevicinity of the workpiece.

Jigs are sometimes used to apply a controlledpressure to a joint in an assembly. One compo-nent can then be deliberately and elastically dis-torted to bring it into close and uniform contactwith its mating part. This is an advantage whenvery narrow joints are required and when solid-state diffusion constitutes an important part ofthe joining process. Compressive loading on thejoint also aids expulsion of air and vapor fromthe joints, which are otherwise trapped in pock-ets and produce voids. Another application isoutlined in Fig. 1.22. Because this is a fluxlesssoldering process, the applied load serves a dualfunction. First, it helps puncture the oxide filmson the surfaces of the filler. Second, when the-filler metal melts, the applied force acts againstany dewetting capillary force of the liquid toensure flow. The combination of these two fac-tors leads to improved joint filling. Additional


considerations about fluxless joining processescan be found in Chapter 3, section 3.3.

1.3.2.2 Form of the Filler Metal

Filler metals are available in many differentforms. These include configurations that nor-mally can be produced from an ingot by me-chanical working—for example, wire, rings, andfoil. Such geometries are not restricted to ductilealloys. If the constituents are individually duc-tile, then the preform can be partitioned. Thismethod is discussed further in Chapter 4, section4.1.5. The development of rapid-solidificationprocesses has led to the availability of foils andwire of joining alloys that are inherently brittle.These foils are produced directly from the melt:the process involves forcing molten metalthrough a hole or slot onto a rapidly spinning,water-cooled, metal wheel. Figure 1.23 showssuch a strip casting process in operation, and Fig.1.24 gives some typical foils produced by thisroute. The high rate of heat extraction that occursin this process causes the molten metal to so-lidify almost immediately on striking the wheel,resulting in the formation of a strip of the alloywith a fine crystalline, or occasionally amor-phous,microstructure.The dimensions of the castmaterial can be controlled by varying the nozzledimensions, the ejection pressure, the speed ofrotation of the wheel, and other parameters of thecasting process [Jones 1982]. The refined mi-crostructure of the rapidly solidified alloys gen-erally improves strength and ductility compared

with the same alloys produced by conventionalcasting and mechanical working (see Fig. 1.25).A good example is the Bi-43Sn eutectic solder,which is normally brittle. However, when pre-pared as a foil by rapid solidification, the duc-tility of this alloy is comparable to that of othersolders.

Solders are also available as finely dividedpowders that can be mixed with a binder to forma paste capable of being screen printed onto asubstrate or applied to the workpiece, via a dis-penser, in an automated production line. How-ever, powders and pastes containing powdershave an extremely high ratio of surface area tovolume of filler, which generally produces highoxide fractions and in the absence of suitableprecautions would be detrimental to the qualityof the resulting joints. Paste manufacturers go togreat lengths to produce solder spheres with pol-ished surfaces and thin oxide skins specificallyfor this very reason.

Some of the more common filler metals areavailable as wire or rod that incorporates a flux.Most readers are probably familiar with flux-cored solder wire. In this form, the filler metalscan be readily used in air without any additionalprecautions.

In more specialized joining processes, the sol-der can be deposited as a coating on the com-ponents by screen printing, electroplating, andby vapor deposition techniques such as evapo-ration. Where it is not possible to deposit theactual alloy, sequential layers of the constituentelements can be applied. The former is generallypreferred because the melting point of an alloy

Fig. 1.22 (a) Sensor comprising three piezoelectric ceramic elements. These aremetallized and soldered onto ametallized substrate.The component is then encapsulated in a polymer to provide protection against the environment. (b) To ensure the

designed degree of acoustic coupling between the piezoelectric elements and the substrate, the soldered joints must be of a specifiedthickness. This was achieved using tungsten spacer wires in each joint and a spring-loaded jig to apply a compressive stress to eachelement during the process cycle. Also shown is the mask used to apply the metallization pattern to the substrate.


is well defined, whereas there is no guaranteethat melting will take place at the desired tem-perature in the case of the composite layers, un-less significant solid-state diffusion has occurredfirst to form the appropriate low-melting-pointphases.

The use of some form of preplaced filler metalhas a number of advantages. Most particularly,because the thickness and area of filler metal arepredetermined, the volume of molten filler maybe carefully controlled. Also, the number of freesurfaces is reduced from four (corresponding toa foil preform sandwiched in the joint) to justtwo, thereby considerably reducing the propor-tion of oxides and other impurities deriving fromexposed surfaces.

Tin-lead solder is the only solder that can bereadily applied to faying surfaces as an alloydirectly by electroplating. More recently, and

aided by the development of advanced pulse-platingmethods, coelectroplating of gold-tin sol-der has been developed. Prior to this, the solderhad to be realized as a deposit of gold, overlaidwith tin (the respective electropotential of theseelements makes it difficult to reverse this order).This order is undesirable because the tin is ableto oxidize during storage and on heating to theprocess temperature.Also, it is difficult to realizethick solder deposits using this strategy becauseof the need to heat above 420 °C (788 °F) todestabilize the layer of AuSn intermetallic thatwill otherwise form as a barrier layer at the in-terface between the two metals. The newmethoduses pulse plating to extract a controlled ratio ofgold to tin atoms from a cyanide-free, weaklyacidic solution and offers the prospect of highlyreproducible deposits over large areas to any de-sired thickness [Sun and Ivey 2001].

Some ingenuity has been applied to other, less-direct methods of selective application of ap-plying filler metals. One of these exploits achemical reaction to selectively apply tin-leadsolder to fine-pitch pads on printed circuit boardswithout the normal complexity of precision sten-cil printing. The so-called “Super Solder” systemuses tin powder combined with a lead salt of anorganic acid to make a paste. On heating, theorganic acid decomposes, thereby providing apartial fluxing action, and the lead so liberatedcombines with the tin to form the solder alloy insitu. The paste is applied over the entire surfaceof the circuit board. On heating above 183 °C(360 °F), any tin-lead formed where there is anexposed copper land will wet to it. The excesspaste can be simply washed off as there are noFig. 1.23 Production of foil directly from amolten charge by

strip casting. Source: Vacuumschelze GmbH, Ger-many

Fig. 1.24 Examples of foil strip produced by rapid-solidifi-cation casting technology. Source: Fleetwood et

al. [1988]

Fig. 1.25 Dendrite arm spacing decreases with increasingcooling rate and hence fine-grained microstruc-

tures have improved mechanical properties. The data pertain tohypereutectic cast iron.Adapted from Seah,Hmanth, and Sharma[1998]


flux residues to promote adhesion. By introduc-ing the lead content of the solder in essentiallyliquid form, the diameter of the tin particles canalso be decreased. By this means, it is claimedthat this process is suitable for component pitchesas fine as 80 μm (3 mils). The solder deposit istypically 100 μm (4mils) thick [Fuse, Obara, andIrie 1992].

In the electronics industry there exists a verydifferent method of applying the solder to theworkpiece. It is wave soldering, which is a pro-cess used to manufacture many millions ofprinted circuit boards (PCBs) each year. It isfundamentally different in that molten solder isapplied to the faying surfaces. There are manyvariations of the process, but, in essence, moltensolder is pumped continuously over a weir sothat there exists a stationary wave whose heightis precisely controlled. The circuit board, alreadypopulated with electronic components, is fluxed,preheated, and then passed through the very topof the solder wave. The transit time is typically1 to 2 s, making it a very rapid assembly method.Because the entire area (volume) of the joint isimmersed inmolten solder during transit throughthe wave, this removes from the process the needto achieve any spreading, so the solder need onlywet the PCB lands and component terminations.

1.3.2.3 Heating Methods

Heat must be supplied to the joint to raise thetemperature of the filler metal and joint surfacesabove themelting point of the filler. The joint sur-faces need to be heated; otherwise the filler metalwill be incapable of wetting them and thereforewill “ball up.” To prevent this situation, it is goodpracticealways toheat thefillermetalvia thecom-ponents to be joined and never vice versa.

The available methods of heating are: localheating, in which only that part of the compo-nents in the immediate vicinity of the joint isheated to the desired temperature, and diffuseheating, where the temperature of the entire as-sembly is raised.

Common local heat sources include solderingirons, gas torches, and resistance heating usingthe assembly as the resistive element. More so-phisticated heating techniques, such as inductionheating and laser heating, also fall within thiscategory. Although some methods of local heat-ing are applicable to joining in a controlled at-mosphere, this is not usually the case with asoldering iron or torch, and a flux must then beused.

In local heating, the rate of heat energy inputmust be high to overcome the heat conductedaway by the components and jigging.Ahigh rateof heat input can achieve the desirable charac-teristic of fast heating and cooling of the joint.Fast heating coupled with short heating cyclesminimizes erosion of substrate surfaces andtherefore restricts the formation of undesirablephases, while rapid cooling ensures a fine grainsize to the solidified filler and thereby superiormechanical properties. However, these potentialbenefits can be offset by the thermal distortionthat might be produced in the components beingjoined. Local heating can be used to create speci-fied temperature gradients that will restrict theflow of the molten filler metal to the immediatevicinity of the joint.

Diffuse heating sources include furnaces (bothresistance and optical), hot plates, and inductioncoils. The features of diffuse heatingmethods arethe opposite of those of the local heating meth-ods. For example, the total energy requirement ishigher, as the temperature of the entire assemblyhas to be raised, which also significantly in-creases the process cycle time.On the other hand,there is less risk of thermal distortion and accu-rate control of temperature is easier to achieve.Diffuse heating methods tend to impose fewerconstraints on the atmosphere surrounding theworkpiece, because the source of heat is rela-tively remote from the components. For ex-ample, a torch is not generally compatible witha special atmosphere.

If diffuse heating is to be used in the fabricationof complex assemblies, the designer must ensurethat all of the component parts are able to with-stand the peak process temperature. With localheating, heat sinks canbeused toprotect sensitiveareas from excessive thermal excursions. A re-lated consideration when using diffuse heating ina situation where several joints must be made isthat the melting point of the filler metal used forthe preceding joining operations must be higherthan thepeakprocess temperature used in the suc-ceeding cycle. Several different filler metals willtherefore be required to fabricate a multijointedproduct in a step-joining process.

A diffuse heating method often used in theelectronics industry is reflow soldering.Aprintedcircuit board is prepared by populatingwith com-ponents and at each joint location a controlledvolume of solder and flux is applied, often in theform of a paste. Heating is carried out in thesaturated vapor of a very precisely formulatedorganic fluid that has a condensation temperature


some tens of degrees above the liquidus tem-perature of the solder. Heaters are used to va-porize the fluid, and the circuit board is placedin the vapor stream produced. When this inter-sects with this cold body of the printed circuitboard, the vapor condenses and liberates its heatof vaporization, thereby heating the board andcomponents. The process is particularly effectivewhen the thermal mass of adjacent parts differsgreatly as the coldest components condensemorevapor and therefore receive the greatest heat in-put. Temperature gradients are therefore auto-matically minimized. Through the correct choiceof fluid, it is also possible to design the workingfluid so that flux residues are removed and thefinished assembly emerges from the processchamber dry and clean. The principal drawbacksare the cost of the working fluids and their at-tendant health and safety issues.

The oldest method of heating joints is by na-ked flame. The gases predominantly used noware acetylene and propane, burnt in oxygen.These gases are inexpensive, widely available,easy to use, and can bemade oxidizing, reducing,or neutral by adjustment of the oxygen-to-gasratio. These three combustion conditions are alsoreadily discernible by eye, allowing a skilledoperator to adjust the torch to satisfy the require-ments of the job at hand. The thermal charac-teristics of some common fuel gases burnt inoxygen are given in Table 1.6.

A relatively new development in flame tech-nology is the microflame (Fig. 1.26). This is es-sentially a conventional gas torch that uses ahypodermic needle as the gas tip. By this meansthe flame diameter can be reduced to submilli-meter dimensions. This tool is particularly usefulfor precision hand (or robotic) soldering, braz-ing, and indeed welding of three-dimensionalcomponents, such as jewelry items, musical in-struments, and “heavy” electronic parts, ex-amples ofwhich include coils and connector pins.The flame has an equivalent heat output of some-where in the region of 500 to 2000 W (370 to

1475 ft • lbf/s), so it is a considerably moreintense heat source than a soldering iron. A fur-ther convenience of these miniature gas torchesis their compatibility with a gas source derivedfrom the electrolysis of water, for example, usingan iron cell containing a concentrated sodiumhydroxide solution. Both the energy source (elec-tricity) and fuel (deionized water) are readilyavailable; the gas supply is effectively instanta-neous and does not require pressurized cylinders.This type of equipment is available commer-cially from a number of European manufacturersand possibly several others, worldwide.

1.3.2.4 Temperature Measurement

The liquid-solid metallurgical reactions thatoccur during soldering operations are highly tem-perature dependent. Therefore, reliable measure-ment of temperature is essential. Thermocouplesand pyroelectric elements are the most commontypes of temperature sensor.

Anumber of precautions should be takenwhenemploying thermocouples. Regular calibrationchecks should be made to determine if the ther-moelectric characteristics of the thermocouplematerials have altered and to test for electricalinterference affecting the display system. Cor-rect temperature measurement requires goodthermal contact between the thermocouple andthe object being monitored. This tends to presenta problem in vacuum joining processes wherethermal contact by mechanical means—namely,

Table 1.6 Thermal characteristics of commonfuel gases burnt in oxygen. In each case, theflame temperature is in the region of 3000 °C(5430 °F)Fuel gas Thermal output, kJ/cm3/s (kW/cm3)

Acetylene 15Methane 7Propane 6Hydrogen 9

Fig. 1.26 Microflame torch to solder small objects requiringa small, intense heat source. Source: Dipl. Ing.

Ernest Spirig


resting the thermocouple against a surface—tends to be inadequate. The thermal mass of thethermocouple and its protective sheath impedesthe thermocouple junction from sensing the truetemperature of the component surface. These ef-fects can be minimized by embedding the ther-mocouple within the workpiece to improve ther-mal transfer.

Even when thermocouples are used for tem-peraturemeasurement in gas atmospheres, wherethe thermal coupling is better than it is in avacuum, a change in the measured temperaturewill lag behind that actually occurring. This de-lay, which can be of the order of seconds, isdifficult to measure accurately, but it must betaken into account if a thermocouple is beingused to monitor the temperature of assembliesexposed to high heating and cooling rates.

Pyrometers have one important advantageover thermocouples: they are noncontacting sen-sors of temperature.Measurements may bemaderemotely from the workpiece, and the responsetime of the instrument can be accurately deter-mined. Traditional pyrometers are primarily de-signed for operation above about 750 °C (1380°F). However, recent advances in pyroelectrictechnology have led to the commercial develop-ment of thermal imaging bolometers that arecapable of measuring radiated energy down toand even below ambient temperature with a fastresponse time. Bolometers are now worth con-sidering not only for high-temperature brazing,but also for general application to soldering andbrazing processes. Apart from their flexibility,bolometers are capable of rapidly measuringtemperature over a large surface area. By con-trast, a thermocouple provides only a highlylocalized measurement of temperature at its tip;this might not be representative of the entirejoint region.

1.3.2.5 Joining Atmosphere

For a molten filler metal to wet and bond to ametal surface, the latter must be free from non-metallic surface films. Although it is possible toensure that this condition is met at the beginningof the heating cycle, by prescribed cleaning treat-ments, significant oxidation will generally occurif the components are heated in air. Steps musttherefore be taken to either prevent oxidation orremove the oxide film as fast as it forms.

The approach adopted will depend largely onthe atmosphere surrounding the workpiece. Sol-dering processes are conducted in one of three

types of atmosphere, defined according to thereaction that occurs between the atmosphere andthe constituent materials:

• Oxidizing (e.g., air, usually in the presence ofliquid flux)

• Essentially inert (e.g., nitrogen, vacuum)• Reducing (e.g., carbon monoxide, halogen

containing)

The implications associated with using each ofthese atmospheres are considered below.

Oxidizing Atmospheres. Air is the mostcommon oxidizing atmosphere. The principal ad-vantages of joining in air are that no specialgas-handlingmeasures are required and that thereare no difficulties associated with access to theworkpiece during the joining process. However,because most component surfaces and those ofthe filler metal are likely to form oxide scalewhen heated in air, normally fluxes must be ap-plied to the joint region.

An active flux is capable of chemically and/orphysically removing an oxide film. The flux maybe applied either as a separate agent or as anintegral constituent of the joining alloy. The sub-ject of fluxes is discussed in detail in Chapter 3,section 3.2.

Gold and some of the platinum-group metalsdo not oxidize when heated in air. Although sil-ver will oxidize at air ambient temperature, theoxide dissociates on heating to about 190 °C(375 °F). These precious metals are thereforesometimes applied as metallizations to the sur-faces of the components being joined in fluxlessprocesses. The use of wettable metallizations isdiscussed in Chapter 4, section 4.1.2.1. Soldersthat contain significant proportions of the pre-ciousmetals are generally less susceptible to oxi-dation, enabling mild fluxes to be used.

Inert Atmospheres. From a practical viewpoint, an atmosphere is either oxidizing or re-ducing. This is because it is not possible to re-move and then totally exclude oxygen from theworkpiece, except perhaps under rigorous labo-ratory conditions. Thus, when defining an atmo-sphere as inert it must be taken as meaning thatthe residual level of oxygen present is not suf-ficient to adversely affect the joining process un-der consideration. An atmosphere that might besuitable for soldering to silver is likely to beinadequate for nickel.

Because the “inertness” of an atmosphere isjudged relative to the specific application, it isnecessary to define a quantitative measure of theoxygen present. This parameter is the oxygen


partial pressure. Partial pressure provides a mea-sure of the concentration of one gas in an at-mosphere containing several gases. The partialpressure of a gas in a mixture of gases is definedas the pressure it would exert if it alone occupiedthe available volume. Thus, dry air at atmo-spheric pressure (0.1 MPa, or 14.5 psi) containsapproximately 20 vol% O2, so that the oxygenpartial pressure in air is 0.02 MPa (2.9 psi).

Typical inert atmospheres among the com-mon gases include nitrogen, argon, and hydro-gen. Hydrogen is included here because it is notcapable of reducing the oxides present on themajority of metals at normal soldering and braz-ing temperatures. The oxygen partial pressure instandard commercial-grade bottled gases is ofthe order of 10 mPa (1.5 � 10–6 psi). Higher-quality grades are available, but their cost is usu-ally too prohibitive to permit their use in mostindustrial applications.

Vacuum is frequently used as a protectiveenvironment for filler metal joining processes.Vacuumoffers several advantages comparedwitha gas atmosphere, particularly the ability to mea-sure and control the oxygen partial pressuremorereadily. In a substantially leak-free system, theoxygen partial pressure is one-fifth of the vacuumpressure, which is relatively easy to determine,as compared with direct measurement of oxygenpartial pressure. Although a roughing vacuum of100 mPa (15 � 10–6 psi) will provide an atmo-sphere with the same oxygen partial pressure asa standard inert gas, it is possible to improve onthis value, by several orders of magnitude, usinga high-vacuum pumping system.Alternatively, alow oxygen partial pressure may be achieved byobtaining a roughing vacuum, backfilling withan inert gas and then roughing out again. Theeffect of the second pumping cycle will be toreduce the oxygen partial pressure to less thantypically one-thousandth of that in the inert gas,that is, approximately 10 μPa (1.5 � 10–9 psi).This estimate assumes that the furnace chamberis completely leak tight and does not outgas frominterior surfaces, nor does any oxygen or watervapor backstream through the pump.

The disadvantages of using a vacuum systemfor carrying out a joining process are, principally,restricted access to the workpiece and the inad-visability of using either fluxes or filler metalswithvolatileconstituents,suchascadmium,asthevapors can corrode the vacuumchamber, degradeits seals, and contaminate the pumping oils.

A frequently overlooked consideration in re-duced-pressure atmospheres is adsorbed water

that naturally exists on surfaces that are exposedto ambient atmospheres. The continuous stream-ing of water vapor that desorbs from surfaces andflows past the workpiece as the pressure in avacuum chamber is reduced is a source of oxi-dation; this is discussed in quantitative detail inChapter 3, section 3.1. In a vacuum system op-erating at 10 mPa (1.5 � 10–6 psi), the desorbingwater vapor constitutes the major proportion ofthe residual atmosphere.An adsorbedmonolayerof water vapor of just 100 mm2 (0.16 in.2) in areadesorbs to a gas pressure of 4 mPa (6 � 10–7 psi)per liter of chamber volume. The surfaces of thechamber should therefore be smooth tominimizethe surface area and also dry. In order to reducethis problem further, the walls of the vacuumchamber should be heated and the system shouldbe vented to a dry atmosphere. To effectivelydesorb water vapor, the bakeout temperatureshould be at least 250 °C (480 °F), which maybe difficult to achieve in practice owing to designconstraints and the employment of rubber andother organic seals.

Another source of oxidizing contamination ina vacuum system is oil vapor mixed with air andwater vapor, backstreaming from a rotary pump.This can occur whenever the pressure inside thevacuum chamber drops below 1 Pa (1.5 � 10–4

psi), but can be largely eliminated by employinga foreline trap or by isolating the pump from thechamber once the required pressure reductionhas been obtained.

Thewidespread practice of relying on an opengas shroud to provide an inert atmosphere isoften unsatisfactory because it is extremely dif-ficult to control such an atmosphere reliably. Forexample, turbulence in the inert gas shroud canresult in a supply of air actually being directedat the workpiece. Recent advances in furnacetechnology now permit open furnaces, often beltfurnaces intended for reflow soldering of PCBs,to achieve very high specification atmospheresin the working zone through careful design of thegas flow at the open portals.

A reducing atmosphere is one that is ca-pable of chemically removing surface contami-nation from metals. Gases that provide reducingconditions are generally proprietarymixtures thatliberate halogen radicals. Specific gas-handlingsystems are usually needed for these in order tosatisfy health and safety legislation.

For a few metals, hydrogen is satisfactory asa reducing atmosphere. No less important formeeting its functional requirement than the oxy-gen partial pressure of the gas is its water con-


tent. Hydrogen is a relatively difficult gas to dry,and the water vapor present can present a seriousproblem. A frost point of –70 °C (–95 °F) isequivalent to a water content of 0.0002% byvolume—that is an oxygen partial pressure ofabout 10 mPa (1.5 � 10–6 psi). There is also therisk of explosion when dealing with hydrogen athigh temperatures, and hydrogen can embrittlesome materials. A more detailed treatment ofreducing atmospheres and their use is given inChapter 3, sections 3.1 and 3.3.

1.3.2.6 Coatings Applied toSurfaces of Components

Only occasionally is the desired joining al-loy (chosen on the basis of melting tempera-ture and physical properties) metallurgicallycompatible with the substrate in the sense thatthe filler will wet the substrate uniformly, with-out the consequential formation of embrittlingphases by reaction. A solution is to apply a sur-face coating that will promote wetting by thefiller and react with it in a benign manner.Coatings can be applied by a variety of tech-niques and to thicknesses that suit the particu-lar application.

Onmetals, coatings are usually applied by wetplating methods, which are quick, economical,and flexible with regard to the coating thickness.Electroplating cannot often be used directly formost nonmetals, and it is more common insteadto rely on vapor-deposited coatings. If the sub-strate is refractory in character, adhesion ofmetalcoatings tends to be poor unless themetal is itselfsufficiently refractory so that it will form a strongreactive bond to the substrate. Widely used met-allizations are chromium or titanium as the re-active layer, overlaid by gold or a platinum groupmetal to provide protection from the atmosphere.These and other metallizations and the principleson which they are designed are described inChapter 4, section 4.1.2.1.

1.3.2.7 Cleaning Treatments

The surfaces of the components to be joinedand the filler metal preforms must be free fromany nonmetallic films, such as organic residuesandmetal oxides, to enable themolten fillermetaltowet and alloywith the underlyingmetal. Fluxesare often capable of removing surface oxides,provided they are reasonably thin.

Organic films can be removed with solvents,which obviously should not react with the un-

derlying materials. Thick oxides and other non-metallic surface layers can be removed chemi-cally. However, mechanical cleaning is generallypreferable because chemicals tend to leave resi-dues, which then also have to be removed. Aprocedure that has proved effective with solderwire is to wipe the latter with a fibrous tissuesoaked in solvent. The first few wipes leave ablack mark on the tissue, which is a combinationof native oxides and manufacturing residues (in-cluding lubricating oils). As shown by the datapresented in Fig. 1.27, the solder wire should bewiped multiple times and, ideally, used within1 h (see Chapter 3, section 3.3).

Mechanical abrasionexposes a freshmetal sur-face.The roughness of the abraded surface can bereadily controlled, and this can be used to advan-tage in promoting the spreading of the moltenfillermetal.The rougher the surface, the better thewetting and spreading of the molten filler tend tobe, for the reasons given in Chapter 1, section1.3.5. However, rough surfaces create problemswhen it is required to cover them with thin metalcoatings. For example, thickness uniformity ofthin metallic films is difficult if not impossible toachieve on a rough-textured substrate.

Soft components, such as solder foils, can bedifficult tomechanically cleanusingabrasivepar-ticles because these tend to get embedded. Theseand thin vapor-deposited and electroplated met-allizations can be protected against atmosphericcorrosionby theapplicationofanoblemetalover-coat. If correctly stored and handled, such com-ponentswill not require cleaning prior to bondingand, for obvious reasons, must never be abraded.

1.3.2.8 Heat Treatments Prior to Joining

Prejoining heat treatments are occasionallyuseful in providing stress relief and thereby pre-

Fig. 1.27 Improvement in shear strength of a fluxless jointmade with In-48Sn wire as a function of the num-

ber of wipes made with a paper tissue soaked in ethanol, im-mediately prior to melting of the solder


venting unpredictable distortion during heatingof the components to the bonding temperature.Other situations where prejoining heat treat-ments can be beneficial include those involvingcomponents with nonmetallic surface films thatare thermally unstable. In the case of silver, forexample, the oxide will readily dissociate whenheated above 190 °C (375 °F) in an ambientatmosphere. Likewise, silver sulfide dissociateson heating above 842 °C (1548 °F).

Heating cycles may be used to produce solderalloys from layers of the constituent elementsapplied to surfaces by screen printing, electro-plating, or vapor deposition. By heating the sub-strate above the melting point of the constituentwith the lowestmelting point, alloyingwill occurby solid-liquid interaction. The joint can then beformed in a subsequent heating cycle that is usu-ally referred to as the reflow stage.

1.3.2.9 Heating Cycle of theJoining Operation

The prepared components and fillermetal, pos-sibly mounted in jigs, are joined by applyingheat. The heating cycle involves four importantprocessing parameters: the heating rate, the peakbonding temperature, the holding time above themelting point of the filler, and the cooling rate.

In general, it is desirable to use a fast heatingrate to limit reactions that can occur below theprescribed bonding temperature. However, themaximum heating rate is normally constrainedby adverse temperature gradients developing in

the assembly. These can produce distortions inthe components and give rise to nonuniform re-actions between the filler and the two joint sur-faces.Also, temperatures are difficult to measurereliably during fast heating schedules. A betterpractice, when joining in a vacuum or specialatmosphere furnace, is to heat the assembly rap-idly to a preset temperature that is just below themelting point of the filler metal and then hold atthis temperature for sufficient time (which canrange from a few seconds to over one hour, de-pending on the size of the assembly and the heat-ing method) to allow the assembly to thermallyequilibrate and for water vapor to flush out of thejoint. Following this dwell, the assembly maythen be rapidly heated to the bonding tempera-ture. This method is used to make joints to PCBsusing lead-free solders where only a few degreesof superheat are permitted. Profiles of typicaltemperature cycles are shown in Fig. 1.28.

The bonding temperature should be such thatthe filler is guaranteed to melt, but at the sametime should not be so high that the filler degradesthrough the loss of constituents or by reactionwith the furnace atmosphere. The optimal tem-perature is normally determined bymetallurgicalcriteria, most importantly the nature and extentof the filler-substrate interaction. The peak pro-cess temperature is frequently set at about 50 and100 °C (90 and 180 °F) above the melting point,because accurate temperature measurement andcontrol is not always readily achievable, espe-cially in reduced-pressure atmospheres, wherebulky jigging is used or where conduction be-

Fig. 1.28 Profiles of typical temperature cycles. (a) Heating cycle with a controlled profile incorporating dwell stages to reducethermal gradients. (b) Heating cycle defined solely by attainment of a peak temperature


tween the heat source and the workpiece is poor.Moreover, the reported melting temperatures ofsome fillers are not based on accurate measure-ments, and it is prudent to make some allowancefor this uncertainty. The minimum time that theassembly is held above the melting point must besufficient to ensure that the filler has melted overthe entire area of the joint and the maximum timeis usually a compromise based on practical andmetallurgical considerations. Extended dwelltimes tend to result in excessive spreading by themolten filler, possible oxidation gradually takingplace, and deterioration of the properties of theparent materials.

The cooling stage of the cycle is seldom con-trolled by the operator, but tends to be governedby the thermal mass of the assembly and jig.Forced cooling can lead to problems, such asexacerbating mismatch stresses. Occasionallyone or more dwell stages are required, either toprovide stress relief to the bonded assembly, orto induce some requisite microstructural change.An example of the latter would be the agingtreatment of a precipitation-hardening alloy. Inthis case, the solution treatment and quenchingstages are carried out in tandem with the actualreflow operation.

A heat treatment temperature of about 75% ofthe freezing point of the filler metal in Kelvinusually provides the optimal relief of residualstresses. An example of an assembly that re-quired a stress-relief treatment in order to avoidcatastrophic cracking is shown in Fig. 1.29.

1.3.2.10 Postjoining Treatments

Various types of postjoining treatments can beapplied. A cleaning schedule is generally used to

remove flux residues and the tarnishing that areproduced when joining in air. Flux residues mustbe removed, as they are usually corrosive, es-pecially when moist and can affect the long-termreliability of the component in service. Bothchemical and mechanical means of flux removalare employed. Tarnishing tends to be removedchemically, often with acids, followed by thor-ough rinsing. If a heat treatment is not integratedinto the cooling stage of the bonding cycle, aseparate heat treatment may be carried out sub-sequently.

1.3.2.11 Postjoining Cleaning

Formany years it has been standard practice toclean assemblies after soldering. This need arosefrom the corrosive nature of traditional rosin-based fluxes and the subsequent reliability andaesthetic degradation that can occur if the fluxresidues are not removed. Particularly in the elec-tronics industry, the preferred cleaning agent hasbeen a blend based on chlorofluorocarbons(CFCs). These liquids were used in vapor clean-ing equipment of low complexity and thereforelow cost. In addition to being relatively efficientsolvents, CFCs had good compatibilitywithmostengineeringmaterials including solders, no flam-mability risk, and low toxicity. The process hadtheadditionalbenefit that theproductemergeddryfrom the cleaning operation. The ease of cleaningin this manner was largely responsible for it to bethe norm to clean all parts and assemblies—itwasless expensive and easier to clean than to risk sub-sequent problems in service.

The discovery in the mid-1970s that CFCswere contributing to stratospheric ozone deple-tion led to elimination ofCFCs from solder clean-ing and ultimately to a ban on their use andmanufacture. The response by industry and par-ticularly those engaged in electronics manufac-ture was threefold.

First, alternative cleaning chemicals andmeth-ods were devised. Many of these are now com-mercially available systems. As a result of thissubstitution endeavor it is perhaps ironic that ithas been proven that cleaning with CFC-basedsolvents is actually fairly ineffective, especiallywith modern surface-mount and chip-on-boardtechnologies.

Second, new and improved flux chemistrieswere formulated, including “water soluble” and“no-clean” fluxes (see Chapter 3, section 3.2).

Finally,manymanufacturers undertook a thor-ough reappraisal of the need for cleaning, par-

Fig. 1.29 Large-area silicon chip soldered into a metallizedceramic package using an alloy based on the Au-

3Si composition. The expansion mismatch between silicon andthe ceramic makes it necessary to cool the assembly at a con-trolled rate in order to prevent fracture of the components.


ticularly given the short life of some productsowing to technical obsolescence or the tran-sience of fashion; examples include wrist-watches, portable computer games consoles, andmobile telephones.

Each of these options is considered further inthe paragraphs that follow.

Cleaning of ElectronicAssemblies.The cur-rent scene of cleaning methods is very analogousto the situation regarding lead-free solders as areplacement for lead-tin eutectic; namely, an al-most universal cleaning route has now been re-placed with a plethora of alternatives. This ispartly due to commercial pressures as the variousmanufacturers compete for market share. Theavailable cleaning methods can be broadlygrouped into four categories:

• Solvents• Semiaqueous formulations• Aqueous formulations• All of the above with ultrasonic or high-

pressure jet assistance

Comparative studies have been made of the rela-tive effectiveness of each of these approaches[Richards et al. 1993], from which the followingconclusions can be made:

• There are many CFC-free cleaning optionsthat are technically and commercially viable.The choice of which to use requires carefulconsideration of the type of assembly forwhich cleaning is required, the level of clean-liness it is necessary to obtain, and the natureof the residues it is permitted to leave.

• Cleaning process times are significantlylonger (10–20 min) compared with that forCFCs (3 min). Although the intrinsic clean-ing step is not significantly longer, the overallprocess time is substantially increased owingto the need to employ multitank processingand a drying stage.

• Mechanical assistance, most practicably inthe form of ultrasonic agitation or high-pressure jet sprays, always results in a cleanerproduct for a given cleaning time. This op-tion is not always possible as some parts canbe degraded or damaged by aggressive wash-ing [Richards, Burton, and Footner 1993].

• The ability to clean flux from narrow gaps,such as from underneath components, isroughly related to h3, where h is the width ofthe gap. Consequently, components such asleadless ceramic chip carriers (LCCCs) with

stand-offheightsof50μm(0.002in.)aremuchmore difficult to clean under than plasticleaded chip carriers (PLCCs), which havestandoff heights around 100 μm (0.005 in.).

• Different substrates and components have dif-ferent levels of inherent cleanliness. This isdue to the relative chemical affinity and de-gree of mechanical adhesion between theflux/flux residues and the substrate. Thus, forexample, ceramic circuit boards have onlyhalf the contamination of fiber-reinforcedlaminate (FR4) boards as reflowed, specifiedin terms of μg NaCl equivalence.After clean-ing, the results are not greatly different.

• Best cleaning results are obtained by opti-mizing the thermal profile of the solderingcycle for the cleaning system used. Somecleaners are far more effective at removingflux residues than unspent flux, and vice versafor others. Usually,manufacturers sell a pack-age comprising a flux and recommendedcleaning system and will be able to advise onthe most appropriate thermal cycle as well.

Comparative studies reveal that the solvent, flux,and equipment manufacturers have a clear un-derstanding of the capability of their products toclean (see Fig. 1.30). There is negligible tech-nical merit in deviating from their recipes.

Water-Soluble and No-Clean Fluxes. Wa-ter-soluble fluxes are designed such that residuesand unspent flux are miscible with water. Thus,the assembly can be cleaned in water. To avoidwater stains on the product, it is usual to com-plete the cleaning process with an organic rinse,usually some form of alcohol.

Another family of fluxes has been developedthat carries thedesignation“no-clean.”Thechem-istry of these materials has been carefully for-mulated to ensure that unspent flux and flux resi-dues are chemically bound or otherwise bufferedand remain relatively inert in the presence ofmoisture. Modern versions of these fluxes havebeen further tailored so that the residue can pro-vide additional functionality.Aprime example isprovided by no-clean fluxes developed for flip-chip underfill applications, where the flux resi-due is designed to have a particular modulus andexpansion coefficient that helps to extend thefatigue life of the solder joints.

As a general observation, the soldering pro-cess window for water-soluble and no-cleanfluxes tends to be somewhat narrower than for aprocess utilizing a more conventional flux for-mulation.


Alternatives to Cleaning. There are threealternatives to cleaning—simply do not clean atall, use fluxless processes, or switch to conduc-tive adhesives [Lea 1998a; Lea 1998b].

Electrically conductive adhesives have madeconsiderable technological progress and arewidely used to interconnect to LCD displays.However, they have not yet reached the point ofmaturity where they can be used reliably for in-terconnects that carry more than a few micro-amperes of current or high frequencies. There arealso issues concerning their stability in service:deterioration of their adhesion and the slow evo-lution of volatile species, which is of particularconcern when conductive adhesives are used tointerconnect bare die or are in the vicinity ofnaked optical facets. No doubt, improved sta-bility will be achieved as better-formulated prod-ucts become available.

At face value, fluxless processes are highly at-tractive because there is no flux and therefore nocleaning required. However, fluxless processesare extremely difficult to implement and are in-compatible with low-cost volume manufacture.To obtain satisfactory wetting and spreading, theprocess atmosphere needs to be denuded of oxy-gen andwater vapor to levels that can only be ob-tained in closed vessels. More details on fluxlesssoldering are given in Chapter 3, section 3.3.Some care needs to taken to ensure that a “flux-less” process does leave a clean product after sol-dering. The authors encountered one process in

which the solder perform was etched in 10% hy-drochloric acid and then rinsed indeionizedwaterprior to use. The process was claimed to be flux-less because nofluxwas used, and all the acidwassupposed to be rinsed off the solder. However, asthe solubility of hydrochloric acid in water de-creases faster than does the dilution ratio, in prac-tice the dried solder preform and final productwere both proved to be severely contaminatedwith chlorides. It transpired that better-qualityjoints with lower levels of ionic contamination ofthe product were obtained by using a commercialno-clean flux without cleaning!

The ultimate alternative to cleaning is simplynot to do it. Cleaning adds to the capital andconsumable process cost and results in an as-sembly yield loss. Provided the flux residues donot adversely impact the reliability of the prod-uct, then the merits of cleaning are truly ques-tionable. This is particularly true for short-lifeconsumer products, for example, musical greet-ing cards. Because the principal risk to leavingflux residues on the product is corrosion arisingafter the residues react with moisture, anotherfrequent approach is to coat the entire product inlacquer. This provides a sufficient barrier tomois-ture ingression over the recommended life of theproduct. Obviously for life- or mission-criticalproducts, cleaning will probably always be un-dertaken, but in many instances acting in rec-ognition of a finite life is a technically and eco-nomically sound approach.

Fig. 1.30 The relative effectiveness of different cleaning agents on glass-reinforced epoxy laminate FR4 printed circuit boards (PCBs)as measured by residual sodium chloride contamination. The “reflowed” bar references as-made and uncleaned PCBs.

Adapted from Richards et al. [1993]


1.3.2.12 Statistical Process Control

All processes are subject to variation, andachieving stability of processes is an importantstep in any quality-improvement program. Theapplication of statistics to monitor and controlthe variability is termed statistical process con-trol (SPC) [Ledolter and Burrill 1999]. Manyindustrial soldering processes are subject to SPC.A modern PCB assembly line can achieve jointdefect rates of a few ppm. This means, quiteliterally, that only one or two soldered joints inevery million made would fail a quality inspec-tion. This degree ofmanufacturing quality is onlyachieved through SPC.

A fundamental tool in SPC is a graphical dis-play,knownasacontrol chart.This chart providesthe basis for decidingwhether the variation in theoutput of a process is due to common, randomlyoccurring variations or to unusual causes, whichwould require investigation and action. The con-trolchart isachronologicalplotofparticularchar-acteristics, such as joint strength or peak reflowtemperature, sampled at periodic intervals. Thisinformationfurnishesdataon theprocessstabilityand provides an understanding of improvements,where made. Whenever a significant deviationfrom the norm is detected, a decision can bemadeto adjust a process variable in order to bring theoutput back to the required quality level. Obvi-ously, to accomplish this there must be a properunderstanding of the relationship between theprocess variables and the output.

There are different types of control charts, de-signed for different situations, which are classi-fied by the type of data they contain. Control

charts designed to monitor the proportion of de-fective items are referred to as p-charts, whilecharts that track the number of defects on theproduct are known as c-charts. Both are used todescribe attribute data, that is, a record of thepresence and absence of certain characteristics.Quantitative data are monitored using a mean orx-bar chart, while process variability ismeasuredusing range charts (r-charts) and standard devia-tion charts (s-charts).

The basis of SPC is, for each parametertracked, to select upper and lower control limits(see Fig. 1.31). These are set at a multiple num-ber of standard deviations from the mean suchthat there will be a high probability that the datawill fall between these limits when the processis working as desired. Only when the processmetrics drift further from the mean is interven-tion required.

It is possible to apply SPC to virtually anyprocess or machine output. In many instances, itis a highly effective basis for controlling manu-facturing processes. As with any tool it is nec-essary to use some discretion and critical thoughtto ensure that SPC is appropriate and the cost ofimplementing and sustaining it is justified.

1.3.3 Health, Safety, andEnvironmental Aspects ofSoldering

Soldering encompasses the use of a large num-ber of different materials, covering metallic andnonmetallic elements for the fillers and the par-ent materials, and organic and inorganic chemi-

Fig. 1.31 Statistical process control chart for peak reflow temperature (indicated by solid squares) measured at a test point on aprinted circuit board, showing the upper and lower control (i.e., intervention) points for this process


cals used in fluxes, controlled atmospheres, andfor removing flux residues. Several of these ma-terials are hazardous in varying degrees to theoperators or to the environment [Sax and Lewis1989]. Accordingly, they must be handled, used,and disposed of according to national codes ofpractices or regulations governing hazardoussubstances. Official listings produced by na-tional health and safety authorities classify ma-terials according to their toxicity level, for ex-ample, the exposure limits for hazardousvapors and dusts.

The main problem with solders and solderingfluxes ariseswhen they are heated tomake a joint.The fume contains a cocktail of gases that cancause eye and nose irritation, dermatitis, asthma,and respiratory problems. The fume contains fineparticles, in the range 0.1 to 1 μm (4 to 40 μin.),which is the most dangerous size distribution forcausing long-term lung damage. The recom-mendedsolution is to ensure that theworkplaceorworkchamber isventilatedusinganappropriatelydesigned extraction system that is able to exhaustthe gases and trap the particulates [Jakeway1994]. Preventing exposure to the hazard by ap-propriatemeasuresshouldalwaysbegivenhigherpriority than protective measures.

All materials that are likely to be encounteredin a joining context will have an assigned valueof maximum exposure limit, usually in weightper unit volume (normally in mg/m3). The formof the material is also relevant. Powders anddusts are more hazardous than nonvolatile liq-uids and monolithic solids: they are ranked ac-cording to the maximum inhalable quantity inmg/m3, time weighted over a period of time,either short term, meaning minutes, or over alonger period of many hours. For correct inter-pretation of the rules, regulations, and audits,reference should always be made to a qualifiedsafety practitioner, as there are often also legalaspects to consider.

Care must also be taken in the storage of ma-terials both prior to use and in the procedures forthe subsequent disposal of residues, exhaustemissions, and other associated effluent, such assolutions containing rinsed fluxes. These are usu-ally subject to statutory controls. For many or-ganic chemicals and gases, in this context solderfluxes, binders used in pastes, and halogenatedgases, there may also be fire risks to consider.The flammability is rated according to flash-point, which is the lowest temperature at whichthe substance can be spontaneously ignited whenit is in a saturated condition.

Appendix A1.1:Solid-State Joining withGold, Indium, andSolder Constituents

Gold has three desirable characteristics thatrender it a most suitable metal for making dif-fusion-bonded joints. These are its low modulus,rapid self-diffusion, and absence of an oxide skinwhen heated in air.As a result, diffusion bondingof gold can be achieved at room temperaturewith plastic deformation of as little as 20%. Thetemperature/pressure curve for a process time of1 h is given in Fig. 1.32 [Humpston and Baker1999]. Extrapolation of the graph predicts that asuccessful gold-gold bond can be achieved with-out pressure above approximately 450 °C (840°F). Certainly it is the authors’ experience thatgold rods merely placed in contact bond readilyat 500 °C (932 °F).

Indium can also be used to make pressure-welded joints. Because indium tends to be cov-ered by a relatively thick layer of oxide, the pro-cess usually relies on extensive physicaldeformation to ensure adequate virginmetal con-tact. Heating is not normally necessary becauseof the extensive deformation and, at room tem-perature, the metal is already very close to itsmelting point. A typical temperature/pressureprocess curve for indium is given in Fig. 1.33[Plotner et al. 1991].

Diffusion bonding and pressure welding can,perhaps surprisingly, be achieved using standardsolders. The authors have witnessed one high-volumemanufacturing linewhere a hermetic sealwas made between two gold metallized surfacesusing a wire of ordinary (fluxless) solder. Two

Fig. 1.32 Temperature/pressure curve for diffusion bondingof gold, for a process time of 1 h. The line on the

graph differentiates between joints of acceptable (above) andunsatisfactory (below) tensile joint strength after fabrication.


solders used in this manner include Sn-40Pb andIn-48Sn. In both instances, the trick to obtaininga satisfactory joint lies in ensuring that the fayingsurfaces of both the components and the solderare as free from nonmetallic contamination aspossible. The joint wasmade by forming a lengthof freshly cleaned solder wire into a ring, buttwelding the ends, and simply pressing it betweenthe two component surfaces, at room tempera-ture. Sufficient load was applied to cause flow ofthe solder out to the joint edge and a substantialnarrowing of the joint gap; so the process wasclearly pressure welding. The resulting joint pro-vided a key seal on a hermetic cavity, and theproduct carried a 20-year guarantee against her-meticity failure by the manufacturer. The meritof using solder preforms for this application isthat they are readily available at low cost com-pared with specially deposited coatings of goldor indium.

Appendix A1.2:Relationship amongSpread Ratio, SpreadFactor, and ContactAngle of Droplets

Expressions describing the spread of a moltenmetal droplet are derived under the following setof idealized conditions:

• The original metal pellet is in the form of aspherical bead of radius a (and diameterD � 2a).

• The droplet resolidifies after spreading on thesubstrate as a spherical cap of radius R andheight h, its interface with the substrate hav-ing a diameter 2A, as shown in Fig. 1.34.

• The volume of the original pellet is equal tothe volume of the resolidified droplet. Thismeans that any volatilization of the moltendroplet and reaction with the substrate do notmeasurably affect its volume.

The volume of the spherical cap is:

V �1

6�h (h2 � 3A2)

Spread Ratio and Contact Angle

The spread ratio, Sr is defined as:

Sr �Plan area of spread on the substrate surface

Plan area of the original spherical pellet

A and a are related by the conservation of thevolume of the droplet, that is:

V �4

3�a3 �

1

6�h (h2 � 3A2)

Therefore:

a �1

2�h(h2 � 3A2)�1/3

Fig. 1.33 Temperature/pressure curve for diffusion bondingof indium, for a process time of 1 h. Good-quality

joints are obtained from the conditions above the boundary lineand lesser-quality or no joint from those below.

Fig. 1.34 Spherical cap geometry


and

Sr �4A2

�h(h2 � 3A2)�2/3

From the geometry (Fig. 1.34), A � R sin �and h � R(1 – cos �):

Sr �4A2/h2

(1 � 3A2/h2) 2/3

�4 cot2 �/2

(1 � 3 cot2 �/2) 2/3

for 0° < � < 180°

Spread Factor and Contact Angle

The spread factor is defined by:

Sf �D�h

D�

(h3 � 3A2h)1/3 �h

(h3 � 3A2h)1/3

� 1 �1

(1 � 3A2/ h2) 1/3

� 1 �1

(1 � 3 cot2 �/2) 1/3

for 0° < � < 180°

Contact Angle and theDimensions of the SolidifiedPool of Filler

From Fig. 1.34, it can be seen that:

A2 � (R � h)2 � R2

according to the Pythagorean theorem. Rear-ranging this equation:

R �A2 � h2

2h

Therefore:

sin � �2

(A/h) � (h/A)

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CHAPTER 2

Solders and Their Metallurgy

This chapter presents an overview of solderalloy systems that one is likely to encounter.Necessarily, the survey must include consider-ation of the parent material with which the solderis used because the suitability of a solder for aparticular joining process will depend largely onits compatibility with the base materials.

Extensive reference is made to phase dia-grams in order to highlight particular points. Anintroduction to alloy constitution and phase dia-grams designed for those with no background inthe subject is presented in section 2.3, whichcovers interpretation of phase diagrams and theassociated terminology. Methods for their deter-mination are summarized in the literature [Hump-ston and Jacobson 1993, Chapter 3].

For a solder to be compatible with a particularparent material, it must exhibit the followingcharacteristics:

• A liquidus temperature below the meltingpoint (solidus temperature) of the parent ma-terials and any surface metallizations. Usu-ally a margin is required between these twotemperatures in order to achieve adequate flu-idity of the molten filler. Strictly speaking,the fluidity of solders is not a strong functionof temperature, but the overall flow behavior,which is of immediate practical interest, doesoften depend substantially on the applicationtemperature.

• Capability of producing joints at tempera-tures at which the properties of the base ma-terials are not degraded. For example, manywork-hardened and precipitation-hardenedalloys cannot withstand elevated tempera-tureswithout loss of their beneficialmechani-cal properties. Work hardening involves sub-jecting the alloy to mechanical deformationsuch as rolling or hammering, when reason-ably cold. As the temperature is raised, the

deformation damage is removed by atomicrearrangement in the metal. Precipitationhardening is accomplished by creating afinely divided phase within the material,which can be thought of as akin to a com-posite material. The dispersed phase is pre-cipitated by means of an appropriate heatingschedule and its strengthening effect is like-wise degraded by high temperatures or pro-longed exposure at low temperatures, both ofwhich tend to coarsen this phase.

• The ability of the parent materials, or a met-allization applied to the parent materials, tobe wetted in order to ensure good adhesionthrough the formation of metallic bonds. Asexplained in the preceding chapter, it is notpossible to get spreadingwithout wetting, butan absence of spreading does not automati-cally imply a lack of wetting. For example,all solders will wet platinum, but only gold-base solders will spread on this metal.

• Limited erosion of the parent metals at thejoint interface.The associated alloying,whichmust occur to form a metallic bond, shouldnot result in the formation of either a largeproportion of brittle phases within the joint orof significant concentrations of brittle phasesalong interfaces or other critical regions ofthe joint. Even ductile phases can have weakinterfaces with solidified filler alloys. Simi-larly, the products of interalloying must notgenerate other forms of weakness such ascorrosion or voids in the joint.

• Elimination of constituents or impurities thatmight embrittle or otherwise weaken the re-sulting joint. Likewise, the parent materialmust not contribute constituents or impuritiesto the solder that will have a similar effect.

Besides being compatible with the parent ma-terial, the solder and the joining process used



must be mutually suited. For example, solderscontaining zinc, lead, or other volatile constitu-ents are not usually appropriate for furnace join-ing at elevated temperature, especially whenthese involve reduced pressures.

The degree of temperature uniformity that canbe achieved over the joint area will have an in-fluence in determining the minimum tempera-ture difference that can be tolerated between themelting temperature of the filler and the meltingor degradation temperature of the parent mate-rial. This consideration is particularly relevant tothe joining of aluminum and stainless steel com-ponents, for contrary reasons. Stainless steel pos-sesses relatively poor thermal conductivity,whereas aluminum has a low thermal heat ca-pacity, both of which make it difficult to obtainuniform heating throughout a bulk assembly us-ing local heat sources that do not encompass theworkpiece.

The condition of the surface of the parent ma-terial may affect its compatibility with the solder,especially when fluxes are not used. As an ob-vious example, filler metals will less readily wetan oxidized surface than a freshly cleaned metalsurface. This consideration often determines theacceptable shelf life of components prior to join-ing. The term “atomically clean” should not gen-erally be used to describe cleaned surfaces. It isan abstract ideal, not normally relevant in prac-tical situations!

In order to establish whether a particular par-ent metal (or nonmetal with a surface metalliccoating) is compatible with a given solder, it isnecessary for the appraisal to be carried out un-der conditions that are likely to be representativeof those used in any practical implementation ofthe joining process. Parameters such as processtime and temperature can be critical in this re-gard. Storage shelf life of the components (andthe filler when in the form of paste) is anotherrelevant factor that needs to be taken into ac-count but is often neglected during transfer of aprocess from the laboratory to the factory. Thisis particularly true where the parent metal has amultilayer metallic coating, a situation that iscommonly met with in soldering. Although theoutermost layer, often gold, will have indefiniteshelf life, if it is not fully dense or sufficientlythick, oxygen will be able to percolate throughit and oxidize less noble metals beneath the goldlayer. The problem then only becomes apparentwhen the gold dissolves in the solder, exposingthe underlying metal that has been rendered non-wettable during storage.

The properties of the solder and the resultingjoint must also be compatible with the servicerequirements. These are likely to involve a com-bination of at least some of the following con-siderations:

• The strength and ductility of the joint shouldmeet certain minimum requirements over therange of service temperatures. Soldered jointsare primarily used in the assembly of elec-tronic and optical products where dynamiccharacteristics, such as resistance to thermalfatigue and creep, are usually of greater im-portance than quasi-static mechanicalstrength.

• The design of the joint should not introducestress concentrations in the assembly thatmight arise through solidification shrinkageor formation of intermetallic phases. Like-wise, the design should not cause undue dis-tortion of the assembly from bimetallic ef-fects, excessive thermal expansion mismatchstrain, or the filler metal being too unyield-ing.

• The joint is normally required to be resilientto the service environment, in terms of cor-rosion and oxidation resistance and compat-ibility with vacuum, in accordancewith func-tional requirements. Zinc-base solders areparticularly limited in this regard.

• The filler must be compliant with statutoryneeds. These include hallmarking regula-tions for precious metals and health restric-tions on lead and cadmium for certain culi-nary and medical applications and nickel inapplications that involve prolonged skin con-tact, such as spectacle frames.

• Aesthetic requirements are usually impor-tant, for example, color, color matching injewelery and utensils, the ability of joints toaccept surface finishes such as paints, elec-troplatings, and so on. Good fillet formationis often demanded for aesthetic reasons andalso as a criterion of acceptable joint quality.The reader is cautioned that the latter infer-ence can be misleading, particularly for largearea planar joints (see Appendix A4.2).

• Joints are normally required to possess cer-tain thermal and electrical properties. Theseoften represent essential constraints in elec-trical and optical assembly, which directlydetermine fitness for purpose.

The simultaneous attainment of several of thedesired characteristics is frequently achievablewith common filler metals, provided the basic


design guidelines and process conditions are sat-isfied. Lead-tin eutectic solder, which is taken toinclude the Pb-60Sn alloy, accounts for the larg-est proportion of joints involving filler metalsthat are encountered. It is the more exceptionaland demanding service requirements that havegiven rise to the development of the hundreds ofadditional filler metal compositions. However, ingeneral, soldering is not a difficult process tomaster. The commercially available fillers andmatching fluxes have been designed so that, whenused in conjunction with the common engineer-ing parent materials, they meet many if not all ofthe requirements listed previously.

2.1 Survey ofSolder Alloy Systems

The survey begins with consideration of lead-tin solder on account of its preeminent positionamong the lower-melting-point filler metals. Thereview then extends to consider the less commonsolders, arranged in rough order of usage anddecreasing melting point. To readers familiarwith brazing technology, one important differ-ence of solders will soon be apparent. That is,solders are largely based on unique eutectic com-positions. The addition of other alloying ele-ments usually destroys the eutectiferous charac-ter of these alloys and the desirable solderingproperties that stem from it. Thus, the alloy com-positions are normally exact rather than encom-passing a broad range, an important exceptionbeing the indium-lead series of alloys, which arenot eutectiferous. In any case, there is little in-centive to introduce such modifications becausethere is a ready choice of alternative eutecticsolders having similar melting points, as illus-trated by Fig. 1.3. For detailed coverage of theavailable solder alloys, the reader should consultreference publications [e.g., KleinWassink 1989,Manko 2002, Schwartz 2003] as well as suppli-ers’ data sheets and manuals.

All solder alloys are based on combinations oftwo or more constituents, chosen from nine com-mon elements, namely antimony, bismuth, cad-mium, gold, indium, lead, silver, tin, and zinc. Ofthese, the cadmium-containing alloys havelargely been removed from manufacturers’ cata-logs because they offer no clear advantages overother solders, and their use is subject to restric-tions arising from the toxicity of cadmium fumes.As with beryllium, it transpires that cadmiummetal is relatively safe, and it is the oxide that is

hazardous to health. In recent years, the use oflead-containing solders for joining domestic wa-ter pipes has been largely banned. This has ledto the development of lead-free solders forplumbing applications [Irving 1992]. These sol-ders comprise typically 95% tin with the balancebeing one or more of silver, antimony, bismuthcopper, nickel, and zinc. Solders that do not con-tain lead are also actively being used in elec-tronics manufacture as part of a portfolio of mea-sures by companies toward environmentalresponsibility. Further information on this ini-tiative is given in Chapter 5, section 5.1.

High-melting-point solders contain either goldor lead as the principal constituents. The major-ity of lower-melting-point solders, that is, thosewith melting points below 150 °C (300 °F), fallinto two groups depending on whether the activeconstituent of the alloy is tin or indium. Thisclassification is determined on the basis of thelikely reaction of the solder with the substratematerial, which, in the temperature range cov-ered by solders, is necessarily a metal becausethere is insufficient thermal activation availablefor wetting nonmetals.

In most cases, wetting of a component by asolder results in the formation of intermetalliccompounds, either within the filler or at the in-terface between the solder and the parent mate-rial. These intermetallic phases have a pro-nounced effect on the mechanical properties ofthe joint. As constituents of solders, indium andtin have a dominant role in determining the in-termetallic compounds that formby reactionwiththe components: the compounds invariably con-tain either indium or tin. When both indium andtin are present, the composition of the parentmaterials (i.e., the material on the surface of thecomponents) determines which of these ele-ments is the predominant constituent of the re-sulting intermetallic compounds. For solder al-loys that contain neither element, intermetalliccompounds are not always formed and those thatdo necessarily depend on the composition of boththe solder and of the parentmaterials. Themecha-nism by which these intermetallics form andsome of the implications of their presence injoints are discussed in section 2.3 in this chapter.

Each of the major elements in the lowest melt-ing point solders—bismuth, indium, lead, andtin—confers different properties to the filler met-als, often in a manner that is not obvious. Forexample, the fact that solders tend to be based oneutectic composition alloysmeans that theirmelt-ing points are determined by the eutectic reaction

Chapter 2: Solders and Their Metallurgy / 51

rather than by the melting points of the indi-vidual elements present. There is no simple re-lationship between the melting temperatures ofthe constituent elements and the eutectic tem-perature. Therefore, while the elements fall intothe following sequence, lead > bismuth > tin >indium, when ranked in descending order ofmelting point, the solder alloys do not fall intothis pattern; those containing bismuth have thelowest melting points of all. Despite such com-plexities, it is possible to make some generali-zations about the role of each element in a solder.To this list is also appended antimony, not thatthere are common solders based on this element,but because it is sometimes a minor constituentin other solder alloy families.

Tin is a preferred constituent of many solderalloysbecause it confersfluidity,benefitswetting,enhancesmechanicalandphysicalproperties,andpossessesexceptionally lowvaporpressure.Afterlead, cadmium, and zinc, tin is the least expensiveingredient of solder alloys. Nevertheless, thepriceof tin isstill some30timesthatof lead,whichaccounts for the popularity of lead-rich lead-tinsolders. By way of comparison, bismuth and an-timonyareabout thesamepriceastin,withindiumand silver being over an order ofmagnitudemoreexpensive still.However, tin-bearingsolders tend

to form brittle compounds on reaction withmanyparent materials and metallizations used in engi-neering, particularly copper and gold. Silver isoneof the fewexceptions,with silver-tin interme-tallic phases being comparatively ductile. There-fore, considerable caremustbe taken indesigningthe joining process and specifying the service en-vironment so as to restrict the formation of inter-metallic compounds to concentrations belowthose that would otherwise weaken and embrittlethe joints. This point is discussed in connectionwith the phase diagrams of the relevant alloy sys-tems in section 2.3 in this chapter.

Indium and lead are the two softest and mostductile constituents of solder alloys. Despite theirinferior mechanical properties, solders with highlead concentrations find wide application be-cause they are the least expensive and the easiestto use of the high-melting-point solder alloysavailable. The indium-bearing solders are par-ticularly attractive for use with gold metalliza-tions because these are not readily dissolved andthe interfacial phases that form are compara-tively ductile, so that joints are not embrittled bytheir presence. The low level of gold erosionstems from a combination of the steep slope ofthe liquidus line on the phase diagram betweenindium and gold (see Fig. 2.1) and the formation

Fig. 2.1 Gold-indium phase diagram


of a thin, continuous intermetallic compoundAuIn2 between the molten solder and the goldmetallization. This layer of compound then actsas a barrier against significant further gold dis-solution from taking place and results in the pro-file of the erosion curves shown in Fig. 2.2. Thus,indium-containing solders can be reliably usedin conjunction with very thin gold metalliza-tions.

Silver is a constituent of several solder alloys,but only as a minor proportion, not simply onaccount of its price premium, but because higherconcentrations (more than about 5%) result in asudden increase in the liquidus temperature to-ward those of the silver-bearing brazes. Smalladditions of silver are used primarily to enhancemechanical properties of solders and joints andto promote fluidity by destabilizing native sur-face oxides on the molten solder. Silver oxide isnot stable in air above 190 °C (374 °F). Unfor-tunately, however, owing to industrial pollutionsilver tarnish often contains sulfide, which is farmore stable against thermal degradation. TheAg-96Sn solder has among the best mechanical andphysical properties of any low-melting-point sol-der alloy [Harada and Satoh 1990]. Silver-containing solders tend to be preferredwhen join-ing silver-coated components because thepresence of silver in the solder reduces the rateand extent of scavenging from the metallization,as shown by the data in Fig. 2.3.

Bismuth is the most brittle constituent of thecommon solders and, for this reason, few solderscontainmore than 50% of this element. Bismuth-bearing alloys comprise the majority of the low-est melting point solders, as can be seen fromFig. 1.3. Bismuth exhibits the unusual charac-teristic of expanding on freezing, enabling sol-ders to be tailored to have essentially zero liquid-

to-solid volume contraction by appropriatelyadjusting the bismuth concentration. It has beenclaimed that this property can confer benefits inmaking hermetic soldered joints [Dogra 1985].Although alloys such as the bismuth-tin eutecticsolder can be prepared in a manner that rendersit soft and ductile, by rapid solidification, it willsubsequently embrittle, even at room tempera-ture, owing to changes in the atomic lattice spac-ing of the bismuth phase that is responsible forthe solid-state expansion [Hare, Corwin, and Re-imer 1985]. The mechanical properties of bis-muth-tin solders may be improved considerablyby the addition of 0.5% Ag. This low concen-tration of silver does not appreciably affect themelting range of the solder, but acts as a veryefficient grain refiner that improves the tensileductility of the solder by a factor of three andreduces the susceptibility of the solder to strain-rate-dependent deformation behavior [McCor-mack et al. 1997]. The low melting point andinferior fluidity of the bismuth–bearing soldersimpose constraints on the joint design and pro-cessing conditions. For example, their relativelylow melting temperature means that aggressiveinorganic fluxes are needed to chemically cleanthe surfaces of the parent materials.

Antimony is often found as a deliberate minoraddition (<10%) in many solders, particularlylead-tin alloys. This is because the addition im-proves some key mechanical properties. In lead-tin solder, antimony causes solid-solution

Fig. 2.2 Erosion of a gold metallization by molten indium asa function of reaction time and temperature. Similar

results are obtained for indium-base solders, including gold-indium, silver-indium, indium-lead, and indium-tin.

Fig. 2.3 Substantial reduction of the dissolution rate of silverin lead-tin eutectic composition solder obtained by

small additions of silver to the alloy. Adapted from Bulwith andMackay [1985]


strengthening, up to a concentration of about 3%(see Fig. 2.4). As the proportion of antimonyincreases, SbSn cuboids form in the solder and,while this produces further improvements inproperties such as creep, these benefits are offsetby a tendency toward brittleness that is respon-sible for high and undesirable scatter on reli-ability data [Tomlinson and Bryan 1986].

Gold is the most expensive major constituentof solders and, for that reason, the applicationsof gold solders tend to be limited to high-valueelectronics, photonics, and jewelry manufacture.One of the chief attractions of these solders istheir melting point, which falls within the 300 to500 °C (570 to 930 °F) gap that separates theupper limit of the lead-base solders from thelower limit of the available aluminum-bearingbrazes.

Zinc forms a eutectic alloy with aluminum atthe composition,Al-94Zn (melting point 381 °C,or 718 °F). It is the only low-melting-point metalthat doesnotgrossly enhance thecorrosionof alu-minum alloys because its electrode potential isclose to that of aluminum, as indicated by the datainTable2.1.Accordingly, zinc-aluminumsoldersfind use in the joining of aluminum alloys. How-ever, the refractory nature of aluminum oxide re-quires theuseofaggressivefluxes topromotewet-ting by zinc-base solders, and this creates otherproblems, namely highly corrosive flux residues.

This outline of solder families is brief andfairly generalized. Many solder alloys haveunique properties, confirmed in numerous stud-ies, that show that there is often no regular pat-tern among solder composition, melting point,and joint properties. Indeed, the ranking orderfor even simple mechanical tests can be radicallychanged merely by altering the test conditionsused [Tomlinson and Fullylove 1992].

A list of some of the more common binary,ternary, and quaternary composition solder al-loys is given in Table 2.2, and the associatedphase diagrams of those in widest use are shownin Fig. 2.5 to 2.15. Most of these alloys areductile and can be mechanically worked to pro-duce preforms of virtually any desired geom-etry. The majority of solders are based on eutec-tic compositions, for reasons that are explainedin section 2.3 in this chapter. Many solder manu-facturers also offer off-eutectic compositions inthese and other alloy systems. They are in-cluded in the product range because they areeither less expensive, easier to fabricate as wireand foil, or have a melting range, which issometimes a technical advantage. It is usuallythe case that fluidity and mechanical propertiessuffer on moving to an off-eutectic composition,but sometimes these characteristics are desir-able. For example, the high-lead solders areexpressly used below their liquidus temperaturefor the purpose of bridging wide gaps. On theother hand, the indium-lead alloys offer a con-tinuum of solder compositions with intermedi-ate melting temperatures.

Although not included in Table 2.2, referencehas been noted in the literature to one other high-melting-point solder. Details are sparse, but itappears to be a quaternary alloy based on thecopper-silver eutectic braze with substantial ad-ditions of indium and tin to depress the liquidusand solidus temperatures. The melting point ofthe solder is claimed to be 450 °C (842 °F), and,therefore, it technically qualifies as a solder[Hirakawa, Tanahashi, and Terasawa 1995]. Fur-ther details on conventional brazes based on cop-per-silver and other alloys can be found in theplanned companion publication Principles ofBrazing.

Fig. 2.4 Shearstrengthof�-brassjointsmadewithlead-tinsol-der containing varying concentrations of antimony

Table 2.1 Electrode potential of selectedelements at 25 ºC (77 °F)Element Electrode Potential, V

Gold �1.50Silver �0.80Copper �0.34Hydrogen 0.00Lead �0.13Tin �0.14Nickel �0.25Cadmium �0.40Iron �0.44Zinc �0.74Silicon �1.30Aluminum �1.66Magnesium �2.37


Fig. 2.5 Indium-tin phase diagram

Table 2.2 Low-melting-point eutectic composition alloys used as soldersComposition, wt% Eutectic temperature

Ag Bi In Pb Sn Other °C °F

. . . 49.0 21.0 18.0 12.0 . . . 57 135

. . . 33.7 66.3 . . . . . . . . . 72 162

. . . 52.0 . . . 32.0 16.0 . . . 95 203

. . . 67.0 33.0 . . . . . . . . . 109 228

. . . . . . 51.0 . . . 49.0 . . . 120 248

. . . 55.5 . . . 44.5 . . . . . . 123 253

. . . 57.0 . . . . . . 43.0 . . . 139 2823.0 . . . 97.0 . . . . . . . . . 144 291. . . . . . 96.0 . . . . . . 4.0Zn 144 2915.0 . . . 80.0 15.0 . . . . . . 142–149(a) 288–300(a). . . . . . 99.5 . . . . . . 0.5Au 156 313. . . . . . 75.0 25.0 . . . . . . 165–170(a) 330–340(a)1.5 . . . . . . 36.0 62.5 . . . 179 354. . . . . . . . . 38.0 62.0 . . . 183 361. . . . . . 50.0 50.0 . . . . . . 178–210(a) 352–410(a). . . . . . . . . . . . 91.0 9.0Zn 198 388. . . . . . . . . 85.0 . . . 15.0Au 215 4193.5 . . . . . . . . . 96.5 . . . 221 430. . . . . . . . . . . . 99.3 0.7Cu 227 441. . . . . . . . . . . . 95.0 5.0Sb 235–245(a) 455–475(a). . . . . . . . . 85.0 3.5 11.5Sb 240 46425.0 . . . . . . . . . 65.0 10.0Sb 232–234(a) 450–453(a). . . . . . . . . 88.9 . . . 11.1Sb 251 484. . . . . . 25.0 75.0 . . . . . . 255–265(a) 490–510(a)2.5 . . . 5.0 92.5 . . . . . . 300–310(a) 572–590(a)2.5 . . . . . . 97.5 . . . . . . 304 5791.5 . . . . . . 97.5 1.0 . . . 309–310(a) 588–590(a). . . . . . . . . 99.5 . . . 0.5Zn 318 604

(a) These alloys are not eutectic compositions, but have been included on account of their narrow melting range and their industrial exploitation.


2.1.1 Lead-Tin SoldersThis survey necessarily startswith lead-tin sol-

ders because they account for about 95% of allsoldered joints. Soldered joints are made on avast, global scale. It is estimated that more than60,000 tonnes per annum (6 � 107 kg/yr or54,000 tons/yr) of lead-tin solder are used in theelectronics assembly industry alone. The ben-eficial characteristics of lead-tin alloys have beenappreciated since at least Roman times, and theelder Pliny, writing in the first century A.D., inhis Historia Naturalis [Rackham 1952] specifi-cally mentions an alloy containing two parts ofblack lead (modern lead) and one part of whitelead (i.e., tin) used for soldering pipes. He alsoremarks that the price of this alloy is 20 denariiper pound. It is interesting to note that this worksout at roughly $100/kg, assuming that gold hasmaintained its purchasing power since Pliny’sday. Today’s price for the same solder is lowerby an order of magnitude.

Lead-tin alloys offer the following advantagescompared with other solders:

• Superior wetting and spreading character-istics, compared with most other solders,especially those containing bismuth, anti-mony, and indium. The difference is quan-tified in Fig. 1.14.

• Relatively inexpensive to produce and use.Lead, in particular, is inexpensive comparedwith other solder metals, notably bismuth, in-dium, and silver. Moreover, the fabricationcostsare lowincomparisonwith thoseofmostsolders as lead-tin alloys are readily coldworked.

• Greater versatility. Lead-tin solders readilywet a wide range of metals to produce soundjoints with minimal substrate erosion. Sol-dering to thick gold metallizations representsone of the few situations in which restrictionsneed to be observed in order to prevent em-brittlement through the formation of unfa-vorable intermetallic phases. The reaction be-tween lead-tin solders and goldmetallizationsis discussed in detail in section 2.3 in thischapter.

• Ready application as coatings on compo-nents by electroplating. The lead-tin alloysystem is the only one for which low-costelectroplating technology has been devel-oped successfully.

• Satisfactory mechanical properties for manyapplications. In terms ofmechanical strength,stiffness, and fatigue resistance, only silver-tin solders are superior and even then bymuchless than an order of magnitude and in par-ticular circumstance.

Fig. 2.6 Bismuth-tin phase diagram


In most other respects, such as corrosion re-sistance and electrical and thermal conductivity,there is little to distinguish lead-tin alloys fromthe majority of other low-melting-point solders(see Table 5.19 in Chapter 5, section 5.7).

Lead-tin solders are seldom pure, but oftencontain other elements, either as incidental im-purities or as deliberate additions designed tomodify specific properties. Three metal addi-tions are routinely encountered; antimony, bis-muth, and silver.

Antimony is frequently present to amaximumconcentration of about 1%. It has a beneficial rolein improving the mechanical properties of jointsthough a solid-solution-strengthening mecha-nism.Antimony can substitute for double the pro-portion of tin, without greatly widening the melt-ing range of the solder and, therefore, is oftenfavored by solder manufacturers as it decreasesmaterials costs. At higher concentrations of anti-mony, there are solder alloys based on the ternarysystem Pb-Sb-Sn. The liquidus projection of thisalloy system is given in Fig. 2.13, from which itcan be seen that there is a ternary eutectic at 240°C containing 11.5% Sb.

Bismuth is added for similar functional rea-sons as antimony. The maximum concentrationis usually limited to about 3%, as higher levels

result in a significant widening of the meltingrange and a noticeable impairment to wettingand spreading behavior. In addition, there is arange of low-melting-point ternary solders basedon Bi-Pb-Sn alloys in which bismuth is presentin high concentrations. The liquidus surface ofthis alloy system is given in Fig. 2.16. Smallquantities of bismuth and also antimony areadded to lead-tin solders to prevent degradationat low temperatures through amechanism knownas “tin pest.” Further details of this failuremecha-nism are given in section 2.2 in this chapter.

Silver. Lead-silver-tin solders find applicationfor soldering silver-coated surfaces. The addi-tion of 2% Ag to the lead-tin eutectic alloy hasa marked effect in reducing the erosion of silvercoatings, as shown by the data in Fig. 2.3. Thepresence of silver in lead-tin solders results in theformation of a fine dispersion of the relativelyductile intermetallic compound Ag3Sn, on so-lidification, which boosts the mechanical prop-erties of joints. Small quantities of copper andgold have a similar beneficial effect, as describedin section 2.2 in this chapter. However, theseelements are not usually present as deliberatealloying additions since they are incorporatedautomatically when soldering copper- and gold-coated components, respectively.

Fig. 2.7 Silver-indium phase diagram


All of the aforementioned additions to lead-tinsolder, in small quantities, enhance the strengthof joints to copper because they alloy preferen-

tially with some of the tin and thereby serve toreduce the thickness of the copper-tin interme-tallic phases. These form at copper-solder inter-

Fig. 2.8 Lead-tin phase diagram

Fig. 2.9 Silver-tin phase diagram


faces and can be a source of mechanical weak-ness (see section 2.3 in this chapter) [Quan et al.1987].

Attempts have been made to improve the me-chanical properties and resistance to failure oflead-tin solders through creep and fatigue by a

Fig. 2.10 Antimony-tin phase diagram

Fig. 2.11 Silver-lead phase diagram


number of strengthening mechanisms. Thesenovel alloys are discussed in Chapter 5, sections5.5, 5.6, and 5.8.

2.1.2 Other Tin-Base SoldersTwo tin-base solders find wide industrial ap-

plication. These are silver-tin eutectic and a Ag-Sb-Sn solder often referred to as “Alloy J.”

Silver-Tin. Silver-tin is a eutectic alloy thathas a fixed point generally accepted to be 221 °C.The composition of the eutectic is close to Ag-96.5Sn.Alloys offered by manufacturers will de-viate from this by up to 1.5% on account of cost,but this results in awidening of themelting range.The silver-tin eutectic forms the basis of virtuallyall of the “lead-free” alloys offered as replace-ments for lead-tin solder (see Chapter 5, section5.1). In many respects, this solder is the closestbinary alloy to lead-tin solder. Themelting pointsare not greatly different, and their wetting andspreading characteristics are comparable to thoseof lead-tin. So are their mechanical properties,although silver-tin alloys are generally slightlysuperior in this respect for service temperaturesbelow 150 °C (300 °F). Both solders form closelyrelated intermetallic compounds on wetting ofthe common engineering parent materials and

metallizations. The hue and luster of solder filletsare visually indistinguishable, but the surfaces oflead-tin solder fillets are often smoother as thissolder contains a more even ratio of the twoconstituents and hence a finer grain size.

Alloy J. The majority of solders are referredto in manufacturers’ catalogs by a name or ter-minology that gives a clue to the alloy compo-sition, although part of the product range of onewell-known solder manufacturer is denoted by anumbering system that can only be decoded withaid of a reference card.More recently, it has beennoted that, with the advent of lead-free solders,marketing departments have had some role in thenaming of new products. These names are de-signed to convey an impression of green andenvironmentally friendly products rather thantechnical information about the product. One ofthe cryptic solder names in the literature is “Al-loy J,” and particular reference is made to thisrelatively prominent alloy here to elucidate itskey features.

Alloy J has the composition 25Ag-65Sn-10Sb. It has a narrow melting range, usuallycited simply as 233 °C (451 °F) [Olsen and Span-jer 1981, Mackay and Levine 1986]. Containing65% Sn, Alloy J qualifies as a tin-base solder.The alloy was originally developed as an alter-

Fig. 2.12 Indium-lead phase diagram


native solder for silicon semiconductor die-attach with ostensibly more favorable mechani-cal properties, melting range, and price than thegold-tin and gold-silicon solders. The name de-rives from the systematic search of theAg-Sb-Snsystem that was undertaken for suitable candi-dates—Alloy J was the tenth alloy compositionevaluated and the first one found largely to meetthe target requirements, that of substantially su-perior performance compared to lead-tin eutecticunder conditions of power cycling of semicon-ductor die, as shown in Fig. 2.17 [Mackay andLevine 1986].

Reference to the ternary phase diagram forthis alloy system (Fig. 2.15) shows that the 25Ag-65Sn-10Sb solder is a three-phase mixture of tin,Ag3Sn, and SbSn. The two intermetallic com-pound constituents are relatively hard and brittle.When produced and used conventionally, thepresence of these intermetallics makes the fab-rication into convenient forms such as wire andpreforms extremely difficult and is detrimental tothe joint properties. This problem has been re-solved through the production of the alloy usingrapid-solidification technology (see Chapter 1,section 1.3.2.2). The rapid-solidification processyields an alloy with a substantially finer micro-

structure with a grain size below 0.3 μm (12 μin.)compared with 10 μm (400 μin.) when producedby conventional ingot casting. The rapid-solidification process also extends the solid solu-bilities of both silver and antimony in tin so thatthe volumetric proportion of intermetallic phasein the alloy is correspondingly decreased. Thesetwo features result in improved ductility and areduction in hardness.

Provided the soldering cycle is short and thesubsequent service temperature is not excessive,the benefits of a refined grain size persists injoints made using solder prepared by rapid-solidification processing. The reasons for this arenot fully understood, but the net effect is thatsilicon die soldered to ceramic packages usingrapidly solidified Alloy J have improved resis-tance to fatigue failure on power cycling oversolder produced by conventional casting, as in-dicated by the data presented in Fig. 2.18 [Pi-namaneni and Solomon 1986].

2.1.3 Zinc-Bearing SoldersOne of the major applications of solders con-

taining zinc is for joining aluminum and its al-loys [Finch 1985]. This metal plays an additional

Fig. 2.13 Liquidus surface of the Pb-Sb-Sn system


role in helping to disrupt alumina surface layersand so helps to promote wetting by the solder. Aflux-cored, zinc-base solder that is intended forjoining aluminum is available commercially [Ru-bin 1982]. Such joints are obviously susceptibleto corrosion and tend to be used for making elec-trical connections to functional aluminum partsrather than structural assemblies. The key tomak-ing a successful soldered joint to aluminum is theformulation of the flux used. This topic is dis-cussed in Chapter 3, section 3.2.2.1. When sol-dering to aluminum parts, a frequently over-looked consideration is the combination of highthermal conductivity and expansivity of thismetal, coupled with low heat capacity. There-fore, particular attention needs to be paid to theheating method to minimize thermal gradientsand thermally induced distortion.

One of the zinc-bearing solders that is widelyused for joining aluminum is the eutectic com-position alloy Al-94Zn (melting point is 381 °C,

or 718 °F). The aluminum-zinc phase diagramis given in Fig. 2.19. Often a few ppm ofgallium is added to the solder to aid wetting,molten gallium being one of the few metals thatwill not ball up on alumina, which is prevalenton the surface of aluminum alloys. Spreading ofthe solder is promoted by a reaction betweengallium oxide and alumina, rather than betweenthe respective metals. The relatively small de-pression in the melting temperature of alumi-num achieved by alloying with a large propor-tion of zinc has one negative consequence,namely the occurrence of significant erosion ofthe aluminum components in the region of thejoint. This can be a problem in situations whereit is required that component geometry is notperceptibly changed by the joining operation.Zinc-bearing alloys are not suitable for joiningprocesses that entail reduced pressure atmo-spheres, owing to the inherent volatility of thiselement.

Fig. 2.14 Liquidus surface of the Ag-Cu-Sn system. Adapted from Petzow and Effenberg [1988]


All of the zinc-bearing solders contain highpercentages of zinc, generally in the range 75 to95%, and their solidus temperatures cover thetemperature band 197 to 419 °C (387 to 786 °F).The alloying elements present in significant per-centages are aluminum, cadmium, copper, andtin. A representative list of these solders is givenin Table 2.3. Pure zinc is not used as a filler metalbecause it tends to ball up on heating when usedin normal atmospheres, whereas when smallamounts of other elements are added, wetting ofsurfaces is more readily achieved. The cadmium-containing alloy listed in Table 2.3 is now gen-erally avoided because cadmium fume is classedas a health hazard.

The choice of alloying additions is made ongrounds of reducing the erosion of the substratemetals and, as far as possible, of improving thespreading and flowing characteristics. In gen-eral, the additions cannot be entertained as melt-ing-point depressants because theymostly do notreduce the liquidus temperature, only the solidustemperature.

Over the years, considerable research has beendevoted to developing improved new solders

based on zinc [Harrison and Knights 1984] be-cause these alloys have the following potentialattractions:

• They can be used at sufficiently low tem-peratures so as not to destroy the work-hardened strength of the copper alloys. Yet,at the same time, the joints that are producedare typically two to four times stronger thanthose obtained with the common solders.

• The zinc solders are compatible with galva-nized steel components.

• They are inexpensive in relation to most sol-ders, being approximately one-quarter thecost of the lead-tin alloys, for a given volumeof filler metal.

• They are lightweight, having approximatelyhalf the density of lead-tin eutectic solder.

• They possess high thermal conductivity, ex-ceeding 100 W/m • K (60 Btu/h • ft • °F).

• They are not hazardous: unlike cadmium andlead, the effects of small amounts of zincabsorbed in the human body are noncumu-lative and temporary. Zinc is one of the vitaltrace elements in the human diet (as are many

Fig. 2.15 Liquidus surface of the Ag-Sb-Sn system. Adapted from Petzow and Effenberg [1988]


elements that are toxic in larger concentra-tions).

The factors that have greatly limited the adop-tion of the zinc solders include:

• The potential strength of joints to copper al-loys are compromised by the presence of em-brittling copper-zinc compounds, which formas interfacial phases. For this reason, the zinc-base solders that have been formulated foruse with copper alloys often contain numer-ous minor alloying additions to modify thegrowth of these phases.

• Zinc alloys in joints to “heavy” base metals(e.g., those based on copper or iron) are sus-ceptible to galvanic corrosion, a problem thatthey share with aluminum filler alloys.

• The low intrinsic material costs are largelyoffset by high fabrication costs arising fromthe low ductility of most zinc-base alloys.

• The relatively high volume contraction onsolidification, which is typically 1 to 1.5vol%, is detrimental to joint filling and cancause stress concentration in joints.

• The zinc alloys generally exhibit poor flow-ing characteristics. The high oxidation rate of

Fig. 2.16 Liquidus surface of the Bi-Pb-Sn system

Fig. 2.17 Degradation of joint quality, as measured by through-thickness electrical resistivity for silicon semiconductor die attachedjoints using three different solders and subject to power cycling


zinc, which is of the order of 2 to 3 μm/min(80 to 120 μin./min) at 400 °C (750 °F), in air,coupled with high surface tension and vis-cosity (both approximately double those oflead-tin alloys), are the main factors.

• The high vapor pressure of zinc means thatthese solders cannot be used in reduced pres-sure atmospheres because the volatility of

this element exacerbates the formation ofvoids in joints and contaminates furnaceequipment.

• The high affinity of zinc for oxygen requiresthe use of aggressive fluxes when solderingin ambient atmospheres which, in turn, leadsto cleaning and corrosion problems. The fluxefficacy is highly specific to the particular

Fig. 2.18 Cumulative failure data for TO-220 silicon semiconductor die, subject to power cycling. Joints were fabricated usingpreforms prepared using conventional casting (dark bars in chart) and mechanical working or rapid solidification tech-

nology (white bars on chart). Adapted from Pinamaneni and Solomon [1986]

Fig. 2.19 Aluminum-zinc phase diagram


parent metal/solder combination on which itis used. This makes it extremely difficult toachieve tolerant soldering processes usingzinc-base filler metals, although consistentresults do appear to have been achieved withnickel and copper substrateswith fluxes basedon L-glutamic acid and dimethylammoniumchloride using process temperatures as low as250 °C (480 °F).

Formany situations, the disadvantages of zinc-bearing solders outweigh their prospective ben-efits, which accounts for the fact that they havenot found much favor in industry. Nevertheless,further development continues. One example isthe development of fluxes based on organic tincompounds. These chemicals are designed to de-compose at zinc soldering process temperaturesand coat the substrate and filler metal with a thinfilm of tin. By this means, contact angles as lowas 20° can be obtained on copper substrates whensoldered in air using tin-zinc eutectic solder[Vaynman and Fine 2000]. Another example isthe development of zinc-base solders as lead-free substitutes for the high-lead family of fillermetals. In particular, an alloy of compositionAl-3Ga-3Mg-90Zn, which has a melting rangeof 309 to 347 °C (588 to 657 °F), has beenproposed as a hard solder for die-attach appli-cations in the electronics industry [Shimizu et al.1999]. Flux pastes of zinc-base solders have alsobeen developed [Noguchi 1999].

2.1.4 Gold-Bearing Solders

Gold is unusual in that it is the only elementon which both brazes and solders are based; thatis, this element is the major constituent of bothsolders and brazes. Further details about gold-base brazes are given in the planned companionvolume Principles of Brazing; also discussed isthe use of gold solders for jewelry applications.The gold-bearing solders are all gold-rich alloysof eutectic composition and have melting pointsbetween 278 and 363 °C (532 and 685 °F). Theyare listed in Table 2.4 and their associated phasediagrams are given in Fig. 2.20 to 2.23. In viewof their high cost, the applications of these alloystend to be limited and specialized. One of thechief attractions of these solders is their meltingpoint, which falls within the 300 to 500 °C (570to 930 °F) gap that separates the upper limit ofthe lead-base solders from the lower limit of theavailable aluminum-bearing brazes.

Gold-bearing solders have the advantage ofbeing suitable for joining to gold-metallized com-ponents. Their compatibility with gold metalli-zations is explained more fully later in this sec-tion. Three of these solders (the gold-antimonyalloys being the exception) are usedwidely in theelectronics industry as high-temperature soldersfor attaching semiconductor devices into pack-ages and building hermetic enclosures for sen-sitive compound semiconductors and optical de-vices. Gold-antimony alloys are brittle, evenwhen prepared using rapid-solidification tech-nology and, as far as the authors are aware, arenot much used as filler metals.

The principal gold-base solders are consid-ered briefly in the paragraphs that follow.

Gold-Silicon and Gold-Germanium. On ac-count of their closely similar characteristics assolders, these alloys may be considered together.Gold-silicon alloys are primarily used in the formof a foil preform for bonding silicon semicon-ductor chips to gold-metallized pads in ceramicpackages. The alloy compositions used as sol-ders are slightly gold-rich with respect to theeutectic composition, generally at or close toAu-2wt%Si. This is deliberate because the eu-tectic alloy is too hard and brittle even to hot rollto foil. By making the alloy gold-rich, the pro-portion of the ductile gold phase in the micro-structure is sufficient to improve the mechanicalproperties to tolerable limits. Rapid solidifica-tion is unable to produce ductile foil becauserapid cooling of molten gold-silicon alloys re-sults in the formation of a large volume fractionof (metastable) gold-silicon intermetallic com-pounds, principally Au3Si, which render the foiltoo brittle to handle [Johnson and Johnson 1983].An alternative method of applying the solder tosilicon components is simply to coat the backsurface of the silicon die with a thin layer of gold,applied by a vapor-phase technique. On heatingthe gold-metallized silicon to above 363 °C (685°F), the resulting interdiffusion between the goldand silicon generates liquid filler metal in situ.

The Au-2Si solder is an exception to the ruleof eutectic composition alloys, possessing fa-vorable characteristics as filler metals. In par-ticular, it suffers from high viscosity when mol-ten. This characteristic is a direct consequence ofthe low temperature of the eutectic transforma-tion, relative to the high melting points of bothconstituent phases, together with the silaceousdross that tends to form on the surface of thealloy and the absence of suitable fluxes thatmightenhance wetting. This can be seen by comparing


the spreading behavior of the Au-2Si alloy, as afunction of temperature above its melting point,shown in Fig. 2.24 with similar data for other

common solder alloys given in Fig. 1.14. Poorfluidity of the solder increases the risk of inad-equately filled joints, whichmar the performanceand reliability of the product.

It is generally recommended that gold-siliconfoil is given a light etch in hydrofluoric acid andthen used within 30 minutes. This process stepremoves the surface silica and silicon and greatlyimproves the wetting and spreading behavior.While technically successful, this approach doesrequire handling of this rather aggressive acidand special operator training is usually requiredto satisfy health and safety regulations. Further-more, when using gold-silicon solders, precau-tions must be taken to ensure that the initial cool-ing rate of solidified joints does not exceed 5 °C/s(9 °F/s). If this condition is not satisfied, Au3Siis formed, and its subsequent decompositionwithtime to gold and silicon can produce crackswithin the joint, due to the associated volumecontraction [Johnson and Johnson 1984].

A number of alloying additions are known tobe capable of promoting solder spreading[Humpston and Jacobson 1990]. For the Au-2Sisolder, one of the most effective promoters ofspreading is tin. Results of spreading tests con-ducted under identical conditions, but with dif-ferent concentrations of tin in the solder showed

Table 2.3 Zinc-bearing alloys used as soldersMelting range

Solder composition ºC ºF

Zn-6Al 381 718Zn-5Al-2Ag-1Ni 381–387 718–729Zn-7Al-4Cu 379–390 714–735Zn-10Cd 266–399 511–750Zn-25Cd 266–370 511–700Zn-1Cu 418–424 784–795Zn-30Cu-3Sb-1Ag 416–424 781–795Zn-10Sn-1Pb 197–385 387–725Zn-70Sn 199–311 390–592Zn-2Ni 419–560 786–1040

The cadmium-containing alloys are no longer used for health reasons.

Table 2.4 Gold-bearing soldersEutectic temperature

Composition, wt% °C °F

Au-20Sn 278 532Au-25Sb 356 673Au-12Ge 361 682Au-3Si 363 685

Fig. 2.20 Gold-silicon phase diagram


that increasing the level of tin gave a progressiveimprovement in solder spread, as illustrated inFig. 2.25. Hardness measurements revealed that

this was accompanied by a softening of the alloy,with its hardness decreasing by more than 150HV, a welcome feature because it makes the al-

Fig. 2.21 Gold-tin phase diagram

Fig. 2.22 Gold-germanium phase diagram


loy more amenable to mechanical working intofoil and wire for solder preforms.

The simplified phase diagram of theAu-Si-Snalloy system, reproduced in Fig. 2.26, indicatesthat, by restricting the concentration of tin tobelow 8%, the melting temperature of the alloyis determined by the reaction among gold, sili-con, and the Au5Sn intermetallic compound,which occurs over a narrow melting range of356.5 to 358.0 °C (673.7 to 676.4 °F). This isclosely similar to the solidus temperature of theAu-2Si alloy (363 °C, or 685 °F), and switchingof this latter alloy for one additionally containingtin will therefore not upset subsequent manu-facturing steps.

Semiconductor die attach represents one ofthe main areas of application of gold-siliconand gold-germanium solders. Neither will di-rectly wet bare silicon. Silicon is a relativelyrefractory element that, when exposed to air,becomes covered with a stable, continuous, andextremely adherent layer of the native oxide(silica). For this reason, all conventional sol-der alloys are unable to wet silicon even withfluxes. In order to achieve wetting, it is nec-essary to either coat the back of precleaned diewith a solderable metallization, such as gold,or to manually scrub the molten filler betweenthe die and the package and thereby physically

displace the silica layer. The first option is notalways compatible with other stages of semi-conductor fabrication and so scrubbing is re-sorted to, using a protective shroud of inertgas. With the advent of automatic machines toperform this operation, it has evolved to be arapid joining method that delivers consistentquality joints.

A successful method of promoting wetting ofoxidized metal surfaces is to incorporate ele-ments into the filler that are capable of reactingwith the oxide to form an adherent bond [Crispinand Nicholas 1984, Xian and Si 1991]. Researchhas shown that addition of just 1% Ti to theAu-2Si and Au-20Sn solders is highly effectivein promoting wetting of bare silicon at solderingtemperatures of around 400 °C. Concentrationsof titanium of up to 2% were found to have littleeffect on the melting ranges of the alloys and areactually beneficial in reducing their hardness.Figure 1.29 shows a silicon die soldered into agold-metallized package using a gold solder withthe composition Au-2Si-8Sn-1Ti; the joint wasmade by heating in a nitrogen atmosphere with-out scrubbing. Wetting of the ceramic packageand the silicon die by the solder was found to behighly uniform.

The gold-germanium eutectic composition al-loy is sometimes recommended in place of the

Fig. 2.23 Gold-antimony phase diagram


gold-silicon solder where cost margins are par-ticularly tight. As can be seen from the phasediagram given in Fig. 2.22, it offers the benefitof lower gold concentration with little change inmelting point (note in some older texts the melt-ing point of the gold-germanium eutectic is er-roneously given as 256 °C). The gold-germa-nium binary eutectic contains 12.5 wt% Gecompared with 3.2 wt% Si in the case of thegold-silicon eutectic alloy. However, the price ofgermanium is much higher than that of silicon(because this element has to be produced as aminor by-product of zinc), but is considerablybelow that of gold. Historically, the price of ger-manium is about 5% that of gold, although in2002, it stood at 10% of the gold price, so thatthe saving in materials cost with respect to gold-silicon is fairly significant. As a solder, gold-germanium is no easier to use than its siliconequivalent. The Au-12Ge solder reportedly ex-hibits excellent wetting characteristics on cop-per, silver, and nickel surfaces [Hosking,Stephens, and Rejent 1999]. Where this solder isto be used in conjunction with steel parts, theparts should be plated with nickel prior to sol-dering to avoid the formation of iron germides.Hosking, Stephens, and Rejent [1999] have alsoshown that although Au-12Ge solder is brittle attemperatures up to 170 °C (340 °F), above about220 °C (430 °F), this alloy is reasonably ductile

Fig. 2.24 Spreading behavior of Au-2Si solder as a functionof the excess temperature above its melting point

Fig. 2.25 Spreading behavior of Au-2Si solder containingtin. Left to right: 0%, 4%, 8% Sn

Fig. 2.26 Gold-rich portion of the Au-Si-Sn phase diagram. The system can be divided into three regions bymelting point. The circleindicates the alloy composition selected as the solder. Adapted from Prince, Raynor, and Evans [1990]


and it is possible to substantially remove expan-sion mismatch stresses by a postsoldering heattreatment at this temperature. The same will pre-sumably be true of the Au-2Si solder, but thesoftening temperature is not known for this alloy.

Gold-Tin.TheAu-20Sn eutectic solder is hardand moderately brittle. These properties arisefrom the fact that its constituent phases are twogold-tin intermetallic compounds, namely,AuSn(�) andAu5Sn (�). TheAu5Sn phase is stable overa range of compositions and, consequently, hassome limited ductility that is imparted to thesolder alloy (approximately 2% at room tem-perature). Although difficult, this solder can behot rolled to foil and preforms stamped from it.By using rapid-solidification casting technology,it is possible to produce thin ductile foil, up toabout 75 μm (3 mil) thick and having an amor-phous microstructure. However, this state issomewhat unstable, and within about 30 min atroom temperature the rapidly solidified strip isindistinguishable in its mechanical propertiesfrom foil prepared by conventional fabricationmethods [Mattern 1989]. Nevertheless, this crys-tallization can be suppressed for about a year ifthe quench-cooledmaterial is stored under liquidnitrogen (�196 °C, or �320 ºF), and for closeto one month at �20 °C (�4 °F), so that it ispossible to manufacture shaped foil preformswhile the strip is still ductile, which can then beeither immediately placed in a jig or returned tocold storage. The alloy can also be readily gasatomized to form spherical powder and, becauseit is relatively inert, will survive for long periodsin an organic binder medium (optionally con-taining a flux) without degradation. For this rea-son, the solder alloy is often used in the form ofpaste, provided the components being joined arecompatible with the chemicals involved.

An alternative method of introducing the Au-20Sn solder into an assembly is to selectivelycoat the joint area with a thick layer of gold,overlaid with a thinner layer of tin in the thick-ness ratio of 3Au to 2Sn. If electroplated, thelayers must be deposited in this order owing totheir electrochemical potentials. By subjectingthe tin to a light acid etch and then immediatelyapplying an outer layer of gold by evaporation orimmersion plating, the predeposited solder issuitable for use without flux and can be endowedwith a shelf life of several months. Other meth-ods of producing ductile preforms of gold-tinsolders are described in Chapter 4, section 4.1.5.

The principal applications of the Au-20Snsolder are die attach of gold-metallized chips

and the hermetic sealing of ceramic semicon-ductor packages. The die-attach process that isused when the sealing operation is performedwith the Au-20Sn solder is usually a higher-melting-point gold-silicon solder. Thus, thegold-silicon and gold-tin solders are used in theinitial joining operations of a step-soldering se-quence, which ends with the soldering of thepackages containing the chips onto printed cir-cuit boards. This final soldering operation ismost frequently performed using a lead-tin orlead-free solder at a process temperature in theregion of 240 °C (464 °F).

Of the gold-bearing solders, only theAu-20Snalloy has a significant degree of fluidity whenmolten. In their major field of application, that ofsemiconductor manufacture, the use of fluxes topromote spreading is not usually permitted. In-stead, the techniques and methods normally ap-plied to fluxless joining are invoked to effectreliable joints. The options available and under-lying principles of fluxless joining are describedin detail in Chapter 3, section 3.3.

The gold-tin eutectic solder tends mostly to beused for soldering to thick gold metallizations.There are some difficulties with this, notably thedissolution of gold increases the freezing pointand prevents good spreading. Identification ofwettable, but insoluble barrier metallizations forthis solder is therefore desirable. Copper, nickel,chromium, and nichrome are wholly unsuitableas they all dissolve readily in the Au-20Sn alloy.Palladium is a good wettable metallization forthe molten solder, but, on extended aging in thesolid state, Kirkendall voids form at the interfacebetween the residual palladium and the Pd3Sn2intermetallic compound [Anhock et al. 1998].The solubility of platinum in Au-20Sn solder isstrongly temperature dependent. At normal sol-dering temperatures with short process cycles, athin layer of platinum (200 nm, or 8 μin.) pro-vides a readily wettable and stable barrier layer[Wada and Kumai 1991]. However, if large su-perheats or prolonged thermal excursions arenecessary, then published work suggests that co-balt is a better choice. The dissolution rate ofplatinum in gold-tin eutectic at 330 °C (626 °F)is 10 nm/s (0.4 μin./s), but nearly an order ofmagnitude slower for cobalt. Both gold and tinhave very low solubility for cobalt. The me-chanical properties of cobalt-tin intermetallicsare not known, but, as the layers formed duringnormal soldering cycles will be extremely thin,they are unlikely to influence joint propertiessignificantly [Park et al. 2002].


A high-melting-point quaternary filler metalbased on the gold-tin eutectic became availablerelatively recently, intended for joining to thick-film metallizations [DuPont Electronics 1998].The design strategy appears to have been to sub-stitute some of the gold by silver and copper,both of which form solid solutions with gold.The exact composition and melting range havenot been disclosed, but the recommended pro-cess temperature is 400 °C (750 °F), comparedwith 350 °C (660 °F) for the Au-20Sn binarycomposition from the same supplier. This im-plies the melting range is probably somewherearound 300 to 350 °C (570 to 660 °F). Amongthe claimed benefits of the substitution are highermelting point, greater ductility, and improvedwettability.

A noneutectic gold-indium alloy of composi-tion 82Au-18In is offered commercially by anumber of manufacturers. The alloy is gold-richto the nearest eutectic and deliberately so be-cause of the poor mechanical properties of theeutectic composition alloy. The demand for thissolder and its inclusion in catalogs stems from itsmelting range of 451 to 485 °C (844 to 905 °F),which makes it the highest melting point solderavailable. However this gold-indium alloy is adifficult solder to work with, as it is inferior togold-silicon and gold-tin alloys with regard to itswetting and flow characteristics, as indicated inTable 2.5. Also, there is no developed flux forthis solder and it must be used fluxless, with allthe attendant problems and limitations, althoughthe process temperature is sufficiently high thatthis is one of the rare exampleswhere a hydrogen-containing atmosphere is chemically active to-ward surface oxides on the solder (see Chapter3, section 3.1.3).

2.1.5 High-Lead Solders

High-lead solders can be loosely defined asalloys containing 95% or more Pb and having

solidus temperatures above 300 °C (570 °F). Arepresentative list of these solders is given inTable 2.6. They possess small melting ranges,usually, about 5 to 10 °C (9 to 18 °F), and theirusage arises largely on account of their meltingpoint. They are, in effect, inexpensive alterna-tives to the Au-20Sn eutectic alloy or, with care,can be used in a step-soldering sequence inwhichthe next filler metal is the gold-rich solder. Thesehigh-lead solders differ in one important respectto the Au-20Sn alloy in that they are extremelysoft and creep readily. This can be either an ad-vantage or drawback, depending on the appli-cation and the function that the solder joint isrequired to provide.

The high-lead alloys possess other advantagesas solders. They are suitable for use with gold-plated components without the risk of cata-strophic embrittlement through the formation ofcontinuousAuSn4 intermetallic phase at the jointinterface.Although there is a small percentage oftin in someof the high-lead solders, the alloy con-stitution allows for appreciable gold dissolutionbefore the formation of undesirable gold-tinphases because the small amount of tin present ismostly incorporated in the lead solid solution.With small adjustments of composition, somehigh lead-base solders have wide melting rangeand consequently have excellent gap-bridgingcharacteristics. In the development of electronicpackaging fabrication technology, the Ag-92.5Pb-5Sn alloy applied in the form of a propri-etarysolderpastewas found toproduceconsistenthermetic seals for feed-throughs into hybrid mi-crowave packages despite joint clearances vary-ing from below 20 μm (0.8 mil) to more than 150μm (60mil), as can be seen in Fig. 2.27 [Jacobsonand Sangha 1997].

An obvious omission from any list of high-melting-point lead-rich solders is pure lead. Thereason for this is pure lead has atrocious wetting

Table 2.5 Contact angle of Au-12Ge andAu-18In solders on common metal substrates

Contact angle, degrees

Substrate Au-12Ge(a) Au-18In(a)

Copper 4 35Silver 4 58Nickel 5 43

(a) The process temperature for gold-germanium was 430 °C (806 °F) andgold-indium 550 °C (1022 °F), with 5 min hold at peak temperature in a vacuumof 7 mPa ( 5 � 10�5 torr). Adapted from Hosking, Stephens, and Rejent [1999].

Table 2.6 Selected high-melting-pointlead-rich solders

Melting range

Solder composition ºC ºF

98.5Pb-1.5Sb 310–322 590–61295Pb-5Sn 308–312 586–59492.5Pb-5In-2.5Ag 300–310 572–59097.5Pb-1.5Ag-1Sn 309–310 588–59093Pb-3Sn-2In-2Ag 304–305 579–58197.5Pb-2.5Ag 304 58195Pb-5In 300–313 572–59593.5Pb-5Sn-1.5Ag 296–301 565–57490Pb-5In-5Ag 290–310 554–59092.5Pb-5Sn-2.5Ag 287–296 549–565


and spreading behavior. This is a consequence ofthe fact that lead does not alloy readily with mostengineering metals and metallizations, so thatthere is only marginal thermodynamic drivingforce for wetting and spreading. Indeed, this isthe sole reason for lead being unsuitable on itsown as a solder and requiring alloying additions.

The list of high-lead alloys given in Table 2.6is by nomeans exhaustive.Many other variationsexist, but the basis of their formulation—namely,small additions of silver, tin, indium, antimony,and other elements—is similar. Data sources citeslightly different solidus and liquidus tempera-tures for many of these alloys, but the discrepan-cies are mostly less than 5 °C (9 °F).

Although the mechanical properties of lead-rich solders are often maligned, they will actu-ally deliver a joint with respectable mechanicalproperties. Chadwick peel-test results with a98Pb-2Sn solder can have peel-fracture initia-tion strengths of more than 35 kN/m (200 lb/in.)when optimallymade [Beal 1984]. Lead-rich sol-ders containing 98% Pb appear to be a goodinitial choice for applications where mechanicalproperties are important, as indicated in Table2.7. At lower lead contents, the deformation char-acteristics are slip bands that are balanced indistribution between the grains and the grainboundaries. At 98% Pb, the grain boundaries ap-pear to be more robust, leading to more heavilylocalized strain effects and superior creep resis-tance. Above this concentration of lead, the sol-der progressively behaves more like a homoge-neous single-phase material.

2.1.6 Indium SoldersIndium-base solders share the common char-

acteristics of low melting point and being ex-tremely soft and ductile. The mechanical prop-erties are mostly a reflection of the fact that atroom-temperature indium solders are operatingat a very high homologous temperature. That is,25 °C (77 °F) is close to their melting point whenexpressed inKelvin. Indium remains ductile evenat cryogenic temperatures. At high homologoustemperatures, the rate of solid-state diffusion inmetals with simple crystal structures is suffi-ciently fast that microstructural changes can oc-cur in timescales that are comparable to changesin the service environment of joints in compo-nents. This is exemplified by the stress-straincurve of a thick indium soldered joint and thecontinuum between stress-strain and creep datagiven in Fig. 2.28 and 2.29, respectively [FreerGoldstein and Morris Jr. 1994; Darveaux andLinga Murty 1993]. This means that recoveryand recrystallization occur as fast as work hard-ening is induced, andmechanical failure of jointsmade using indium-base solders tends to be stressoverload or unidirectional creep.

The ready creep of indium solders implies thatjoints are unlikely to fail on thermal cycling un-less the load-displacement curve is asymmetric.It follows that indium solders are well suited formaking joints between dissimilar materials thatwill be subject to thermal cycling. In fact, creepin joints made with indium solders can usuallyoccur sufficiently fast to ensure that the stress isalways close to zero, with roughly 80% stressrelaxation occurring within seconds of a stepchange in strain. Then, the solder will not workharden. If joints are suitably designed to takeadvantage of these beneficial thermomechanicalcharacteristics, indium solders can provide su-perior life (see Fig. 2.30, also Humpston et al.2001). When indium soldered joints do fail, themechanism is predominantly by classical creeprupture, having its origins in the nucleation

Fig. 2.27 Microsection through a joint of varyingwidthmadeusing the 1.5Ag-92.5Pb-5Sn solder in the form of

a paste to two electronic packaging components in lightweightAl-70Si (Osprey CE7 alloy), plated with nickel and gold. Theblackish region constituting the joint is the solder, which com-pletely fills the gap, which varies in width from less than 20 tomore than 150 μm (0.79 to 5.9 mil).

Table 2.7 Mechanical properties and corrosionresistance of high-lead solders containing tin

Stress-rupturetime, 142 N (32lbf) load, days

Corrosionweight loss

ASTM D 13842 weeks, mg

Leadcontent, %

Peel fracture initiationkN/m lbf/in.

100 17.5 100 122 10.199 19.25 110 8.3 16.198 36.75 210 28.6 9.297 20.12 115 0.55 16.296 21.88 125 0.79 10.4


and coalescence of cavities, arising from the ma-terial redistribution associated with stress relax-ation. In contrast, tin-base solders, when subjectto conditions of low-cycle (low-strain rate) fa-tigue, fail because the plastic deformation in thesolder leads to microstructural changes in theregion of maximum strain. Internal cavities thendevelop, which coalesce to form voids and thengrow into cracks.

Indium-base solders used in optoelectronic andphotonics applications can also fail by a processknown as phase segregation. This occurs whenthere is an extreme and unidirectional electricaland/or thermal gradient sustained across a joint.The result is migration of indium toward onejoint interface and the accumulation of voids atthe other until the joint fails. This mechanism isgenerally only observed in applications such asdie attach of microwave power amplifiers andlaser diodes where, although the absolute powerlevels are modest, the small physical size of theparts results in very high energy flux.

Pure indium is not often used as a solder be-cause the wetting and spreading characteristicsare mediocre as are the mechanical properties ofjoints made using it. One exception stems fromexploitation of the complex oxide that forms onindium.Very-high-purity indium is readily avail-able because this metal is chemically extractedfrom zinc residues as a minor by-product. Pro-vided the indium is of purity better than99.99995%, it will wet and spread over unmet-allized oxide ceramics and glass, in air, withoutflux. The resulting joints do not have the samestrength (5 to 10 MPa, or 725 to 1500 psi) andfracture toughness as conventional solderedjoints, but are nevertheless hermetic and usablein a limited range of applications [Indium Cor-poration of America 1994].

The lowmelting point of indium solders stemsfrom the lowmelting point of indium itself, whichis 157 °C (315 °F). It is interesting to note thatthe cited melting point of indium has increasedby nearly 5 °C (9 °F) over the last 40 years withthe development of improved refining methods;the melting point of the metal being very sen-sitive to low levels of metallic impurities. Albeitlargely as a point of metallurgical curiosity, themelting point of indium is also unusually sen-sitive to pressure; the application of 4000 MPa(580 ksi) will cause the melting point to roughlydouble to 300 °C (570 °F; see Fig. 2.31).

Alloying additions made to indium to formlower-melting-point eutectics confer a numberof practical benefits. Table 2.8 lists some of themore common indium solders and their meltingpoints. The merits of alloying are principally thegeneration of multiphase microstructures, whichimprove the mechanical properties of the solder,and the formation of mixed-composition oxides,which are generally easier to chemically removethan pure indium oxide. Indium oxide formsreadily as a sticky and tenacious film that re-quires specially formulated fluxes to effect its

Fig. 2.28 Shear stress strain curve for a 500 μm thick indium-tin (In-48Sn) joint held at 40 °C (104 °F) ambient

and strained at a rate of 5 � 10–4/s. Adapted from Freer Goldsteinand Morris Jr. [1994]

Fig. 2.29 Continuum between stress-strain and creep datafor indium-tin eutectic solder (In-48Sn) at room

and elevated temperature. Adapted from Darveaux and Murty[1993]

Fig. 2.30 Resistance of a flip-chip daisy chain between sili-con and alumina, versus the number of cycles (N)

of a thermal shock test �25 to –196 °C, 30 s dwell for the soldersindicated. Adapted from Shimizu et al. [1995]


removal. Even so, as can be seen from Fig. 1.14,indium solders are appreciably less fluid thantheir tin-base counterparts and require greatersuperheats to achieve comparable flow.

Indium-lead alloys are a useful class of solderswith readily adjustable melting ranges. Consid-eration of the phase diagram for this alloy sys-tem, which is given in Fig. 2.12, reveals thatindium and lead form an essentially misciblemixture that extends between the melting pointsof the two parent materials, that is, 157 to 328 °C(315 to 622 °F). Thus, by appropriate selectionof the composition it is possible to obtain a sol-der, albeit onewith amelting range andmediocremechanical properties, that has any desired soli-dus temperature between the two limits. Al-though indium-lead alloys solidify by peritecticreaction, the solidification rate of most solderedjoints is usually sufficiently slow to prevent prob-lems associated with microstructural coring, butthis must be considered if a fast heating method,such as laser soldering, is being employed.

Indium-containing solders are often recom-mended for joining to gold-coated componentsbecause gold is less soluble in these alloys than

in lead-tin solders. The restricted solubility ofgold in a joint containing indium is largely as-sociated with the formation of a continuous layerof the intermetallic compound AuIn2 at the sol-der/gold interface, which effectively suppressesfurther reaction between these metals (see Fig.2.2). In the presence of lead, the interfacial layertakes the form of separate grains of the AuIn2compound embedded entirely in primary lead.The lead provides an easy diffusion path betweenthe solder and the gold coating and so permits thereaction to continue, albeit at a modest rate overextended timescales [Yost, Ganyard, and Kar-nowsky 1976; Jacobson and Humpston 1989].

Intermediate melting temperature indium-base solders, such as the In-15Pb-5Ag compo-sition, are significantlymore ductile than lead-tineutectic and can yield joints with correspond-ingly superior thermal fatigue performance [Ed-wards, Nixon, and Lakes 2000]. Although themelting range of this alloy is close to the eutectictemperature of lead-tin solder, its price differ-ential is a factor of about 80, which restricts itsuse to high-added-value and more specialist ap-plications.

2.2 Effect of Metallic Impurities

As might be expected, metallic impuritiespresent in filler metals can be either highly ben-eficial, benign, or totally undesirable, dependingon the solder, the impurity, and the application.The effects of impurities have been most widelystudied for lead-tin solder.

Lead-tin solder is generally prepared to high-purity specifications because the presence of

Fig. 2.31 Melting point of indium as a function of pressure. Adapted from Kennedy and Newton [1963]

Table 2.8 Composition and melting ranges ofsome common indium-base solders

Melting range

Composition, % °C °F

In-50Pb 178–210 352–410In-30Pb 165–172 329–342In-3Ag 141 286In-5Ag-15Pb 142–149 288–300In-40Sn-20Pb 121–130 250–266In-49Sn 120 248In-67Bi 110 230In-34Bi 72 162In-32.5Bi-16.5Sn 60 140


many other elements can have a deleterious ef-fect on joint formation, even in concentrations aslow as 5 ppm [Schmitt-Thomas andBecker 1988;Becker 1991]. With some elements, joint for-mation may be inhibited or even prevented. Thedetrimental characteristics of the elemental ad-ditions and the critical concentrations that pro-duce these adverse effects are summarized inTable 2.9. There was no minimum level of alu-minum that did not adversely affect solderability,except on brass substrates (see below): the low-est concentration investigated was 0.0005%.

A detailed study of the effect of different el-ements on thewetting properties of eutectic lead-tin solder was made by Ackroyd, Mackay, andThwaites [1975]. They assessed the solderabilityon copper, brass, and steel coupons using area-of-spread tests, rotary dip tests, and wetting bal-ance measurements. High-purity Pb-60Sn solderwas deliberately adulterated with low concen-trations (<1%) of aluminum, arsenic, bismuth,cadmium, copper, phosphorus, sulfur, antimony,and zinc and used under an ambient atmospherewith selected active and nonactivated fluxes. Ofthe substrate materials, brass exhibited a con-sistent enhancement in spread area with increas-ing impurity addition, irrespective of the elementin the above selection. By contrast, the solder-ability of copper and steel almost invariably de-teriorated in the presence of subpercentage lev-els of the impurity and then continued to declineor tended to a steady value at higher concentra-tions.

The effect of major ternary additions on thewetting angle of Pb-80Sn solder on copper, brass,and mild steel substrates is given in Table 2.10[Raman 1977]. Bismuth was found to be gen-erally beneficial; other additions such as anti-mony and indium exhibited a dependence on thesubstrate. The ternary additions consistently im-

proved the strength of butt joints to copper andbrass substrates, butmarkedly decreased the jointstrength for the mild steel substrates.

One important area where impurity elementshave a positive role in solders is in suppressing“tin-pest” in lead-tin solders. Pure tin can existin two allotropic forms: white tin, which is thecommon metallic form and gray tin, to whichwhite tin can transform below 13 °C (55 °F).Owing to the 26% increase in volume that ac-companies the phase change from white to graytin, the solid metal disintegrates into a crumblymass having essentially no strength. The trans-formation from white to gray tin is clearly di-sastrous for the mechanical properties of sol-dered joints. “Tin-pest” can affect electricalsystems used in subzero environments, for ex-ample, in high-flying aircraft and also solderedcontainers and conduits of refrigerants.

The transformation does not occur spontane-ously, but is always preceded by an incubationperiod, which may be as much as several yearsif the tin is exceptionally pure. The maximumrate of conversion occurs at about �40 °C (�40°F). Figure 2.32 shows the rate of transformationas a function of time at various temperatures. Theincubation period is much shorter if white tin ismechanically worked at low temperatures or isinoculated with either gray tin or other elementsand compounds having similar diamondlike crys-tal structures, such as silicon and ZnSb [Macin-tosh 1968]. Copper, too, accelerates the process[Rogers and Fydell 1953], whereas lead at higherconcentrations than the eutectic retards it slightly[Williams 1956].

Relatively small additions of certain elementshave been shown to suppress the allotropic trans-formation of tin. The addition of 0.15% anti-mony will prevent it in tin-base solders. Bismuthis even more effective and a 0.05% addition is allthat is necessary [Bornemann 1956]. Existingsolder specifications accommodate sufficientquantities of impurity elements to prevent “tin-pest.” However, in recent years there has been a

Table 2.9 Lowest impurity concentrationsproducing detrimental effects, in terms ofwetting, in Pb-60Sn solderImpurity element Impurity concentration, % Detrimental effect

Aluminium <0.0005 OxidationAntimony 1.0 Reduced spreadingArsenic 0.2 Reduced spreadingBismuth 0.5 OxidationCadmium 0.15 Reduced spreadingCopper 0.29 High melting rangePhosphorus 0.01 DewettingSulfur 0.0015 High melting rangeZinc 0.003 Oxidation

Adapted from Ackroyd Mackay, and Thwaites [1975]

Table 2.10 The effect of major alloyingadditions on the wetting angle of Pb-80Snsolder

Wetting angle, degrees

Solder/addition Copper Brass 0.15% C steel

80Sn 20Pb 26 50 954Sb 38 43 825Bi 22 36 705Cd 30 38 1025In 34 37 98


trend toward high-purity solders in order to im-prove certain characteristics in wave solderingmachines. In some solder specifications, themaximum level of antimony permitted has beenreduced from 0.5 to 0.12%, so that “tin-pest”could occur in soldered assemblies unless otherimpurities, notably bismuth, remain at appropri-ate levels [EP&P 1992].

Metallurgical reactions of solderswith the par-ent materials can also have an effect on the prop-erties and performance of joints made with fillermetals. The most common examples are modi-fication of melting range and physical properties.This subject is dealt with further in section 2.3in this chapter.

Other examples of favorable effects of impu-rities on different solders include:

• Small quantities of the transition metals co-balt and iron (<0.5%) increase the mechani-cal strength of Ag-1.7Cu-93.6Sn solder byabout 25% at room temperature. Cobalt re-fines the Ag3Sn phase, while iron promotesrefinement of primary tin dendrites. The ben-efits are lost at elevated temperature (>150°C, or 300 °F) owing to rapid solid-state dif-fusion [Anderson et al. 2002]. Rare-earth el-ements have a similar effect, as discussed inChapter 5, section 5.8.

• Zinc improves the spread and flow ofAu-2Sisolder, but only at concentrations above 0.5%.Zinc has the additional benefit of softeningthe alloy, reducing the hardness from 320 to128 HV for a 0.75 wt% addition.

• Tin also improves the spreading character-istics and reduces the hardness. However, toachieve these benefits, the concentration oftin ideally needs to be several percent (seesection 2.1.4 in this chapter).

• The addition of germanium to gold-siliconsolder also reduces the alloy hardness withconsequential benefits for joint ductility. Thesoftening of the multicomponent alloy arisesas a result of grain refining of the siliconphase, the grain size being decreased byroughly half for comparable cooling condi-tions.

• Similarly, a 0.5% silver addition providesgrain refinement of bismuth-tin solder.

• Silver, gold, and antimony, at moderate con-centrations (<4%), act as strengthening ele-ments in lead-tin solder.

• The addition of 0.5 to 0.6%Bi to theAu-20Snsolder is said to improve the wettability whenthe alloy is used for joining gold-plated prod-ucts [Rapson and Groenewald 1978].

• Although the presence of aluminum is un-desirable in conventional solders, as it greatlyexacerbates oxide formation, the same is nottrue of all solders. A case in point is the lead-free alloy Sn-8.5Zn-1Ag. Minor additions ofaluminum to this solder, up to a maximum of1 wt%, actually improves wetting of copper.In this instance the increased oxidation isoutweighed by modification of the metal-lurgy such that the formation of copper-zincintermetallic compounds is suppressed in fa-vor of copper-tin phases because of the readyassociation between aluminum and zinc inthe molten state [Cheng and Lin 2002].

This list is by no means exhaustive, and someadditional examples are given in Chapter 3, sec-tion 3.3.8.4. In general, most of these effectshave been discovered by accident rather than bydesign. Manufacturers supply solders of suitablepurity for most applications, but the process de-signer must always be aware of the possibility ofunfavorable composition changes to the solderas a result of alloying with the components beingjoined. This topic is examined in the next section.

2.3 Application ofPhase Diagrams to Soldering

The selection of a solder for a particular ap-plication is often based exclusively on the melt-

Fig. 2.32 Allotropic transformation of white tin into gray tinas a function of time and temperature. Adapted

from Bornemann [1956]


ing point and mechanical properties of the solderand its ability to wet the parent materials. Thesolder is regarded as a uniform layer ofmetal thatsimply bridges the gap between the componentsand binds them together. If only life were thatsimple! In reality, the formation of the desiredmetallic bond between the solder and a compo-nent requires a degree of alloying. The ensuingmetallurgical reactions usually lead to a hetero-geneity of phases in the joint. To further com-plicate matters, kinetic factors tend to accentuatethe development of this nonuniformity. Such in-homogeneities often determine the quality andoverall characteristics of joints, such as their me-chanical properties, the ease and extent of solderspreading, the nature of any fillets formed, andso on.

Metallurgical reactions do not cease once thejoint has been made, but continue to proceed, toa greater or lesser extent, during the service lifeof the assembly. The rate-controlling step forreaction between two solid metals is the diffu-sion of atoms between the reacting phases. Therelative position of the product of the reactionand the reacting phases will be governed largelyby the diffusion coefficients of the participatingmetals. For individual metals, it has been estab-lished empirically that the rate of diffusion R,increases rapidly with absolute temperature, T,following an exponential relationship:

R � A exp�QkT

where k and A are constants; k is the Boltzmannconstant, and A is an experimentally determinedfactor for each combination of reacting phasesthat may vary with concentration. Q is the ac-tivation energy for diffusion which, to a firstapproximation, is proportional to the meltingpoint, Tm, of the particular metal [Birchenall1959].

The rate of reaction will therefore be depen-dent on the homologous temperature defined asthe ratio of T/Tm andwill bemore pronounced forlow-melting-point solders that see service at orclose to normal ambient temperatures. The re-actions produce perceptible metallurgicalchanges in the constitution and microstructure ofsoldered joints [Frear, Jones, and Kinsman 1990,Chapter 2].

For a proper understanding of metallurgicalreactions between solders and parent materials,it is essential to have some grasp of the subjectof alloy constitution. The “constitution” of an

alloy refers to features such as its composition,melting range, range of phase stability, solubilitylimits, and related parameters that can be de-duced from the phase diagram of the system inwhich the alloy appears. In the following sec-tions, some attention is given to highlighting thevalue of phase diagrams and suggesting how thisvaluable source of information might be tapped.

Frequently, the appropriate phase diagram forelucidating specific solder/substrate reactions andjoint microstructures is not available in the lit-erature. Experimental techniques and a method-ology for elucidating gaps in phase diagrams aredescribed in the literature [Humpston and Ja-cobson 1993, Chapter 3]. Note that all of thephase diagrams in this book are defined inweightpercentages of the constituent elements as this ismore appropriate to soldering technology thanthe atomic percentage scale. Also, the relativeproportion of the elements in intermetallic com-pounds, such as Cu3Sn, refer to atomic weights.General equations for converting atomic toweight percent of constituents in alloys, and viceversa, are given in Appendix A3.1.

The fundamentals of alloy phase diagrams arecovered inmanymetallurgical textbooks andwillnot be repeated here. Readers without a back-ground in this field are referred to the publica-tions listed in the Preface under the heading “Al-loy Constitution,” which provide an excellentintroduction to the subject. Here it will suffice tostate that a phase diagram is a representation ofthe thermodynamic stability of phases as a func-tion of composition with respect to particularthermodynamic variables such as temperature or,less commonly, pressure. What is important toremember is that the information given by thediagrams relates to essentially equilibrium con-ditions. The phase diagram tells us about theultimate balance of phases within the joint andthose that are likely to be encountered during theprogression toward equilibrium.A joined assem-bly in which the solder and abutting componentsare different materials are never in true compo-sitional equilibrium, as long as the joint remainsdistinct. In most practical contexts, the compo-sition of a joint will be tending toward equilib-rium over most of its width, and therefore phasediagrams are applicable to an assessment of itsconstitution. However, at the edges of the joint,marked compositional gradients will exist, caus-ing a significant deviation from equilibrium.These will be exacerbated by any temperaturegradients that develop during the process cycleand are manifested as the appearance of different


phases in those regions. Even here, phase dia-grams can assist in the elucidation of the met-allurgical reactions and the resulting phases, asshown in the following section.

Phase diagrams can provide the followingpractical information:

• The melting temperatures of the “virgin” sol-der and of the abutting components

• The probable freezing range of the solderfollowing alloying with the components andhence the remelt temperature of the joint

• Whether the solder remains homogeneous inthe joint after reaction with the componentsand, if not homogeneous, the phases that arelikely to be present, or that may form sub-sequently, with their elemental compositionsand melting temperatures

What a phase diagram is unable to reveal is:

• The rate of reactions that might occur be-tween the solder and the components andtheir variation with time and temperature.This applies both when the solder is moltenand when it is solid during service.

• The spatial distribution and morphology ofphases in the joint, although frequently it ispossible to deduce whether intermetallicphases are likely to form as interfacial layersor will be dispersed throughout the solidifiedsolder. This is explained in section 2.3.1 inthis chapter.

• Thewetting characteristics of a particular sol-der/parent materials combination, of evenperfectly clean surfaces. In practice, wettingis likely to be heavily influenced by the ox-ides, impurities, and residues that are inevi-tably present on component and solder sur-faces, but that are extraneous to the alloyphase diagram.

• Physical properties of joints, in particular themechanical and corrosion characteristics.However, it is often possible to predict thelikely range of certain physical properties bycomparison with other known alloy systems.

The simplest diagrams that are encountered ina joining context are those relating to binaryalloys where, for example, the solder is a singlemetal being used to join components of anothermetal. This situation is represented by the use ofpure tin to solder copper pipes.

Although the available literature on phase dia-grams may appear to be reasonably comprehen-sive, it is worth bearing in mind that reliablediagrams exist for, roughly, only 50% of binary

combinations, 5% of ternary systems, and 0.5%of quaternary mixtures. A compendium of au-thoritative alloy phase diagrams is being pre-pared under the auspices of the International Pro-gramme for Alloy Phase Diagram Data (IPAPD)and the first volumes have already appeared. Thiswork is ongoing, and updates are to be found inthe Bulletin of Alloy Phase Diagrams and Jour-nal of Phase Equilibria. On a periodic basis,these publications include a cumulative indexthat lists all phase diagram evaluations publishedby members of IPAPD. It is worthwhile con-sulting the most recent phase diagram available,because older diagrams can contain significanterrors or omissions.

High-order alloy systems are naturally morecomplex and are less well documented, as notedearlier. However, for a given joining process onlya very limited portion of the phase diagram isrequired, and if one is unavailable it is oftenpossible to experimentally determine the neces-sary data (for example, [Humpston and Jacobson1993, Chapter 3]).

Recently, an exciting breakthrough appears tohave been achieved in a method that is able topredict, with claimed accuracy exceeding 99%,whether compound formation will occur for anybinary, ternary, or quaternary system. This willbe of great assistance in reducing the time andeffort to establish the phase relationships in alloysystems, particularly those of higher order thatare often necessary to encompass soldering andbrazing processes. This work also proved thatmaterials properties are quantitatively containedin the elemental property parameters of the con-stituent elements, so that once additional infor-mation retrieval methods are automated, the se-lection of materials for specific applications willbe greatly facilitated, and it is hoped many newmaterials with exciting property combinationswill be discovered [Villars et al. 2001].

The value of alloy phase diagrams for under-standing and optimizing soldering processes canbest be appreciated by describing a few specificexamples in the following sections.

2.3.1 Examples Drawn fromBinary Alloy Systems

Example 1:ABinary Eutectic CompositionSolder Used with Components of One of theConstituent Metals, with No IntermetallicCompound Formation. A representative ex-ample of this type of reaction is a silver-leadsolder used to join components of pure silver.


The silver-lead phase diagram represented in Fig.2.11 shows that in this binary alloy system thereis minimum melting point, that of the eutecticcomposition (Ag-97.6Pb). Alloys with thisunique composition transform between a liquid(L) and two solid phases (S1 and S2) at a fixedtemperature, 304 °C (579 °F), according to thereaction:

L ↔ S1 � S2

For all other compositions, except those ofthe pure metals, there is a separation betweenthe liquidus and solidus temperatures. On eitherside of the eutectic the liquidus and solidusseparate, with the alloys melting over a range oftemperatures. At temperatures within the melt-ing range, the alloy is partly liquid and partlysolid. On cooling below the solidus tempera-ture, all alloys in this system exist as duplex ofS1 and S2, in direct proportion to the alloycomposition.

It can be seen from the silver-lead phase dia-gram (Fig. 2.11) that silver is soluble in the Ag-97Pb solder when molten. Thus, at the joiningprocess temperature (say, 400 °C, or 750 °F), thesolder will dissolve silver from the componentsuntil the “equilibrium” concentration of silver isattained, determined by the intersection of a linedrawn on the phase diagram at the process tem-perature with the liquidus curve. Thus, at 400 °C(750 °F) the dissolution of silver by theAg-97Pbsolder changes its composition to approximatelyAg-93Pb. That is, the dissolution of silver in-creases the liquidus temperature of the solder inthe joint, but not its solidus temperature, becauseeutectic transformations are isothermal.

Because soldering cycles are short, the solderwill not normally dissolvemore silver. (In theory,if left at temperature for a sufficiently long time,the lead would diffuse through the entire volumeof the silver components so that the assemblyhad a uniform composition determined by thetotal quantities of silver and lead in contact. Thisis actually the basis of diffusion-soldering pro-cesses (referred to in Chapter 5, section 5.9)). Atthe commencement of the cooling stage of theprocess cycle, the molten solder no longer cor-responds to the eutectic composition, but is richin silver and, in consequence, now possesses afreezing (i.e., melting) range. On cooling belowthe liquidus temperature, the excess silver willsolidify first, as indicated by the phase diagram.This precipitation tends to occur preferentially at

the interface between the components and thesolder because this interface is usually slightlycooler than the volume of the molten solder. Pre-cipitation continues until the temperature andcomposition of the remaining liquid reach theeutectic point so that final solidification by themolten solder results in the formation of a smallvolume fraction of finely divided eutectic. Thealloy microstructure will therefore comprise pri-mary dendrites of silver with the interdendriticspaces filled with the duplex eutectic mixture.The primary silver phase will contain a localconcentration gradient as the amount of lead itincorporates varies with temperature. The silverphase is then said to be “cored.”

It is possible to quantitatively follow thechange in composition of the solder as it coolsby application of the lever rule. Referring to theenlarged and schematic portion of the phase dia-gram given in Fig. 2.33, at 350 °C (662 °F) theweight fraction of the solder that is solid, underequilibrium conditions, is:

% solid �XO

XY• 100

The remainder of the solder will be molten,that is:

% liquid �OY

XY• 100

where X is the composition of the liquid phase,Y is the composition of the solid phase, and Ois the composition of the alloy.

Fig. 2.33 Application of the lever rule to an enlarged portionof the silver-lead system diagram (not drawn to

scale)


Alloys of eutectic composition are preferredas solders on account of the following charac-teristics:

• Superior spreading behavior when molten.This feature is an immediate consequence ofthere being no temperature range over whichthe alloys coexist as solid and liquid. Wherea pasty mixture can occur, alloying with thematerials of the components will diminishthe available driving force for spreadingwhilethe partly molten alloy is unable to flow dueto its high viscosity.

• Superior mechanical properties, arisingfrom the interspersed or duplex character ofthe eutectic microstructure and the finegrain size. Grain refinement is the onlymetallurgical process that enhances both thestrength and ductility of a metal. The su-perior mechanical properties of the lead-tineutectic solder, with its duplex microstruc-ture, over the adjacent alloy compositionsis shown in Fig. 2.34.

• Joining process temperatures can be chosento be only slightly above the melting point ofthe alloy, precisely because eutectic compo-sition alloys melt completely at a single tem-perature.

• A reduced risk of disturbing located compo-nents, which can easily occur when the solderappears to be solid but is actually in a pastystate. A rapid liquid-to-solid transformationon cooling, without an intervening pasty

stage, minimizes the chance of such an in-terruption. However, this assumes that alloy-ing of the solder with the component mate-rials does not greatly shift the composition ofthe solder from its eutectic point. A disturbedjoint generally has inferior mechanical prop-erties, and the fillets will acquire a roughsurface with a frosty appearance.

For these reasons, most solders are of eutecticcomposition.

Example 2: ABinary Eutectic CompositionSolder Used with Components of One of theConstituent Metals, with Intermetallic Com-pound Formation.The silver-tin phase diagramis given in Fig. 2.9. The eutectic compositionsolder (Ag-96.5Sn) is widely used to join silver-coated components. One of the main reasons forthis choice is that the erosion of the metallizationon such components is low and highly predict-able. The restricted erosion is a consequence ofthe formation of the Ag3Sn intermetallic com-pound () as a layer that separates the moltensolder from the remaining silver, because it hasa melting point above normal soldering tem-peratures. This interfacial layer restricts furtherreaction to a degree determined by temperatureand the duration of the soldering process, as in-dicated by the data given in Fig. 2.35. Using suchdata, it is possible to determine the minimumthickness of silver that is required for a givenvolume of solder. This feature can be used toadvantage in soldering to nonmetals by ensuringthat the application of a silver metallization ofdefined thickness will prevent erosion through tothe nonmetallic base material in a prescribed sol-dering cycle and thereby avoid catastrophic de-wetting.

While the phase diagram can provide guid-ance about whether a new phase will form, itcannot be used to determine the ultimate distri-bution within the joint as this is greatly influ-enced by a combination of factors.Another pieceof important information that cannot be ascer-tained from equilibrium phase diagrams is therate of growth of phases. The different rates atwhich solders can react with substrate metals isvividly illustrated by comparing the rate of ero-sion of gold by molten indium in Fig. 2.2 withthat by molten tin in Fig. 2.36. Superficially, thephase diagrams of these binary systems are iden-tical.

Although it would appear from Fig. 2.2 thatthe indium-gold reaction is self-limiting, this isonly true in relation to the short timescales of

Fig. 2.34 Tensile strength of cast bars of lead-tin alloys. Op-timal mechanical properties are coincident with

the eutectic composition (Pb-62Sn). Adapted from Inoue, Kuri-hara, and Hachino [1986]


typical soldering operations (seconds). Overlonger periods of time, the reaction will proceedto a significant extent. Even when the solder issolid, the extent of the reaction can be appre-ciable as shown in Fig. 2.37, despite the fact thatreaction rate may be one or two orders of mag-

nitude slower than when the solder is molten, theexact ratio depending on the respective tempera-tures. This point must be borne in mind whenconsidering the stability of a joint over the ser-vice life of the associated assembly. The designlife may be as long as 25 years, which is longer

Fig. 2.35 Erosion of silver by molten tin as a function of reaction time and temperature. Adapted from Evans and Denner [1978],augmented by authors’ own data

Fig. 2.36 Erosion of gold by molten tin as a function of reaction time and temperature


than the duration of the joining cycle by a factorof six orders of magnitude.

Example 3: A Binary Peritectic Solder, Il-lustrating Problems Associated with Using aSolder in This Category. The second commontype of phase transformation is the peritectic re-action where a liquid, L, on cooling, partly so-lidifies to form a solid phase S1 and at the peri-tectic temperature the remaining liquid reactswith S1 to form a new solid phase, S2. This maybe written as:

L � S1 ↔ S2

Indium-lead solders are frequently employedin situations where the standard tin-base solderscannot be used, for example, for soldering atintermediate temperatures between the availableeutectic composition solders. The indium-leadphase diagram, which is given in Fig. 2.12, con-tains two peritectic reactions. These are:

L � Pb � � at 171.6 °C (340.9 °F)

L � � � In at 158.9 °C (318.0 °F)

Alloys exhibiting this type of transformation aregenerally undesirable as solders because, during

a peritectic reaction, it is not possible to maintainequilibrium conditions. This is due to the factthat diffusion rates in solids are about two ordersof magnitude slower than in liquids, so that anonequilibriummicrostructure develops consist-ing of islands of the primary solid phase, S1,completely surrounded by a rim of the secondsolid phase, S2.

A quaternary aluminum alloy microstructureexhibiting a peritectic transformation is shown inFig. 2.38. In such an alloy, liquid that is rich inthe lower-melting-point elements will be re-tained below the peritectic transformation tem-perature. Thus, the melting and freezing range ofthe alloy is widened and the remelt temperaturecannot be reliably predicted. Furthermore, itsmi-crostructure will be grossly inhomogeneous andrelatively coarse, to the detriment of the me-chanical properties of joints madewith this alloy.

In higher-order alloys, a number of other typesof phase transformation can occur and these aregenerally referred to as transition reactions. Themajority of these possess features akin to a peri-tectic transformation, and they tend to be avoidedwhen selecting alloy compositions as solders onthe same grounds.

2.3.2 Examples Drawn fromTernary Alloy Systems

It is rare for a soldered joint to be limited to acombination of just two elements, forming a bi-naryalloy system.Usually the solder is analloyofat least two metals, while engineering alloy sub-strates are frequently multicomponent. Com-

Fig. 2.37 Continued growth of gold-indium intermetallicphases at the interface between a gold metalliza-

tion and indium-lead solder at elevated temperature but belowthe solidus temperature. Adapted from Frear, Jones, and Kinsman[1990]

Fig. 2.38 An alloy microstructure characteristic of a peri-tectic transformation. The alloy contains four con-

stituents: aluminum, copper, nickel, and silicon. The primaryphase is totally surrounded by a rim of a second phase as a resultof the peritectic reaction failing to maintain equilibrium condi-tions during solidification.


monly, intermetallic compounds form betweenthe constituents. The volume, distribution, andmorphology of these intermetallic phases in ajoint can have a pronounced effect onmechanicalproperties, in particular. From the relevant phasediagram, it is possible to predict whether the in-termetallic compound will tend to form as a con-tinuous interfacial layer against the parent mate-rials or is more likely to be dispersed throughoutthe joint. If the intermetallic compound has poormechanical properties, then a dispersion of it ispreferred because this not only avoids the sourceof weakness represented by the interfacial inter-metallicphase,butactuallyworkstoadvantagebystrengthening the solder. Examples of these dif-ferent situations are described below.

Some care needs to be taken when referring tothe intermetallic phases that form in a ternarysystem. An example is provided by binary cop-per-tin solder wetted on nickel substrates to forma mixture of copper-tin and nickel-tin interme-tallic phases at the interface. The apparent con-fusion in the literature over the composition ofthese phases can be explained by reference to theCu-Ni-Sn phase diagram that reveals that thebinary compounds Cu6Sn5, Ni3Sn2, and Ni3Sn4not only have very extensive ternary solubility(up to 30%), but there exists a continuous solidsolution between Cu3Sn and Ni3Sn. Hence, thereis considerable scope for composition variationwithin the stable phase fields, and the convenientlabels adopted from binary alloy practice do notadequately describe the real situation [Lin, Chen,and Wang 2002].

A ternary system is most usually representedby an equilateral triangle, with each of the ver-tices corresponding to the three elements. A gridis normally drawn on the triangle to provide alinear scale of composition. Temperatures arethen represented by a series of isotherms, so thatthe position of the liquidus on the diagram ismapped as a topographical surface viewed inplan. Phase stability as a function of temperatureis commonly represented by a diagram resem-bling a binary alloy phase diagram, where eitherone of the constituents or the ratio of two con-stituents is held fixed. A single diagram cannotbe used to track the solidification sequence be-cause the ensuing composition changes can ex-tend outside the plane of the diagram. For a simi-lar reason, the lever rule cannot be applied to thisrepresentation in order to calculate the propor-tions of phases that exist in equilibrium. How-ever, the lever rule can be used in conjunctionwith a series of isothermal sections.

Example 4: Interfacial Compound Forma-tion between a Eutectic Solder and the Com-ponent Metals. As an example of interfacialcompound formation, it is well known that lead-tin solders, when used with copper components,form several copper-tin intermetallic com-pounds, predominantly at the copper/solder in-terface. Despite the many billions of solderedjoints that have been made to copper using lead-tin solder since antiquity, a complete and cred-ible ternary phase diagram for this alloy systemonly become available in 1994. The binary cop-per-tin phase diagram, a liquidus projection, andthe key isothermal sections and isopleths rel-evant to soldering to copper are reproduced inFig. 2.39 to 2.42 [Yost, Hosking, and Frear 1993].

The features to note are:

• Copper is soluble in molten lead-tin solders.• The dissolution of copper by the solder mar-

ginally depresses the melting point of thelead-tin eutectic alloy.

• Liquid lead-tin eutectic solder that has dis-solved a small quantity of copper will so-lidify via a ternary eutectic at 182 °C (360°F). A reduction in melting point of a solderon dissolution of substrate constituents is akey driving force for spreading. The ternaryeutectic point is actually at 61.75Sn-38.05Pb-0.2Cu (wt%).

• The third constituent in the majority of re-actions between lead-tin solder and copper,in addition to the tin and lead phases, is theintermetallic compound Cu6Sn5. It is only forvery lead-rich solder compositions or whenother molten solder compositions have hadthe opportunity to dissolve appreciable con-centrations of copper that the intermetalliccompound Cu3Sn is present as a primaryphase.

The rate of dissolution of copper in moltenPb-60Sn solder is initially rapid, as is the erosionof most other engineering metals and metalliza-tions by this alloy, as can be seen from Fig. 2.43.Consequently, despite the short process cycletimes normally associated with soldering prac-tice, the molten solder will dissolve sufficientcopper to reach the saturation concentration, dic-tated by the process temperature and the com-position of the liquidus surface in the Cu-Pb-Snphase diagram.

On cooling this ternary solder, it would beexpected from consideration of the phase dia-gram that copper-tin intermetallic phases willprecipitate and become distributed within the


bulk of joint. If the peak process temperature iskept below 375 °C (705 °F), the phase diagramshows that the primary phase that precipitates on

cooling will be Cu6Sn5. This has been confirmedby experiment.At normal solder process coolingrates, the precipitates are typically of the order of

Fig. 2.39 Copper-tin phase diagram. � � Cu6Sn5; � Cu3Sn

Fig. 2.40 Liquidus surface of the Cu-Pb-Sn system. Adapted from Yost, Hosking, and Frear [1993]


Fig. 2.41 Expansion of the solder-rich, copper-poor, region of the Cu-Pb-Sn system. Regions are identified by the primary or firstphase that solidifies on cooling in that region. Adapted from Yost, Hosking, and Frear [1993]

Fig. 2.42 Isothermal section of Cu-Pb-Sn system at 150 °C. Adapted from Yost, Hosking, and Frear [1993]


5 to 20 nm (0.2 to 0.8 μin.) in size and thereforecan only be resolved using high-resolution im-aging techniques such as transmission electronmicroscopy [Felton et al. 1991]. However, if thepeak process temperature exceeds 375 °C (705°F) and remains above this level long enough formore than 1.6% of copper to dissolve, then thefirst phase to precipitate will be Cu3Sn.

It is also apparent from the ternary phase dia-gram that copper-tin intermetallic compoundswill form at the interface between the solid cop-per and molten solder. Owing also to the pre-vailing concentration gradient, the copper-richCu3Sn phase forms adjacent to the copper com-ponent with the more tin–rich Cu6Sn5 phase be-tween it and the solder, as shown in Fig. 2.44.

The rate of formation of interfacial copper-tinintermetallic phases is governed by the rate ofinterdiffusion of copper and tin through them and

is well characterized. A distillation of the datathat are available in the literature is presented inFig. 2.45. The Cu3Sn phase layer does not growto any significant thickness during typical in-dustrial soldering processes because the inwarddiffusion of tin through the Cu6Sn5 phase doesnot normally reach sufficient concentration[Dirnfeld and Ramon 1990]. The formation ofcontinuous layers of intermetallic compounds be-tween the copper and the molten solder greatlyrestricts the further dissolution of copper in thesolder [Felton et al. 1991].

Copper-tin intermetallic compounds do notpossess particularly desirable mechanical orphysical properties in bulk form (see Table2.11). These phases will continue to increase inthickness during elevated-temperature serviceof the component, albeit much more slowlythan when the solder is molten as can be seen

Fig. 2.43 Rate of dissolution of a range of engineering parentmetals and metallizations in lead-tin solder as a

function of temperature. Adapted from Klein Wassink [1989]

Fig. 2.44 Phases formed by reaction between a lead-tin sol-der and a copper substrate, following extended

heat treatment

Fig. 2.45 Growth of copper-tin intermetallic compounds ona copper substrate wetted by lead-tin eutectic sol-

der as a function of reaction time and temperature

Table 2.11 Mechanical and physical propertiesmeasured for three common intermetalliccompounds, on bulk specimens

The phases are generally hard with low fracture toughness.However, their thermal expansivities lie between the substrateand solder, which probably plays a role in decreasing the stressconcentration in joints.

Property Cu6Sn5 Cu3Sn Ni3Sn4

Hardness, HV 378 343 365Toughness, MPa • m1/2 1.4 1.7 1.2Electrical resistivity, μ • cm 17.5 8.9 28.5Thermal conductivity, W/m • K 34 70 20Thermal expansivity, 10�6/K 16.3 19 13.7

Adapted from Fields, Low, and Lucey [1991]


from Fig. 2.46. The presence of thick layers ofcopper–tin intermetallic phases is widely re-ported as being detrimental to the mechanicalproperties of the joints, and this has promptedmany studies of the effect of their presence, forexample Dirnfeld and Ramon [1990]. Figure2.47 shows the relationship between the ten-sile strength of joints and the thickness of thecopper-tin intermetallic layer as determined inthese studies, and Fig. 2.48 shows the relation-ship between intermetallic thickness and creeprupture life [Hongyuan, Ying, and Yiyu 1994].Notwithstanding this general concern, there are

very few reports of joint failures that can defi-nitely be pinpointed to the copper-tin interme-tallic phases. This unclear picture is under-standable because tin forms hard intermetallicphases, of comparable thickness, by solid-statediffusion at elevated temperature with virtuallyall the metals commonly used in engineering[Kay and Mackay 1976]. Some relevant dataare reproduced in Fig. 2.49. Not surprisinglythen, joints made to copper testpieces usinglead-tin solder are found not to be signifi-cantly weaker than joints to other commonsubstrate materials—in terms of their strength,ductility, and fatigue resistance—when formedusing conventional process cycles such that theintermetallic layers remain relatively thin.

Not all the binary lead-tin solders form twointermetallic phases on reaction with copper. If

Fig. 2.46 Growth of copper-tin intermetallic compounds ona copper substrate in contact with lead-tin solder

for 100 s at different process temperatures

Fig. 2.47 Effect of thickness of copper-tin intermetallic com-pounds in soldered joints on tensile strength at

room temperature

Fig. 2.48 Relationship between thickness of copper-tin in-termetallic and creep rupture life of joints made to

copper components using lead-tin eutectic solder

Fig. 2.49 Growth of tin-base intermetallic phases by solid-state diffusion at 170 °C (340 °F) on various sub-

strate materials. Adapted from Kay and Mackay [1976]


the tin level is less than 25%, the phase diagramin Fig. 2.41 indicates that only the Cu3Sn inter-metallic will precipitate at the solder/copper in-terface under near-equilibrium conditions [Gri-vas et al. 1986].

Example 5: DistributedCompound Forma-tion between a Eutectic Solder and the Com-ponent Metals. Another industrially importantalloy system comprises gold, lead, and tin, onaccount of the widespread use of gold as a sol-derable metallization for fluxless joining in theelectronics industry.

The liquidus surface of Au-Pb-Sn ternaryphase diagram is given in Fig. 2.50, and a sectionfrom the eutectic Pb-62Sn composition towardgold is shown in Fig. 2.51. The principal featuresof these diagrams are closely akin to those of theCu-Pb-Sn system discussed previously, albeitwith the difference that in the gold-base systemat least three embrittling gold-tin intermetalliccompounds are able to form directly from theliquid. Therefore, it might be expected that, whenlead-tin eutectic solder is used to join solid goldcomponents, the resulting microstructure ofjoints will be similar. However, there are twoimportant differences.

Gold metallizations are seldom more than afew microns (sub-mils) thick, largely on accountof the cost of this metal, so that there is a limitto the concentration of gold that will accumulatein the molten solder. The precise value is deter-mined by both the volume of solder alloy and thethickness of the metallization, but is usually wellbelow the saturation concentration. Also, due tothe high rate of dissolution of gold in moltenlead-tin solders, even thickmetallizations will becompletely dissolved during the heating cycle.

On cooling a Au-Pb-Sn ternary alloy formedby reaction of lead-tin eutectic solder with a goldmetallization, the phases that precipitate are in-dicated by the phase diagram. Because of theirunfavorable mechanical properties, it is gold-tinphases that are of interest. If the gold concen-tration is below 1%, then the gold remains dis-solved as a solid solution in the lead and tinphases. Between 1 and 5%Au, solidification ter-minates by the ternary eutectic reaction so thatthe AuSn4 phase present will be well-dispersedthroughout the volume of the solder and finelydivided. Between 5 and 8% Au, AuSn4 precipi-tates as a secondary phase and is consequentlypresent as large conglomerates in addition to be-

Fig. 2.50 Liquidus projection of the Au-Pb-Sn ternary system. The first phase to form on solidification is labeled for each phase field.The 4% Au isoconcentration line is marked on the figure as is the tie line between lead-tin eutectic solder and pure gold.

Aconcentration of 4% gold in the eutectic is taken as a safe limit becauseAuSn4 cannot form as a primary phase. Adapted fromHumpstonand Davies [1984, 1985]; Humpston and Evans [1987]


ing a constituent of the ternary eutectic. Above8% Au, AuSn4 is the primary phase formed onsolidification and becomes the dominant con-stituent of the solder microstructure. At evenhigher concentrations of gold (above about 13%Au), the solder constitution changes abruptlywith gold-tin phases being the major compo-nents of primary, secondary, and tertiary solidi-fication.

When gold-tin phases form as the primary ormajor phase, they embrittle the soldered joints onaccount of their intrinsically low fracture tough-ness and the weak interface between them andthe lead-phase in lead-tin eutectic [Harding andPressly 1963].Amaximumgold concentration of4% is usually taken as a safe working limit inindustry. As can be seen from Fig. 2.50, high-lead and high-tin content in lead-tin solders canaccommodate slightly higher levels of gold be-fore gold-tin intermetallic compounds becomedominant (i.e., constitute the primary phase) andcause joint embrittlement.

Likewise, from the appropriate sections of thephase diagrams of other tin-base solder alloys incombination with gold, it is possible to derivesafe working limits for them in a similar manner.

Low concentrations of the AuSn4 phase actuallyenhance the mechanical properties of many tin-containing solders, including lead-tin [Wild1968]. This strengthening is illustrated in Fig.2.52 for the Ag-96Sn solder. From a concentra-tion of about 3.5 to 8% Au, the solder micro-structure is characterized by a fine dispersion ofthe AuSn4 phase, present as a secondary phase.However, when the level of gold in the solderrises toward 10%, there is a sudden change inproperties and ingots of the solder become com-pletely unworkable. This change corresponds tothe appearance of AuSn4 as the primary phase;that is, AuSn4 is the first solid phase to form onsolidification and consequently adopts a massiveform.Above 19%Au the primary phase isAuSn2,and at even higher gold concentrations (greaterthan 38%) it is AuSn. The safe working limit forthis alloy system is usually taken to be 8% Au.The ternary Ag-Au-Sn phase diagram and rel-evant pseudobinary section between eutectic sil-ver-tin solder (Ag-96.5Sn) and gold are shown inFig. 2.53 and 2.54. It is noteworthy that althoughtheAg-96Sn alloy containsmore tin than lead-tineutectic solder, it is tolerant to approximatelytwice the volume fraction of gold before the al-

Fig. 2.51 Vertical section through the Au-Pb-Sn ternary system between eutectic lead-tin (Pb-62Sn) composition and gold, with the8% Au concentration marked by a dashed line. The plan view of this section is marked by a dashed line on the liquidus

projection of the Au-Pb-Sn ternary system shown in Fig. 2.50.


loy is embrittled by AuSn4. This is thought to bedue to a better match between the atomic latticesof Ag3Sn and AuSn4 than between lead andAuSn4. The critical levels of gold that give risetoAuSn4 as the primary phase are listed in Table4.3 for a selection of solders. Soldering pro-cesses that involve gold metallizations are dis-cussed in Chapter 4, section 4.1.3 and Chapter 3,section 3.3.8.1.

Themechanical properties of joints containingintermetallic phases can be inferred from a phasediagram according to whether they are com-pounds of exact stoichiometric composition (i.e.,they are in integral atomic ratios of their constitu-ents) or exist over a range of compositions. Exactstoichiometric compounds tend to formwhenoneof the two elements is strongly metallic in char-acter and the other significantly less so, in termsof the density of free electrons that bind the atomsof the metal together. AuSn4, FeSn2, and Cu6Sn5are typical examples of this type of compound.These compounds tend to be hard and brittle.Moreover, because their crystal structures are fre-quently of low symmetry—that is, they deviatefrom simple cubic or hexagonal structures—theinterfaces of these compounds with other phasestend to be weak. These characteristics are trans-ferred to the joint unless the compounds form asafinedispersionwithinthesolder.Therefore, theiroccurrence should be minimized or, even better,avoided wherever possible.

Compounds that are stableovera rangeofcom-positions tend to be ductile andhave crystal struc-tures exhibiting high symmetry, as do most el-emental metals. Therefore, they tend to have abenign effect on joint properties. An example ofsuch a compound is Ag3Sn, which is stable over

Fig. 2.52 Aselection ofmechanical properties of Ag-96Sn asa function of gold addition

Fig. 2.53 Liquidus surface of the Ag-Au-Sn system. The firstphase to form on solidification is labeled for each

phase field.

Fig. 2.54 Vertical section through the Ag-Au-Sn ternary sys-tem between eutectic silver-tin solder (Ag-96.5Sn)

and gold. The plan view of this section is marked by the dashedline in Fig. 2.53.


the composition range from13 to20%Agat roomtemperature and has a hexagonal close-packedcrystal structure. This compound forms as an in-terfacial layerwhensilver-tinsolder isused to joinsilver-coated components in a manner analogousto the reaction between lead–tin eutectic and cop-per.An example is shown in Fig. 2.55, and the as-sociatedbinaryphasediagraminFig.2.9.Therateof growth of theAg3Sn layer decreases exponen-tiallywith time at a fixed temperature as shown inFig. 2.56. The growth of this intermetallic layercorrelateswith themeasured rate of erosionof sil-ver by tin (Fig. 2.35). The growth of the Ag3Snlayer progressively reduces the dissolution of sil-ver. For this reason, the silver-tin eutectic solder,used in conjunctionwith silver-coated substrates,

formsthebasisofasolderingprocess that ishighlytolerant to the processing time and temperature,inasmuch as the risk of totally dissolving a fairlythin (�10 μm, or 0.4mil) silvermetallization canbe minimized.

2.3.3 Complexities Presented byHigher-Order and NonmetallicSystems

More often than not, higher-order alloy sys-tems are encountered in joining, because both thesolder and the parent materials usually each havea minimum of two constituents. Combinationsinvolving five or even larger numbers of ele-ments are not uncommon.

The definition of the plethora of phases thatcan exist in these higher-order systems repre-sents a daunting task. In order to make the prob-lemmore tractable, a reductive approach is oftenemployed. This method usually involves parti-tioning the multicomponent system into a seriesof quasi-binary or quasi-ternary alloy systems,each containing different but fixed proportions ofthe other components and ascertaining these sec-tions of the relevant phase diagrams, in turn.Much care should be exercised in extractingquantitative information from these partial phasediagrams because the tie lines, triangles, quad-rilaterals, and so forth that are used with the leverrule to determine the relative proportions of

Fig. 2.55 Micrograph revealing a continuous interfacial layerof the intermetallic compound Ag3Sn formed on

reaction between tin-silver solder and silver. Magnification: 20�

Fig. 2.56 Thickness of the Ag3Sn intermetallic layer formed by reaction between Ag-96Sn solder and silver as a function of reactiontime and temperature. After Evans and Denner [1978], with authors’ own data


phases present often do not lie in the plane of theselected sections.

2.4 Depressing the MeltingPoint of Solders by EutecticAlloying

The number of commercially available sol-ders is finite, and it is not uncommon to have anapplication where it would be desirable either toextend the lower temperature limit on the use ofa particular family of solders or to identify fillercompositions that melt within a specified range.An obvious case in point is the search for re-placements for lead-tin solder that do not containlead, but that can be used at comparable pro-cessing temperatures.

For the reasons elaborated in section 2.3 in thischapter, eutectic alloys possess several key char-acteristics that make them a natural choice forfillers, namely superior spreading behavior whenmolten, with complete melting occurring at asingle temperature that is lower than, and usuallywell below, that of either constituent metal. Thisproperty of instantaneous melting enables join-ing operations to be carried out at temperaturesonly slightly in excess of the solidification tem-perature, and molten eutectic alloys generallypossess a high degree of fluidity. Eutectic alloysalso possess favorable mechanical propertiesarising from their well-distributed duplex orhigher-order microstructure.

As shown for the particular example of theBi-Pb-Sn solder in Fig. 2.16, a eutectic alloyformed of three constituent metals (that is, a ter-nary eutectic) always has a melting point that islower than the three constituent binary eutectics.What then happens to the melting point if furtherconstituents are added? In particular, can it belowered at will and, if not, how might one de-termine the limits?

Answers to these questions have been pro-vided by analyzing the pattern of behavior ob-served in two simple examples, namely the de-velopment of cadmium solders and gallium alloysystems intended as substitutes for mercury.From their general features it is possible to de-vise certain empirical ground rules of generalapplicability that can be applied to other cases,such as the quest for lead-free solders with asimilar melting point to lead-tin eutectic (183°C, or 361 °F). Other examples, pertaining to

low-melting-point aluminum brazing alloys, aregiven in the planned companion volume Prin-ciples of Brazing.

2.4.1 Liquid Alloys Based on Gallium

There are requirements for metals that are liq-uid at normal ground temperatures. Mercury isthe only metallic element that is molten down tosubzero temperatures (freezing at �39 °C, or�38 °F). It has therefore enjoyed a near mo-nopoly in thermometry for more than two cen-turies as well as in other applications employingliquid metals such as mirrors, high-currentswitches, and slip rings.With the growing aware-ness of health and safety issues in recent years,there has been a move away from the use ofmercury in these applications, mostly to func-tional substitutes, rather than toward finding di-rect replacements. Some direct mercury-free al-loy substitutes have been developed, based ongallium eutectics, gallium being the second-lowest-melting-point metal next to mercury.Thus, for example, Ga-In-Sn, melting at 12 °C(54 °F), and Ga-In-Sn-Zn, melting at 9 °C (48°F), have been used in high-current switches[Walkden, Kowalczyk, and Cooke 1978]. Theprogressive reduction of themelting point of gal-lium alloys, as the number of constituents is in-creased, is shown in Tables 2.12 and 2.13.

2.4.2 Cadmium-Base Solders

Cadmium forms a series of binary eutectic al-loys with other common solder ingredients thatmelt at temperatures down to 123 °C (253 °F), asshown in Table 2.14. The alloying elements se-lected in this example are indium, tin, and zinc,which all melt at relatively low temperatures andform binary eutectics with one another, as well aswith cadmium. In Table 2.14, it can be seen thatternary alloys of cadmium with these elementsspanameltingrangefrom163to93°C(325to199°F) and on down to 90 °C (194 °F) for the qua-ternary eutectic in the Cd-In-Sn-Zn system.

2.4.3 General Features

The melting point behavior follows the fol-lowing pattern for the three alloy systems con-sidered here:

• The melting point drops monotonically withthe addition of each successive constituent.


• The size of the melting-point depression isdependent on the specific alloying addi-tions. Some elements are more effectivethan others in lowering the melting point ofthe alloy.

• The incrementalmelting-point depression ac-companying the addition of each new alloy-

ing element that enters into the eutectic re-action becomes progressively smaller, so thatthemelting point tends to an asymptoticmini-mum.

These features are consistent with elementarythermodynamic and statistical models, as ex-

Table 2.12 Temperature of eutectiferous phase transformations in which one of the participatingphases is gallium

Order Alloy systemMelting point,

°CTemperature depression,

��T

1 Ga 30 02 Ga-Ag

Ga-InGa-SnGa-Zn

25162125

51494

3 Ga-Ag-InGa-Ag-SnGa-Ag-ZnGa-In-SnGa-In-ZnGa-Sn-Zn

1421(a)20(a)1213(a)17

169

10201713

4 Ga-Ag-In-SnGa-Ag-In-ZnGa-In-Sn-Zn

999

212121

5 Ga-Ag-In-Sn-Zn 6(a) 24

(a) Authors’ own measurements

Table 2.13 Effect of addition of a fourth element to the temperature of the Ga-In-Sn eutectic



��T

3 Ga-In-Sn 12 04 Ga-In-Sn�Ag

Ga-In-Sn�AlGa-In-Sn�BiGa-In-Sn�CdGa-In-Sn�CuGa-In-Sn�PbGa-In-Sn�Zn

99(a)

10(a)10(a)10(a)10(a)9

3322223


Table 2.14 Temperature of eutectiferous phase transformations in the Cd-In-Sn-Zn quaternarysystem



��T

1 Cd 321 02 Cd-In

Cd-SnCd-ZnIn-SnIn-ZnSn-Zn

123177266117142199

19814495

204179122

3 Cd-In-SnCd-In-ZnCd-Sn-ZnIn-Sn-Zn

93116163108

228205158213

4 Cd-In-Sn-Zn 90 231


plained inAppendixA2.2 to this chapter, and aregeneric to eutectic alloying.

From a practical point of view, the implica-tions are clear. When seeking a reduction in themelting point of a pure metal or of an existingeutectic alloy, for use in soldering or other ap-plications, there is a trade-off between keepingthe number of constituents low and judiciouslychosen to optimize the melting-point depressionand increasing the number of components, whichis likely to produce only a relatively small furtherreduction in the melting point. Multicomponentsolder are beset by disadvantages, chief amongwhich is a considerably more complex phasediagram, if it is indeed known, and often, a harderand less workable alloy.

On account of the cumulative complicationsassociated with increasing the number of con-stituents in the filler, it is advantageous, wherepossible, to limit the choice to alloying additionsthat are most effective at depressing the meltingpoint. Generally, these are elements that havelow melting points and very limited solid solu-bility in the host metal.

The implications from the foregoing discus-sion are not encouragingwith regard to the searchfor “drop-in” replacements for eutectic lead-tinsolders, as explained below. Further details onlead-free solders are given in Chapter 5, section5.1.

2.4.4 Implications forLead-Free Solders

The wealth of published studies of alternativesolders has generally arrived at a consensus viewthat alloys that provide the closest alternatives tolead-tin eutectic solder, which can meet the ac-ceptance criteria for printed circuit board (PCB)manufacture are tin-rich with the minor constitu-ents being one or more of the following: anti-mony, bismuth, copper, silver, and zinc [Hamp-shire 1993]. Ideally, the replacement soldershould have a melting point close to that of lead-tin, lying in the range between 150 and 200 °C(302 and 392 °F), and be of eutectic composition.Here, the objective is not tominimize themeltingtemperature, but to restrict it to within a pre-scribed temperature window.

Because none of the known binary eutecticalloys of tin satisfy the above criteria, a studywas made of higher-order systems to establishwhether these might contain suitable eutecticcompositions. The results are given inTable 2.15.From the experimental data, conclusions can be

drawn about the solidification of tin-rich alloysand are described below.

Those tin-rich multicomponent alloys that so-lidify eutectiferously do so at a temperature thatis either above 200 °C (392 °F) or significantlybelow 150 °C (302 °F). The available range ofsuitable alloying additions will either only lower

Table 2.15 Alloying sequences that show thatdrop-in replacements for lead-tin solders, basedon tin, are unobtainableA suitable replacement for lead-tin eutectic solder would need tomeet the following criteria:

• Contain only inexpensive and nonvolatile and nonhazardousconstituents

• Be a eutectic with a melting point between 150 and 200 °C(302 and 392 °F)

1. Candidate constituent binary systems and their eutectictemperatures are:

Ag-Sn (221 °C)Bi-Sn (139 °C)Cu-Sn (227 °C)In-Sn (120 °C)Sn-Zn (199 °C)

Another binary alloy that might provide the basis for amulticomponent solder is the noneutectic, namely: Sn-6Sb,melting range: �232–250 °C (450–482 °F).

2. Candidate ternary systems for replacing eutectic lead-tinsolder:

Ag-Cu-Sn, ternary eutectic at 217 °C [Miller, Anderson,and Smith, 1994]

Ag-Sb-Sn(a), ternary transition reaction terminating onthe Ag-Sn binary eutectic (221 °C)

Ag-Sn-Zn(a), ternary transition reaction terminating onthe Sn-Zn binary eutectic (199 °C)

Cu-Sb-Sn, ternary transition reaction terminating on theCu-Sn binary eutectic (227 °C)

Cu-Sn-Zn(a), ternary transition reaction terminating onthe Sn-Zn binary eutectic (199 °C)

All other combinations result in ternary eutectic alloys meltingbelow 150 °C, thus:

Ag-In-Sn(a) (113 °C)Bi-Sn-Zn(a) (130 °C)In-Sn-Zn(a) (108 °C)

or, in the case of the Ag-Bi-Sn and Bi-Sb-Sn alloys, terminateon the Bi-Sn binary eutectic composition (139 °C) [seeKattner and Boettinger, 1994 for the Ag-Bi-Sn system andOhtani and Ishida 1994 and Ghosh, Loomans, and Fine, 1994for the Bi-Sb-Sn system]

3. Quaternary alloys based on the ternary alloys that melt above150 °C

Freezing of these alloys concludes at the binary eutecticmelting point. Thus, for example: Ag-Cu-Sn-Zn(a) undergoesa ternary transition reaction terminating on the Sn-Zn binaryeutectic.



the melting point slightly below that of tin, melt-ing point 232 °C (450 °F), as in the case of the3.5Ag-0.9Cu-95.6Sn ternary eutectic, meltingpoint 217 °C (423 °F) [Loomans and Fine 1999],or will effect a radical reduction in melting point,for example, alloying with indium and/or bis-muth [Yoon et al. 1999; Hassam, Dichi, and Leg-endre 1998]. A eutectic alloy of silver-tin withgold, of composition 3.6Ag-3.6Au-92.8Sn, has amelting point of 206 °C (403 °F), but the costpremium associated with the gold content wouldbe unacceptable for the manufacturing industry.The Sn-9Zn eutectic alloy, melting at 199 °C(390 °F), may be judged unsuitable for mostsoldering applications because of its propensityto corrode with the generation of a conductivewhite zinc chloride surface film when subjectedto a chlorine-containing atmosphere.

Solidification of all other tin-rich multicom-ponent alloys occurs via one or more transitionreactions in a sequence that terminates on one ofthe tin binary eutectics, that is, at a temperatureclose to that of tin and above 200 °C (392 °F).An example is provided by the Bi-Sb-Sn system[Ohtani and Ishida 1994; Ghosh, Loomans, andFine 1994].

It has been shown that some of the multicom-ponent alloys that are being put forward as “drop-in” replacements for lead-tin are handicapped bythe presence of a low-melting-point eutectic frac-tion that can severely compromise the reliabilityof soldered joints if segregation occurs towardthe low-melting-point alloy composition, eithercontinuously at joint interfaces or along grainboundaries [Vincent and Humpston 1994]. Oneof these alloys is the Ag-20In-77Sn alloy beingoffered as a lead-free solder, which substantiallymelts at 177 °C (351 °F). However, a significantfraction of this alloy melts at a much lower tem-perature of 113 °C (239 °F) [Korhonen andKivilahti 1998], associated with a low-melting-point Ag-In-Sn composition, close to the In-48Sn binary eutectic [Artaki, Jackson, and Vi-anco 1994]. Likewise, the In-86Sn-9Zn solderlargely melts between 188 and 198 °C (370 and388 °F), but it contains a proportion of the In-46Sn-2Zn ternary eutecticmelting at 108 °C (226°F) [McCormack, Jin, and Chen 1994].

It may reasonably be concluded that there areno higher-order tin-base eutectics, without gold,that solidify in the target temperature range.Hence, themulticomponent alloys offer nomajoradvantages over the binary eutectic alloys of tin,with regard to melting point. It follows that lead-free solders that are developed in the future are

likely to be based on the silver-tin binary alloysystem, which has a eutectic temperature of 221°C (430 °F). Economic and functional benefitsare obtained by substituting some of the silverfor copper and exploiting the Ag-Cu-Sn ternaryalloy system, which has a eutectic temperature at217 °C (423 °F), as mentioned above [Miller,Anderson, and Smith 1994; Vincent and Hump-ston 1994].More information on the rationale forthis choice is given in Chapter 5, section 5.1.

Appendix A2.1:Conversion betweenWeight and AtomicFraction of Constituents ofAlloys

In an alloy containing N constituents, con-version from weight to atomic fraction of con-stituent, n, may be made using:

at.% n �Pn /An

�i�1

N

Pi' /Ai

� 100

where P is the weight percentage of the con-stituent denoted by the subscript, A is the atomicweight of the constituent denoted by the sub-script, subscript n refers to constituent n, andsubscript i refers to each constituent in turn.

Similarly, in an alloy containing N constitu-ents, conversion from atomic to weight fractionof constituent, n, may be made using:

wt% n �Pn

' An

�i�1

N

Pi Ai

� 100

where P' is the atomic percentage of the con-stituent denoted by the subscript, A is the atomicweight of the constituent denoted by the sub-script, subscript n refers to constituent n, andsubscript i refers to each constituent in turn.


Appendix A2.2:Theoretical Modeling ofEutectic Alloying

The laws of thermodynamics account for alowering of the melting point, when a substance,B, is added to a pure solvent, A, by an amountgiven byRaoult’s law in the form of theClausius-Clapeyron equation. At the liquidus line repre-senting the equilibrium between a solid solutionof B in A and a liquid solution of B in A, theClausius-Clapeyron equation takes the form:

�Hm(1)

R ( 1

T (2)m

�1

T (1)m

) � �ln ( x1

x 1* )

where x1 is the mole fraction of component A ofthe liquidus composition at temperature T

m(2), and

xi* is this mole fraction at the limit of solid solu-

bility at the same temperature. Tm(1) is the melting

temperature of pure A, and � Hm(1) is its latent

heat of fusion. R is the universal gas constant.The latent heat �Hm (or enthalpy of fusion) of

a metal is proportional to its melting point. Thisis because entropies of fusion (�Sm) have similarvalues for all metals (at roughly 10 J/K • mol),and to a first approximation, in the absence ofany phase changes:

�Hm � Tm �Sm � 0 (Eq A2.1)

that is:

�Hm � Tm

Equation A2.1 therefore implies that the attain-able melting-point depression is determined bythe melting point of the addition; the lower themelting point of the addition, the lower will bethe melting point of the resulting eutectic (as-suming that one exists).

The effect on melting point of multiple alloy-ing additions, all with similar melting points, canbe deduced as follows. As a simple approxima-

tion, Eq A2.1 can be extended stepwise to mul-ticomponent alloys, where, for example, the bi-nary alloy A � B may be considered as thesolution “matrix” AB of a pseudobinary alloyAB � C. Then Eq A2.1 may be rewritten in theform:

�Hm(2)

R ( 1

T (3)m

�1

T (2)m

) � �ln ( x2

x 2* )

with the following conditions satisfied:

�H(2)m � �H(1)

m ; T (3)m � T (2)

m � T (1)m ;

x2

x 2*

�x1

x 1*

The relationship x2�x2* � x1�x

1* implies that

for the third component, C, the attainable liqui-dus depression T

m(3) � T

m(2) is in most cases sig-

nificantly smaller than for the second componentB, equal to T

m(2) � T

m(1) . If this sequential pro-

cedure is applied to additional elements, the gen-eral formula for the ith constituent and the cor-responding liquidus temperature, T

mi is

obtained:

�Hm(i�1)

R. ( 1

T (i)m

�1

T (i�1)m

) � �ln (xi�l

xi�1* ) (Eq A2.2)

This expression is a member of a series, of de-creasing size for increasing value of i. The melt-ing-point depression is maximized for a largedifference between the concentration of the “ma-trix” xi�1 and its solid solubility limit x

i�1* .

Despite the fact that this model considerablyoversimplifies reality, it accounts for the two prin-cipal features in common with the experimentalresults on eutectic alloys, namely:

• The progressive reduction in melting tem-perature as the number of alloying constitu-ents is increased

• The asymptotic narrowing of the melting-point depression of the alloy with the intro-duction of each additional constituent


Composition-specific thermochemical data onmulticomponent alloys (�Hm, xi) are needed inorder to apply EqA2.2 in calculating the liquidustemperature depression through progressive al-loying. This information is mostly unavailablefor ternary and higher-order systems. However,from knowledge of thermochemical data for puremetals and binary alloys, it can be inferred thatthe respective values will differ widely from onemetal to another, and therefore large variations intemperature drop are to be expected betweendifferent alloying additions.

The physical picture of this behavior may bemore clearly understood in terms of the entropychanges accompanying progressive alloying. Atthe microscopic level, a material system may beviewed as an ensemble of atoms or moleculesand, on this basis, entropy provides a measure ofthe degree of atomic or molecular disorder in thesystem, according to the relationship:

S � k ln

where S is entropy, k is the Boltzmann constant,and is the degree of disorder, as measured bythe number of different distributions available tothe atoms or molecules in the system.

In the simplest case, where the volume of thematerial is shared by i different species of atom,representing different constituents of an alloy,each present in amounts N1, N2, N3, …, Ni, suchthat:

N1 � N2 � N3 � … Ni � N

then assuming that the different species of atomare equally interchangeable, the number of ways, , inwhich all the atomsmay be arranged amongthe N available sites is:

�N!

N1! N2! N3! … Ni!

and

S � k lnN!

N1! N2! N3! … Ni!

As the number of constituents increases, whilekeeping the total number of atoms, N, constant,

there is a tendency for the entropy to increaseas the number of constituents in thematerial is in-creased, as follows: If Ni �10 for all values of i,then Stirling’s Formula can be applied, whereby:

ln Ni! � Ni ln Ni � Ni

and the change in entropy in increasing the num-ber of constituents from two (binary system) tothree (ternary system) is:

�S2,3 � kN [ln 3 � ln 2] � R [ln 3 � ln 2]

in the particular case where N1 � N2 � 1⁄2 N forthe binary system and N1 � N2 � N3 � 1⁄3 N forthe ternary system.

Then, on increasing the number of constitu-ents from three to four (quaternary system), theentropy rises further by an increment:

�S3,4 � R [ln 4 � ln 3]

�S3,4 � �S2,3

where N1 � N2 � N3 � N4 � 1⁄4 N for the qua-ternary system.

In general:

�Si�1,i � R [ln i � ln (i�1)]

where N1 � N2 � N3�…�Ni�1 � Ni � (1/i)Nand i � 2.

The pattern is established where each succes-sive addition increases the entropy of the systemoverall, but by progressively smaller amounts, asshown by the data in Tables 2.12 to 2.14. In otherwords, each additional constituent has a rela-tively smaller effect on the degree of disorder ofthe system, as one might intuitively expect.

Entropy, S, is related to the Gibbs free energy,G, by the relationship:

(dG

dT)P

� �S

Therefore, as the entropy increases, the depres-sion of the Gibbs free energy of a system as a


function of temperature increases. This in turnwill tend to depress its melting point, althoughthe actual relationship will be governed by thespecific free energies of the constituents in themolten and solid states and of the solution thatthey form. In general terms, the picture providedby this elementary expression is consistent withthat furnished by Raoult’s law and the Clausius-Clapeyron equation.

At first sight, this model may appear inappro-priate for a eutectic alloy. However, the specialcase of a eutectic alloy with a low degree ofintersolubility of the pure metal constituents inthe solid approximates reasonably well, insofaras each phase is tantamount to a pure constituentand is well dispersed throughout the alloy. Thismodel therefore serves as a crude, but neverthe-less graphical illustration of the physical effect ofprogressive eutectic alloying.

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CHAPTER 3

The Joining Environment

WHENCONSIDERING the metallurgical as-pects of soldering in Chapter 2, it is assumed thatcomponents and the filler were perfectly cleanand remained so throughout the process cycle,enabling the constituents to freely interact so thatthe filler metal wets and spreads over the com-ponent surfaces. However, this situation repre-sents the ideal case because oxides and othernonmetallic species are usually present on sur-faces that have been exposed to ambient atmo-spheres and these will interfere with or inhibitwetting and alloying. Any oxygen or moisturepresent in the joining environment will furtherexacerbate this effect, particularly as the kineticsof oxidation reactions are highly temperature de-pendent. Thus, the nature and quality of jointsdepend not only on alloying reactions, but alsoon the processing environment—in particular onwhether the surroundings are oxidizing, reduc-ing, or neutral. The term “surroundings” refers toboth the gas atmosphere itself and any chemi-cals, such as fluxes, that are in the vicinity of theworkpiece. These aspects are considered in sec-tions 5.2 and 5.3, Chapter 5.

Materials used in joining, whether solders,fluxes, or atmospheres, are becoming increas-ingly subjected to restrictions on the grounds ofhealth, safety, and pollution concerns. Theseregulations can limit the choice of materials andprocesses that are deemed acceptable for indus-trial use. This issue is a reoccurring theme andis particularly addressed in Chapter 1, section1.3.3 and Chapter 3, section 3.2 in the context ofsoldering fluxes andChapter 5, section 5.1,whichcovers lead-free solders.

Most nonmetallic materials are not wetted bymost conventional solders, evenwhen these haveclean surfaces. Where wetting does occur, thecontact angle between the molten solder and theparent material is often high, and thus the solderdoes not spread over the component surfaces.

This situation cannot be remedied with the helpof chemical fluxes because these are unable tochange the physical properties of the intrinsicmaterials that govern the wetting characteristics,as explained in Chapter 1.

Wetting and spreading of a solder on nonmet-als can be induced by incorporating within thesolder highly active elements, such as titanium,that react chemically with the base materials toform interfacial compounds that the solder canwet. Although the manner in which reactive fill-ers promote wetting is normally different fromthat of chemical fluxes, they can also be used topromote wetting of oxidized metal surfaces andthereby provide a fluxing action. Owing to thefact that the active constituent of the solder canreduce the oxides of less refractorymetals, it willremove this surface film, enabling wetting toproceed in a conventional manner through al-loying. Reactive solders are described in Chapter4, section 4.1.2.2.Active brazes, which are muchmore common, are discussed separately in theplanned companion volume Principles of Braz-ing; these are commercially available andwidelyused.

3.1 Joining Atmospheres

There are many types of assembly that de-mand soldering under a protective atmosphere.These include assemblies intended for service ina vacuum environment, which must be free fromvolatile contaminants and parent metal compo-nents that are disfigured by oxide scale. The cat-egories of joining atmospheres that are availableand their interrelationships are shown in Fig. 3.1.Generally, fluxes are needed only when carryingout the joining operation in air or other oxidizingenvironments.



Two distinct types of atmosphere are used forsoldering, namely:

• Chemically inert gas atmospheres (e.g., ar-gon, nitrogen, helium, vacuum). These func-tion by excluding oxygen and other gaseouselements that might react with the compo-nents to form surface films and inhibit flow-ing of and wetting by the solder.

• Chemically active atmospheres, both gasesand fluxes, which are designed to react withsurface films present on the components and/or the filler metal during the joining cycle andremove them in the process. These atmo-spheres may either decompose surface films(as does hydrogen when acting on certainoxide or sulfide layers, for example) or reactwith the films to produce compounds that canbe displaced by the molten filler metal. Tra-ditional rosin fluxes predominantly functionin this manner.

Controlled gas atmospheres require a confin-ing vessel, and this invariably means a furnaceof some type. Furnace joining also offers otheradvantages:

• The process may be easily automated for ei-ther batch or continuous production, becausethe heating conditions can be accurately con-trolled and reproduced without the need formuch operator skill.

• Furnace joining offers uniform heating of thecomponents of almost any geometry and is

suitable for parts that are likely to distort ifheated locally.

• The atmospheric protection afforded leads toeconomies on the use of flux and on finishingoperations, such as cleaning and the removalof flux residues.

Against this must be considered the followingpotential disadvantages:

• Capital costs of the equipment, including theassociated gas atmosphere handling orvacuum system, may be significant in rela-tion to the processing costs.

• The entire assembly is heated during the pro-cess cycle, which can result in a loss of me-chanical properties, even to components re-moved from the joint area.

• The range of permissible parentmaterials andsolders tends to be restricted to elements andchemicals of low volatility to avoid contami-nation of the furnace. For a similar reason,most fluxes are undesirable.

Certain metals are embrittled on heating in thepresence of standard gas atmospheres (oxygen,nitrogen, hydrogen, and carbon-containinggases) and must therefore be joined in a vacuumfurnace. These are principally the refractorymet-als beryllium, molybdenum, niobium (colum-bium), tantalum, titanium, vanadium, and zirco-nium, but really only at brazing temperaturesabove 750 °C (1380 °F), which are the same asthose required for using most activated solders.

Fig. 3.1 Interrelationship of joining atmospheres


Compound semiconductors (e.g., gallium ars-enide) can have their electrical functionality poi-soned by hydrogen, while some brasses are in-tolerant of ammonia. Thus, the requirements ofeach component in an assembly must be indi-vidually assessed and the atmosphere chosen tosuit.

3.1.1 Atmospheres andReduction of Oxide Films

Aprincipal process requirement for successfulsoldering is to ensure that the joint surfaces arefree from oxides and other films that can inhibitwetting by the molten solder and the formationof strong metallic bonds. The ability to removea layer of oxide from a given metal depends onthe ease of either physically detaching the filmfrom the underlying metal or of chemically sepa-rating the oxygen ions from the metallic ionspresent in the oxide, that is, the strength of therelevant molecular bonds. Chemical reduction ofmetal oxide by atmospheres is considered first.

Chemical thermodynamics can be used to de-termine the propensity for a metal to spontane-ously oxidize or, conversely for an oxide to dis-associate. An understanding of oxide reductionby gas atmospheres is very relevant to brazing,and for that reason an introduction to the subjectis provided in the planned companion volumePrinciples of Brazing. At typical soldering tem-peratures there is insufficient driving force avail-able to reduce or prevent oxidation of almost allmetals, save those classified as noble and underlimited circumstances, possibly also extendingto silver and copper. For this reason, only a verysuperficial treatment of chemical thermodynam-ics is provided here, sufficient only to grasp themost important fact that gaseous atmospheres,including pure hydrogen, are generally inert to-ward oxides of solders and common substratematerials at normal soldering temperatures.There are, of course, exceptions, and some ofthese are elaborated on later in this chapter.

A measure of the strength of a metal-to-oxygen chemical bond is given by the change intheGibbs free energy that occurswhen thatmetalreacts to form the oxide, as detailed inAppendixA3.1. Here, it is noted that the Gibbs free energy,G, is an important thermodynamic function inchemistry because incremental changes in itsvalue only involve incremental changes in pres-sure, P, and temperature, T, for reversible reac-tions:

dG � VdP � SdT

V is the volume and S is the entropy, which isexplained further in Appendix A3.1 Chemicalreactions, such as oxidation and reduction, whichare reversible, can take place at constant pressureand temperature, so that the Gibbs free energy ofthe material system does not then change in thecourse of the reaction. Table 3.1 shows the Gibbsfree energy of formation of oxides for a selectionof metals at room temperature. This formationenergy is sometimes referred to reciprocally asthe dissociation potential of the oxide. The leaststable metal oxides are those of the noble metals,gold, silver, and members of the platinum group.These metals are therefore the most readily sol-dered, while the refractory metals and the lightmetals—notably aluminum, beryllium, andmag-nesium—have particularly stable oxides so thatthese metals are the most difficult to join. It isprecisely because gold will not form a stableoxide in air that it is widely used as a surfacecoating for fluxless joining processes (see Chap-ter 3, section 3.3.8.1).

Other factors need to be considered in con-nection with oxide reduction. In particular, manymetals form different oxides of varying stabil-ity—for example, cuprous oxide (Cu2O) and cu-pric oxide (CuO). Furthermore, oxides formedon alloy surfaces are not generally pure metal

Table 3.1 Comparative values for free energiesof formation of metal oxides of common solderconstituents and selected metals at roomtemperature (25 °C or 77 °F)

The more negative the value, the more stable the oxide.

Element Common oxide

Free energy of

formation at

25 °C (77 °F),

kJ/mol (rounded values)

Gold Au2O3 �50Silver Ag2O �10Copper CuO �130

Cu2O �150Bismuth BiO �170

Bi2O3 �460Lead PbO �190

PbO2 �210Pb3O4 �570

Tin SnO �260SnO2 �490

Zinc ZnO �300Antimony Sb2O3 �580

Sb2O4 �740Sb2O5 �780

Indium In2O3 �620Chromium Cr2O3 �700Titanium TiO2 �840Aluminum Al2O3 �1580

Chapter 3: The Joining Environment / 105

oxides but rather compound or other forms ofmixed oxide. Often these are of nonuniform com-position and structure, adding further complexityto the subject. This is particularly true of soldersthat are almost inevitably multicomponent al-loys. In its present state of development, chemi-cal thermodynamics is not able to predict accu-rately conditions under which dissociation ofoxides will occur, but can only provide a semi-quantitative indication, particularly when the ki-netics of reaction are taken into account. Owingto such complexities, the thermodynamic prin-ciples for analyzing oxide reduction are consid-ered only for pure metals.

3.1.2 Thermodynamic Aspects ofOxide Reduction

All chemical reactions are reversible, includ-ing oxidation reactions. In general, the oxidationof any metal can be described by an equation ofthe form:

nM �m

2O2 ↔ MnOm

The reaction will proceed spontaneously in ei-ther direction—namely, oxidation of the metal,or, conversely, reduction of the oxide, if it isenergetically favorable to do so. A condensedtreatment of relevant thermodynamic functionsand their relationships, which has by necessityrequired a degree of oversimplification, is givenin Appendix A3.1 at the end of this chapter. Amore rigorous treatment is given in standard text-books on thermodynamics, such as those listedin the Selected References given in the Preface.

The free-energy change (�G) for oxidation re-actions involving a series ofmetals canbe chartedon a diagram as a function of temperature, asshown inFig. 3.2.This representation is knownasanEllinghamdiagramor aRichardson-Jeffes dia-gram. The diagram can be used to determinewhether, in principle, an atmosphere is capable ofreducing surface oxides, although it does not pro-videany indicationof thekineticsof the reactions.The use of the Ellingham diagram in solderingpractice is described in the next section.

At any given temperature, the smaller the equi-librium partial pressure of oxygen in the metaloxide, the stronger the bond between the oxide

Fig. 3.2 Simplified Ellingham diagram showing the free-energy change for oxidation of several metals. Oxide stability is reducedby elevated temperature and decreased oxygen partial pressure. Each dashed line corresponds to the Gibbs free-energy

change as a function of temperature, relating to a particular oxygen partial pressure. mpt, melting point


and the parent metal, that is, the greater is thestability of the oxide. The partial pressure of agas in an atmosphere is defined in Chapter 1,section 1.3.2.5. Thus, the tendency for the oxideto decompose will be greater the lower the oxy-gen content of the atmosphere and the higher thetemperature. The further down the diagram aparticular metal-oxygen reaction curve lies, themore inherently stable is the oxide, and it is cor-respondingly more difficult to reduce; in otherwords, higher temperatures and atmospheres oflower oxygen content are required to effect re-duction.

3.1.3 Practical Application of theEllingham Diagram

3.1.3.1 Soldering in InertAtmospheres and Vacuum

Formanymetals, heating alone in air is not ad-equate to reduce the oxide, because the compo-nents are degradedor evenmelt before the criticaltemperature,Tc, is reached atwhich the oxidewillspontaneously decompose. Moreover, the rate ofoxidationroughlydoubleswitheach25°C(45°F)rise in temperature. Thus, stable oxides becomeprogressively thicker and more tenacious, andconsequently more difficult to remove, over thetime interval that the component is being heatedto the critical temperature. Excessive oxidationcan damage component surfaces, particularly ifthe film spalls off locally, because the rate of oxi-dation will be nonuniform over the surface, pro-ducing an unsightly finish. For these reasons,wherefluxes are not employed, it is usual practicetoheat thecomponents inaprotectiveatmosphereor vacuum, which will both protect the surfacesfrom further oxidation and reduce the partial oxy-gen pressure, and hence the critical temperature.

The conditions of temperature and oxygen par-tial pressure required to spontaneously reduce ametal oxide can be deduced from the Ellinghamdiagram. Reduction will occur when the free-en-ergy curve for metal-oxide formation lies abovethe oxygen partial pressure curve at the tempera-ture of interest; that is, the oxygen pressure in theatmosphere is less than that which will cause themetal under consideration to oxidize. Thus, thecritical temperature for the reduction of PdO de-creases from 920 °C (1688 °F) in pure oxygen atatmospheric pressure to 380 °C (715 °F) if theoxygen partial pressure is decreased to 10�10 atm(10�2 mPa). It can be seen from themore detailed

Ellingham diagram reproduced in Fig. 3.3, thatoxide reduction in vacuum is practicable only forpalladiumandsilveratnormal soldering tempera-tures (<450 °C, or 840 °F). The minimum partialoxygen pressure that can be achieved using high-quality industrial equipment is of the order of10�10 atm (10�2 mPa). Note that it is convenientto use the atmosphere as the unit of pressure inthermodynamic calculations, and this conventionis also applied to Ellingham diagrams.

For metals having oxidizing reaction curvesthat are located below the 10�10 atm (10�2 mPa)oxygen partial pressure curve (that is, a line join-ing the pointOon theT �–273 °Caxis on the left,to the 10�10 atm value on the partial oxygen pres-sure (PO2

) scale on the right of the Ellingham dia-gram,asshowninFig.3.2), itwillbeenergeticallyfavorable for themetal to oxidizeby reactionwiththe residual oxygen and any water vapor presentin the furnace atmosphere. This includes mostcommonsolder constituents, including tin and in-dium. Thus, industrial-quality vacuum and inertatmospheres are incapable of preventing degra-dation of solders during normal heating cycles.Obviously, an atmosphere that is largely free ofoxygen and water vapor will greatly slow furtheroxidation, but cannot prevent or reverse it.

Asmentionedpreviously, caremust be taken toselect an atmosphere that is inert toward all of themetals in the assembly being joined.Vacuum candegrade certain materials, notably brass, even atsoldering temperatures, due to the loss of zincthrough volatilization—a consequence of thehigh vapor pressure of this element. Likewise,lead-containing solders are unstable in highvacuum at temperaturesmuch above 300 °C (570°F) and are not recommended for use under theseconditions.Table3.2 lists theboiling/sublimationtemperatures of selected elements at 10�10 atm(10�2mPa).Formetals tobe joinedunder reducedpressure, theprocess temperaturemust beconsid-erably less than the boiling/sublimation tempera-ture (by a factor of <½ inK/K), if volatilization isnot to be significant.

The oxygen partial pressure in a vacuum fur-nace can be reduced substantially below the gaspressure in the vacuum by repeatedly pumpingout and backfilling the chamber with a dry, oxy-gen-free gas (seeChapter 1, section 1.3.2.5). Caremust be taken to ensure that the inlet system iscompletely leak-tight. Otherwise, some oxygenwill be bled into the furnace, and this will impairor even nullify the benefit of the inert atmo-sphere. A periodic flushing of the chamber withan inert gas will also serve to minimize any


buildup of oxygen released in the dissociation ofoxides during the heating cycle.

The effectiveness of using the process tem-perature and oxygen partial pressure to controloxide reduction, or at least prevent oxidation, is

limited by the presence of adsorbed water vaporon the walls of the vacuum chamber and on otherfree surfaces. The desorption of water vapor ef-fectively increases the oxygen partial pressure inthe chamber, and this has a deleterious effect on

Fig. 3.3 Ellingham diagram for selected oxides. M, melting point of metal; B, boiling point of metal; M', melting point of oxide


the oxide-removal process. Therefore, it is goodpractice to heat the walls of the chamber to pro-mote desorption, while simultaneously remov-ing the vapor from the chamber by alternatelypumping out and/or flushing with dry, inert gasbefore commencing the heating cycle. Addi-tional background information on moisture inclosed spaces in given in Appendix A4.2. Nev-ertheless, as can be seen from the nomogram inFig. 3.2, the quality of vacuum required to pre-vent adsorption on to the faying surfaces needsto exceed that which can usually be economi-cally achieved in an industrial process.

Large-scale industrial processes often rely onliquid nitrogen for several reasons. Not least ofthese is the ease of convenience of delivery andstorage. Furthermore, nitrogen boiled off from acryogenic tank containing the liquefied gas pos-sesses lower levels of oxygen and water vapor(typically <2 ppm combined) than all but thepurest grades of bottled nitrogen. It is also rela-tively inexpensive, being comparable in priceper liter to bottled mineral water. Owing to in-creasingly stringent environmental legislation,joining in inert atmospheres is gaining in popu-larity. This is particularly true in the electronicsindustry, where the trend is toward nitrogen at-mosphere furnaces for both wave and reflow sol-dering processes. In commercial systems, the ni-trogen ambient contains less typically than 10ppm of other species. The running costs associ-

ated with the large volumes of nitrogen that arerequired to achieve this quality of atmosphereare offset by the ability to dispense with post-joining treatments because reduced quantities offluxes and cleaning fluids are required, with theirassociated health and environmental problems.

For certain applications, inert gases other thannitrogen may be more appropriate. Of these lesscommon inert gases, argon and carbon dioxideare probably the most widely used. Both can bepurchased in high-purity form. Carbon dioxide isoften recommended in applications where theatmosphere is confined, but open to air at variousportals, because the greater molecular weight ofcarbon dioxide enables it to displace air moreeffectively than does nitrogen [Esquivel andChavez 1992].Argon is more expensive than theother two gases, and its use is therefore largelyconfined to joining in closed volumes.

3.1.3.2 Soldering inReducing Atmospheres

If the partial oxygen pressure surrounding theworkpiece cannot be sufficiently lowered to ef-fect oxide removal by introducing a vacuum orinert gas environment, then a reducing atmo-sphere might be able to remove the oxide. Thethree most widely used reducing gases are hy-drogen, carbon monoxide, and “cracked” am-monia (that is, ammonia dissociated into nitro-gen and hydrogen).

The basic chemical reduction processes forhydrogen and carbon monoxide are:

yH2 � MxOy → xM � yH2O

and

yCO � MxOy → xM � yCO2

For convenience to the user, the Ellingham dia-gram is provided with a series of side scalesgiving the partial oxygen pressure correspondingto ratios of H2/H2O and CO/CO2, as shown inFig. 3.3. On the left-hand side is shown an axisfor T � �273 °C (0 K) with points marked atvalues of the free energies, �G, at this tempera-ture for the hydrogen/oxygen (point H), the car-bon monoxide/oxygen (point C) and other reac-tions. Each of these points is associated with oneof the side scales shown in Fig. 3.3.

Table 3.2 Boiling/sublimation temperature ofselected elements at a pressure of 10!10 atm(10!2 mPa)

Boiling/sublimation temperature

Element ºC °F

Cd 100 212Zn 150 302Mg 210 410Sb 300 572Bi 350 662In 525 977Mn 550 1022Ag 630 1166Al 725 1337Sn 730 1346Cu 780 1436Cr 800 1472Au 880 1616Pd 905 1661Fe 950 1742Co 1020 1868Ni 1025 1877Ti 1130 2066Mo 1680 3056W 2230 4046

Values are rounded. Note the high position of tin and the low position of man-ganese and zinc in the table in relation to their melting points.


As an example of the use of the Ellinghamdiagram, consider the conditions for reduction ofchromium oxide. The free energy of formation ofCr2O3 as a function of temperature is representedby curveAB in Fig. 3.4. Values of the free energyof formation of water vapor from the reaction ofhydrogen with oxygen are represented by a fam-ily of curves diverging from the point H, eachcurve corresponding to a different molar ratioH2/H2O in the atmosphere. When curve ABcrosses a particular curve belonging to the familyof water vapor curves representing different wa-ter vapor/hydrogen partial pressure ratios, thechromium/oxygen and hydrogen/oxygen reac-tions are in equilibrium because their respectivefree energies are the same. This means that theoxygen potentials for the two reactions are iden-tical. When curveAB lies above the water vaporcurve at a particular temperature, chromium ox-ide will be spontaneously reduced to form chro-mium and water vapor by the hydrogen becausethe latter combination is more stable than Cr2O3.The reverse is true when curveAB lies below the

water vapor curve. For the equilibrium betweenchromium/oxygen and water vapor/hydrogen re-actions to be achieved at a soldering temperatureof 400 °C (750 °F), the H2O/H2 ratio must belower than 10�10—indicated by lines AB andHQ. This condition cannot be achieved in prac-tice, so that hydrogen is ineffective in removingchromium oxide in soldering operations.

It is evident from the Ellingham diagram thatthe stability ofmetal oxides decreases as the tem-perature is increased and the oxygen partial pres-sure is reduced. Commercial supplies of hydro-gen, nitrogen, and other gases will inevitablycontain some oxygen and water vapor, which areoxidizingagents, andcomponent and furnace sur-faceswillusuallycontributedesorbedwater to theprocess atmosphere. In addition, atmosphereswith a high concentration of hydrogen present anexplosion risk, which usually precludes their use.

Although not given in Fig. 3.3, the reactioncurve for tin and its common oxide lies in closeproximity to that for carbon monoxide/dioxide.Thus, at 400 °C, it is theoretically possible to re-

Fig. 3.4 Simplified Ellingham diagram illustrating the graphical method for determining the temperature and H2O/H2 ratio that willspontaneously reduce ametal oxide tometal (here, Cr2O3 to Cr). The set of dashed lines corresponds to theGibbs free-energy

change as a function of temperature for the reaction of hydrogen with oxygen to produce water vapor for different H2O/H2 ratios.


duce tin oxide to tin with aH2O/H2 ratio as low as10�2.However, the reaction rate is so slow that, inpractice, the gases hydrogen and carbon monox-ide are not effective for reducing the oxides ofmost common industrialmetals, including tin, be-low about 500° C (1380 °F) at readily obtainableoxygen partial pressures. The issue of reactionrate is discussed further in section 3.3.3 in thischapter.

3.1.3.3 Alternative Atmospheres forOxide Reduction

Other gases such as chlorine and fluorine aremore effective than hydrogen and carbon mon-oxide at removing surface oxides of particularmetals, as is clearly indicated on the relevantEllingham diagrams [Wicks and Block 1963].The power of chlorides and fluorides as cleaningagents accounts for the effectiveness of chloro-fluorocarbons (CFCs) and the difficulty experi-enced in replacing them by more environmen-tally friendly chemicals, following theimplementation of the Montreal Protocol on Sub-stances That Deplete the Ozone Layer: 1991[McLaughlin et al. 1998; Lea 1991].

Such gases partly operate by converting theoxide to a halide that is volatile at the joiningtemperature and that vaporizes during the heat-ing cycle. These halide atmospheres also chemi-cally attack the underlying metal and physicallyundermine the oxide, as occurs in the fluxing ofaluminum. This point is discussed in section3.2.2.1 in this chapter.

Adifferent option, particularly for tin-base sol-ders, is to use a process atmosphere that containsa chemical flux in its vapor state. Formic acid isone such example. Owing to the absence of aliquid flux, joining under gaseous fluxes is gen-erally misdescribed as a “fluxless” process. Fur-ther information on chemically active atmo-spheres is to be found in section 3.3.6 in thischapter.

3.1.4 Forming Gas as anAtmosphere for Soldering

Despite the scientific knowledge embodied inthe preceding discussion, it is a common mis-conception that “forming gas” (nitrogen/hydro-gen mix) is a better choice than nitrogen as aprocess gas for soldering because it will reducetin and indium oxide. Figure 3.5 shows the resultof melting controlled amounts of four commonsolders in atmospheres of pure nitrogen, nitrogen-

3%hydrogen, and nitrogen-40%hydrogen. Fromthis trial it can be concluded that there is noparticular advantage in using forming gas at smallexcess temperatures above the solder meltingpoint. In fact, the European specification for com-mercial grade forming gas allows it to contain ahigher level of oxygen and water vapor than isfound in bottled nitrogen. In the United States,dry nitrogen and forming gas have roughly simi-lar specifications, but forming gas carries a pricepremium as a specialty gas mixture. Experimen-tal investigators are advised that 40% hydrogenis well above the explosive limit for this gas inair, so its use requires many safety features tocomply with health and safety legislation ormakes for something of a “white knuckle” ridefor the process operator! However, as explainedin section 3.3, if a process using higher than nor-mal soldering temperatures can be adopted, thena dry hydrogen atmosphere may facilitate flux-less soldering, particularly for the higher-melting-point gold-base solders, on account ofthe greater thermal activation available and thehigher intrinsic nobility of the alloys.

It is sometimes claimed that better results interms ofwetting and spreading are obtainedwhenusing forming gas compared with pure nitrogen.Frequently, this has more to do with the fact thatthe quality (i.e., leak tightness) of the furnace andgas control and conveyance system used withforming gas are of a higher standard than for anitrogen furnace, as required by safety protocolsand hydrogen has a thermal conductivity seventimes that of nitrogen (see Table 3.3). Thus, forthe same process temperature setting, the partsare likely experiencing additional superheatwhen reflowed in forming gas.

3.2 Chemical Fluxes for Soldering

Successful soldering is largely dependent onthe ability of the solder to wet and spread oncomponent surfaces. A major barrier to wettingis presented by stable nonmetallic films and coat-ings on the surfaces, in particular oxides andcarbonaceous residues. Oxide films on the fayingsurfaces present more than a physical barrier towetting and spreading. Oxides are typically poorthermal conductors, compared to metals, and actas barriers to heat transfer, thereby exacerbatingtemperature gradients present and delaying fu-sion of the solder with the parent metal. Thethermal contact resistance of electroplated and


then reflowed lead-tin eutectic solder on copperis known to increase with the logarithm of the

oxide thickness [Di Giacomo 1986]. Surfaceswith a coarse-grained microstructure lose sol-

Fig. 3.5 Spread tests of four common solders, melted on NiCr/Au substrates at 10 °C (18 °F) superheat, in controlled atmospheres.There is negligible benefit from a hydrogen-rich atmosphere. Although somewhat subjective, the solders melted in the 40%

hydrogen atmosphere do appear to be “cleaner” with a brightermetallic luster. The holes in the substrates are of identical size and spacingand act as scale markers. The solder disks measured 6 mm (0.24 in.) diameter by 25 μm (1 mil) thick and were only superficially cleanedbefore use. (a) Ag-80In-15Pb solder melted in nitrogen. (b) Same alloy melted in nitrogen-3% hydrogen. (c) Same alloy melted innitrogen-40% hydrogen. (d) Ag-96.5Sn solder melted in nitrogen. (e) Same alloy melted in nitrogen-3% hydrogen. (f) Same alloy meltedin nitrogen-40% hydrogen. (g) Au-20Sn solder melted in nitrogen. (h) Same alloy melted in nitrogen-3% hydrogen. (i) Same alloy meltedin nitrogen-40% hydrogen. (j) Au-12Ge solder melted in nitrogen. (k) Same alloy melted in nitrogen-3% hydrogen. (l) Same alloy meltedin nitrogen-40% hydrogen. Source: BAE Systems


derability much more readily than those withfine-grained microstructures because the oxidefilms grown on the former are more continuous,with fewer defects. Fluxes are chemical agentsthat are used to remove these layers and therebypromote wetting by the molten filler.

In order to be effective in exposing a baremetal surfaces, a flux must be capable of ful-filling the following functions:

• Removal of oxides and other films that existon surfaces to be joined by either chemical or

Fig. 3.5 (continued) (g) Au-20Sn solder melted in nitrogen. (h) Same alloy melted in nitrogen-3% hydrogen. (i) Same alloymelted in nitrogen-40% hydrogen. (j) Au-12Ge solder melted in nitrogen. (k) Same alloy melted in

nitrogen-3% hydrogen. (l) Same alloy melted in nitrogen-40% hydrogen. Source: BAE Systems


physical means, often involving reaction ofthe flux with surface oxides to form metalsalts, which are then dissolved by the flux

• Protection of the cleaned joint from oxidationduring the joining cycle

• Wetting the joint surfaces, but being dis-placed by the molten solder as the latterspreads

While molten, fluxes form a thermal blanketaround the joint that helps to spread the heatevenly during the heating cycle. The flux alsotends to reduce the surface tension between thesolder and the joint surfaces, thereby enhancingspreading.

Ideally, the flux should leave no residues orproduce residues that are easily removed by, forexample, being soluble in water. It should also becompatible with the filler and substrate materi-als. For example, ammonia-containing fluxes arenot suitable for brass components, because in-tergranular corrosion can result through chemi-cal reaction. Chemical fluxes always functionwhile in a gaseous or liquid form, although theactive constituents are frequently solid at roomtemperature.

Fluxes can be introduced to the joint in a num-ber of ways, the most common of which arediscussed here. A flux can be applied in the formof a powder, paste, or liquid immediately priorto the heating cycle. The joint is then heated tothe required bonding temperature, bywhich pointsolid fluxes have become molten, ideally justbefore the filler metal melts. In wave soldering,a liquid flux is applied to a circuit board using afoam or spray dispenser only seconds, or less,before the part enters the solder wave (see Chap-ter 1, section 1.3.2.2).

Aflux can also be placed within or adjacent tothe joint together with the filler metal as a pre-form and the assembly heated to the bondingtemperature. As a properly chosen flux will meltat a temperature below the melting point of thefiller, the molten flux is able to spread over thejoint surfaces and clean them before the fillermetal melts and displaces the flux.

Another method involves introducing the fluxtogether with the filler into a joint already heldat the bonding temperature, in the form of flux-cored solder wire. Although this technique iswidely practiced because it is fast and conve-nient, it is not recommended because the heatedcomponent surfaces are unprotected until thefiller is applied. More aggressive fluxes are thenrequired, which in turn tends to accentuate cor-rosion and cleaning problems.

Alternatively fluxes can be applied togetherwith the filler, prior to the heating cycle, in theform of pastes and creams, which are normallyproprietary formulations. They comprise mix-tures of the fillermetal,which is present as a pow-der of a prescribed shape and size distribution to-gether with a flux and an organic binder that isselected to produce the desired viscosity and toburn off without leaving contaminating residues.These pastes and creams are particularly useful inautomated reflow soldering operations becausethey can be screen printed or dispensed using sy-ringes. The large surface area of the powderedfillermetal in contactwith thefluxmeans that cor-rosion is inevitable during storage; therefore,these products have a finite shelf life and storagerequirements, which are best strictly observed.

The mechanisms of flux action are almost asdiverse as are the flux formulations that are com-mercially available. In many cases, the mecha-nisms have not been fully elucidated, which is inno short measure due to the commercial secrecysurrounding flux compositions. While soldercompositions are mature, flux chemistry is stillradically evolving. Today’s soldering fluxes arevery sophisticated chemicals comparedwith evena few years ago, and new products continue to belaunched almost monthly. Several different flux-ing mechanisms cover the majority of solderingoperations that are encountered. Even these aresufficiently complex not to be understood in de-tail at the present time. However, fluxing mecha-nisms can be classified according towhether theyremove the nonmetallic surface coating by physi-cal or chemical means.

A flux can chemically remove a surface oxidecoating by:

• Dissolving the coating• Reacting with the coating to form a product

that is unstable at the bonding temperature• Reducing the oxide to metal in an exchange

reaction

A surface coating can also be physically re-moved. This usually occurs through erosion of

Table 3.3 Thermal conductivities of solderingatmospheres, relative to airSoldering atmosphere Relative thermal conductivity

Carbon dioxide 0.62Argon 0.68Nitrogen 0.99Air 1Helium 5.8Hydrogen 6.9


the underlying metal. In this mechanism, the fluxdoes not react with the surface coating itself, butis able to percolate through it and react with theunderlyingmetal, thereby causing detachment ofthe coating. In addition, physical fluxing actioncan be achieved by applying mechanical agita-tion,without the need for amaterial fluxing agent,as discussed in section 3.3.5. Many fluxes func-tion by a combination of mechanisms, and forthis reason fluxing action is best illustrated withreference to specific examples.

Environmental considerations have impingedheavily on the use of certain fluxes, particularlythose incorporating organic materials (volatileorganic components, or VOCs), because thesehave tended to rely on CFCs for the removal oftheir residues, which attack the earth’s ozonelayer when they are discharged to the atmo-sphere. These olderVOC formulations have beenlargely replaced in manufacturing industry byfluxes whose residues are soluble in water orcleaning agents that do not contain substancesthat are environmentally harmful [Lea 1991; El-lis 1991] (see Chapter 1, section 1.3.2.11).

Even when soldering using flux, the employ-ment of an inert atmosphere, such as nitrogen, isbeneficial. This is particularly true for the Ag-Cu-Sn lead-free solders, which are used at muchlower superheats than lead-tin solder. The nitro-gen atmosphere helps prevent oxidation of thesolder and substrate before the flux becomes ac-tive and reduces the work that the flux has to doin maintaining oxide-free interfaces ahead of theadvancing solder front. This means that less fluxis required, making it environmentally and eco-nomically advantageous. If large superheats canbe applied, the benefits of a protective atmo-sphere are reduced because the solder spreadsmuch faster, as can be seen from Fig. 3.6 [Buck-ley 2000]. Further information on lead-free sol-ders is given in Chapter 2, section 2.4.4 andChapter 5, section 5.1.

The effect of different concentrations of oxy-gen on the results obtained when fabricatingfluxed printed circuit boards (PCBs) using lead-tin solder has been studied. In general, if theoxygen level can be decreased to below 10 ppmthen the soldering process will generally be in-sensitive to many process variables, because themolten solder has a low surface tension andshorter wetting time. Soldering processes can berealized in an inert atmosphere where the oxygenconcentration is up to 1% (10,000 ppm). How-ever, to be successful, most process parameterssuch as flux type, heating rate, peak temperature,

and time at temperature all need to be carefullyoptimized to achieve low-defect soldering. Atelevated oxygen levels, the rate of dross forma-tion on the solder becomes significant and willinterfere with the solder wetting and spreading,as can be seen in Fig. 3.7 [Nowotarski and DeWilde 1996].

Soldering in very low oxygen level atmo-spheres does not necessarily translate to signifi-cantly improved production costs through maxi-mizing yield. Indeed, just from the standpoint ofthe cost of the increased gas consumption nec-essary to reduce the oxygen level below 20 ppm,at least one study showed that it was economi-cally unfavorable (see Table 3.4) [Verite, Ver-bockhaven, andAlleaume 1997]. Obviously thisanalysis will vary with the value of the product,in this instance the PCB for a mobile telephone,which is produced in high volumes in a highlycompetitive market.

Fig. 3.6 Wetting of copper by Pb-63Sn solder using rosinflux. Soldering with flux generally benefits from a

protective atmosphere (unless the atmosphere detrimentally af-fects the chemistry of the fluxing action), because the flux has towork less to protect the substrate and filler from oxidation.

Fig. 3.7 Effect of oxygen concentration in the atmosphere onthe rate of dross (scum) formation on an exposed

solder reservoir


3.2.1 Fluxes for Tin-Base Solders

The overwhelming majority of soldered jointsmade are to interconnect electronic components.The principal material that is joined in these ap-plications is copper, due to its high electricalconductivity; the solder is almost invariably alead-tin (or lead-free tin-base) alloy. Becauseelectronic components tend to be manufacturedand stored under reasonably cool, clean, and dryconditions, they are likely to have only a thinlayer of copper oxide as a barrier to wetting.This, coupled with the fact that soldering is usu-ally performed rapidly and mostly below 300 °C(570 °F), means that the fluxes required tend notto be highly aggressive chemicals. Major con-siderations pertaining to fluxes intended for elec-trical/electronic applications, other than theircleaning ability, are the nature of the residues andthe ease of their removal. These factors are ofconcern owing to the need to avoid subsequentdeterioration of the joints and ultimately failureof the circuitry through corrosion [Turbini et al.1991].

Indium and zinc solders require fluxes with achemistry different from tin-base solders, be-cause of the composition of the oxides and thedifferent process temperatures involved.As such,they have an even more specialized formulationcompared with fluxes for tin-solders and are bestprocured from specialist suppliers and used inaccordance with manufacturers guidelines.

3.2.1.1 Soldering Fluxes That RequireCleaning

Conventional soldering fluxes contain at leastfour basic ingredients, each ofwhich has an iden-tified role [Klein Wassink 1989; Manko 2002]:

• Acids or halides to provide the cleaning ac-tion (the active constituents)

• An ingredient that is liquid at the solderingtemperature that seals and protects thecleaned surfaces against reoxidation

• Asurfactant that promoteswetting of the jointsurfaces by the active and sealing constitu-ents

• A rheological additive to suit the applicationmethod

In practice, commercial fluxes often containmore than these four ingredients in order to meetthe requirements of the soldering process. Fluxeshave been formulated that are satisfactory formost pure metals and alloys, including stainlesssteel. Noted exceptions are beryllium, chro-mium, magnesium, titanium, and some alumi-num alloys, which are classified as “unsolder-able” in air, unless coated with a different metal(see section 3.2.2 and Chapter 4, section 4.1.2)or a very special flux composition is employed(see section 3.2.2). When selecting a flux for aparticular application, it is usually good practiceto follow the manufacturer’s guidelines, becausethe effectiveness of a particular formulation tendsto be highly sensitive to the combination of met-als and process conditions with which it is used.

The higher the activity of a flux, the greater isits ability to remove surface oxides from metalcomponents, but so is the corrosiveness of theflux and its residues. The active ingredient of asolder flux can be either an inorganic acid (hy-drochloric acid is commonly used) or an organicacid (e.g., carboxylic acids). A common car-boxylic acid used in fluxes is abietic acid(R3COOH, where R is an organic radical), whichis a major constituent of rosin fluxes. In bothcases, the acid reacts with the surface cupricoxide and converts it to compounds that arereadily removed, either chemically and/or physi-cally, from the joint surfaces. Even the basicform of the chemical reactions between thesefluxes and copper oxide is complex, but can besimplistically described by:

CuO � 2HCl � CuCl2 � H2O

CuO � 2R3COOH � Cu(R3COO) 2 � H2O

where R is an organic radical.Copper chloride is soluble in water, while cop-

per abiet [Cu(R3COO)2] is miscible with rosin.Hence, the compounds that now “contain” thecopper oxide dissolve in the excess liquid flux toleave a clean metal surface when the flux is dis-placed by the molten filler metal.

Like most active flux constituents, these acidsbecome progressively more active with increas-

Table 3.4 Normalized cost analysis ofmanufacturing cell phone PCBs, illustrating theeffect that the quality of the solderingatmosphere has on manufacturing economics

AirNitrogen �

1% oxygenNitrogen <20ppm oxygen

Cost of rework 100 37 30Cost of nitrogen 0 22 32.7Total cost per PCB 100 59 62.7


ing temperature until a point is reached whentheir effectiveness ceases, because they eitherthermally decompose or boil off. The corrosiveproperties of acids at room temperature canpresent handling problems and application dif-ficulties especially when the flux must either beintroduced directly to the components or mixedwith solder to form pastes sometime before thejoining operation. However, there are salts thatliberate acid only when heated, thus avoidingthis type of problem. One of these is zinc chlo-ride, which produces hydrochloric acid by reac-tion with moisture at elevated temperatures, ac-cording to the reaction:

ZnCl2 � H2O � Zn(OH)Cl � HCl

By using a mixture of similar salts, the activa-tion temperature of the flux and its corrosivenesscan be adjusted over awide range. In all cases, theresidues are highly corrosive. Other halide fluxesoperate in a very similar manner to zinc chloride.For example, the amine hydrohalides, such as hy-drazine hydrochloride (R2NH2Cl), thermally de-composewith the liberation of hydrochloric acid:

R2NH2Cl � R2NH � HCl

The flux constituent that protects the cleanmetal surface from reoxidation usually alsoserves as the carrier for the other ingredients. Itneed only be effective for a few seconds in manysoldering processes. Alcohols, oils, esters, gly-col, and even water are capable of fulfilling thisfunction at the relatively low temperatures usedfor most soldering operations (<250 °C, or 480°F). As the flux carrier constitutes by far thelargest volume fraction of the formulation, it hasa strong influence on the aggregate propertieseven though it is not intended to be chemicallyactive. The flux carrier is usually a mixture ofchemicals formulated partly to prevent the fluxfrom boiling violently when the workpiecereaches a specific temperature.Asmentioned pre-viously, it may also play a role in mopping up theproducts of reaction.

Surfactants are added to lower the surface ten-sion of the liquid flux. The effect of minute ad-ditions of detergents, soaps, and soluble oils towater is well known, and these are often addedto water-based fluxes to ensure satisfactory wet-ting when either oil or grease are likely to bepresent on the component surfaces. The surfac-

tant needs to be inert toward the other constitu-ents of the flux and also to the clean metal sur-face. For this reason, ion-free organic complexes,similar to those widely used in the electroplatingindustry, tend to be favored.

The rheological agent is provided to impartthe correct degree of “body” to the flux to suit theapplication method, whether syringe dispensing,screen printing, or stenciling.

A common commercial designation of fluxesused for soldering is:

• R: rosin• RMA: rosin mildly activated• RA: rosin activated• OA: organic acid• IA: inorganic acid• WS: water soluble• SA: synthetically activated

This classification is somewhat loose and can bemore misleading than helpful. It does not strictlyindicate the chemical characteristics of the flux,particularly its aggressiveness to specific oxides.However, because this classification is widelyused, it warrants consideration here. An alterna-tive classification of soldering flux types, de-vised by the International Organization for Stan-dardization, is given in Table 3.5. Although thisclassification is scientifically more precise, it isless descriptive; therefore, the traditional termstend to be preferred. The different categories canbe grouped as shown below.

R, RMA, RA. Fluxes with these designationscontain rosin as the principal active chemicalingredient, the difference between them being aprogressive increase in level of chemical activityfrom R through to RA. In modern fluxes, therosin is a synthesized chemical and no longer thenatural product obtained from the sap of pinetrees.

OA. OA stands for organic acid, which pro-vides the cleaning action in this type of flux. Theorganic acid used will tend to be one of thosedescribed previously, such as a carboxylic acid.These fluxes are generally more aggressive thanRA fluxes.

IA. IA fluxes are based on inorganic acids,usually hydrochloric acid and are among themostaggressive of the available fluxes.

SA. Synthetically activated fluxes are formu-lated to have residues that are soluble in chlo-rofluorocarbon solvents. With concern growingabout atmospheric pollution, and CFCs in par-ticular, the SA fluxes have been superceded by


newer formulations that havewater-soluble (WS)residues.

WS. Water-soluble fluxes containing solublehalides have been available for several decades.These are among the most commercially devel-oped fluxes available with chemical activity tai-lored over a wide range, roughly in proportion tothe total halide content. Particular care needs tobe taken to clean off ionic residues, which cancause corrosion and generate electrical malfunc-tion. Water-soluble fluxes tend to have shorterprinting life than rosin fluxes, have less tackinessand are therefore more difficult to use for reflowsoldering.

3.2.1.2 No-Clean Soldering Fluxes

In response to manufacturers’ concern overthe cost of cleaning, a new generation of “no-clean” fluxes has come into use that are said toleave no deleterious residues (in PCB assemblyapplications). These fluxes have been formulatedwith a wide range of chemical ingredients, but allshare the characteristics of leaving residues thatare judged to be benign and so can be left on theassemblies.

No-clean fluxes mostly contain an alcohol sol-vent carrier and a small percentage of active in-gredients, which may comprise resins or simplyorganic acids. Some of the newer no-clean fluxesare water-based and are totally free of VOCs,including alcohols. Fortuitously, using water asthe solvent in place of aVOC, such as an alcohol,has been found to enhance no-clean fluxing. Inparticular, water-based no-clean fluxes appear tocope better with circuit boards that enter the sol-dering operation with surface oxide. Their en-hanced activity with respect to metal oxides has

been ascribed to a more complete ionization ofthe activators in a water base than in alcohol[Hyland and Rao 1996].

The need to clean is obviated by diluting theactivators to low levels. Therefore, these fluxescontain a much lower content of rosin, or ha-lides—referred to simply as “solids”—than dotraditional fluxes. The solids content of no-cleanfluxes is normally in the range 2 to 3 wt%, ascompared with 25 to 35 wt% for normal rosin-based fluxes. These “low-solid” fluxes are moredifficult to work with than the rosin fluxes pre-cisely because they are less active.

An alternative approach used for formulatingno-clean fluxes is to use activators of normalstrength or thereabouts in conjunction with otherconstituents that undergo polymerization whensubjected to soldering temperatures. Gelatin isone such compound. The resulting polymer frac-tion encapsulates residual active species, therebyincapacitating it. However, this mechanism canbe impaired if not all the flux is fully heated,leaving areas of the joint exposed to active resi-dues. It is also important that no-clean flux resi-dues are not cleaned because, unless the cleaningis scrupulous it leaves behind destabilized en-capsulant from which corrosive agents can sub-sequently escape.

From the foregoing outline, it is clear thatno-clean fluxes are less forgiving, being moreintolerant to variability in the soldering processconditions than conventional alternatives. Suf-ficient flux must be provided to achieve the re-quired solderability and, as noted previously, theflux must be heated uniformly and in the correctmanner to react fully and leave innocuous by-products. These considerations are consistentwith the experience of industrial users, who find

Table 3.5 Classification of soldering fluxes using the method adopted by the InternationalOrganization for Standardization

Flux type Flux basis Flux activation Flux form

1 Resin 1 Rosin2 Resin

1 Not activated

2 Halogen activated

A Liquid

2 Organic 1 Water soluble2 Not water soluble 3 Not halogen activated

B Solid

3 Inorganic 1 Salts 1 With NH4Cl2 Without NH4C1

2 Acids 1 Phosphoric acid2 Other acids

3 Alkalis 1 Ammonia and/or amines C Paste

Thus, a resin-based paste flux with a halogen activator is classed as type 122C.


that no-clean fluxes work best in assembly-linesoldering operations when the process condi-tions are optimally tuned and strictly maintainedfrom one assembly operation to another. Whilethis situation is readily achieved in a factoryusing modern equipment, it does mean that no-clean fluxes are not generally suitable for hand-soldering processes. When correctly instigated,the residues left on a PCB are not detrimental inmost applications and accelerated life-test envi-ronments [Anson et al. 1996].

3.2.1.3 Measure of CleaningEffectiveness: The SurfaceInsulation Resistance (SIR) Test

Becauseflux residues, likeother contaminants,can easily affect the performance and service lifeof electronic circuits, effective cleaning proce-dures assume a critical importance. A semiquan-titative test has been devised to assess the effec-tiveness of a cleaning process for removing fluxresidues from circuit boards and to provide assur-ance that boards soldered using no-clean fluxwillnot fail in service due to unbound residues. Thistest measures the surface insulation resistance

(SIR) to current flowbetween pairs of conductingtracks in interdigitizedcombpatternsonboardsasa function of time at a prescribed temperature andlevel of humidity. It is capable of catching elec-trochemical failure mechanisms associated withvestigial ionicresidues,namelyunacceptablecur-rent leakage under humid conditions, corrosion,metal migration, and dendritic growth.

An electric field is applied between the two in-termeshedsetsofcombs,asshowninFig.3.8.ThewidelyacceptedSIR test standard (ISO9455,part17) calls for a test pattern with 400 μm (16 mil)wide tracks, separated by 500 μm (20 mil) widegaps, and a test field of 100V/mm(2.5V/mil) (i.e.,a test voltage of 50 V).

The test environment is a constant tempera-ture of 85 °C (185 °F) and a relative humidity of85%. Surface insulation resistance measure-ments are made twice a day over a test durationof 168 h. If the surface resistance is maintainedabove 108 � throughout the test, the surface isjudged to be “clean.”

The presence of ionic contamination on thesurface of the test board can result in the forma-tion of conducting dendrites, as shown in Fig.3.9. These dendrites can grow rapidly (in min-utes), starting from residues deposited on a cath-

Fig. 3.8 International Electrotechnical Commission (IEC) test coupon for the evaluation of board cleanliness using SIR. Schematicof typical detail of an interdigitated comb is illustrated. Courtesy of Concoat Systems


ode and progressing toward a neighboring an-ode. When a dendrite has bridged the gap, theSIR value will rapidly decrease. As a dendritegrows, it will progressively carry a larger pro-portion of the available current until at somepoint it burns out and the SIR will return to ahigh value. This sequence of events repeatsuntil carbon deposits build up sufficiently topermanently lower the resistance below the ac-ceptance threshold. Examples of SIR test plots(log SIR versus time) are shown in Fig. 3.10.

A recent assessment of the SIR test has high-lighted the fact that the test procedure, as cur-rently followed, can miss detection of failureevents [Hunt 2000]. It is possible that resistancedrops, due to dendrite formation, might escapeobservation if the electrical measurements arewidely spaced in time: half a day is clearly toolong. Moreover, the bridging of a dendrite acrossneighboring conductors, which lasts for evenonly a few seconds before burnout, can repre-sent a large number of operations in equipmentoperating at a clock frequency of MHz andabove. For such equipment, intermittent den-drite bridging would cause a corresponding in-cidence of failures, which the test might notpick up.

This limitation can be overcome by usingmod-ern high-speed multichannel measuring equip-ment in the SIR test. Other recent recommen-dations to improve the test include using a lowervoltage stress (5 V/mm) and finer test structures(100 μm/100 μm, or 4mil/4mil track/gap), whichwould be more appropriate of current circuitryand operating conditions. It is also proposed tochange the test environment from 85 °C (185°F)/85% RH to a more demanding 40 °C (104°F)/93% RH.

As a result of such reviews of the SIR testprocedure, the following compromise set of con-ditions is being proposed for a new Standardspecification (draft IEC 61189-5):

• The test pattern should comprise combs witha 400 μm (16 mil) track width and 200 μm (8mil) gap width.

• The test environment should be 40 °C (104°F) and 93% RH.

• Surface insulation resistance (SIR) measure-ments should be made at 20 min intervals.

• The test voltage should be 5 V, creating anelectric field of 25 V/mm (0.63 V/mil).

• The test patterns should be “overmounted”with dummy components to better replicatethe situation in populated circuit boards,whenusing the SIR test for cleaning and other pro-cess characterization purposes.

3.2.2 Fluxes for “Unsolderable” MetalsAluminum, chromium, and some other metals

are often classified as “unsolderable” becausethey are not wetted by lead-tin solders usingcommon fluxes. Table 3.6 provides an indicationof relative solderability of some engineeringmet-als, alloys, and metallizations. The difficulty ofsoldering to many metals is an exaggerated per-ception because most fluxed solders found inlaboratories, factories, and workshops are de-signed for soldering of electronic componentsand in chemical terms are relatively mild. Theyare certainly not capable of dealing with the na-tive oxide on some of the more refractorymetals.

Aluminum and stainless steel are two engi-neering materials for which there is often a re-quirement to make a solder joint or connectionandwhere flux is permitted. Commercially avail-

Fig. 3.9 Example of conducting dendrites growing across thegaps of a test pattern on a circuit board during a SIR

test. Courtesy of Concoat Systems

Fig. 3.10 Example of an SIR test plot (log SIR vs. time), show-ing features usually associated with dendritic

growth. Courtesy of Concoat Systems


able fluxes exist for both of these “unsolderable”metals.

3.2.2.1 Aluminum Soldering Fluxes

Aluminum forms a natural refractory oxidethat is remarkably stable and tenacious. It is me-chanically durable, with a hardness that is onlyinferior to that of diamond and its high meltingpoint (2050 °C, or 3722 °F) reflects its high de-gree of physical stability.Alumina is also chemi-cally stable to the extent that it cannot be directlyreduced to the metal by aqueous reagents. Onexposure to air, a layer of alumina will formalmost instantaneously on the surface of alumi-num, and this will grow to an equilibrium thick-ness of between 2 and 5 nm (0.08 and 0.2 μin.)at ambient temperature. On heating to 500 to 600°C (930 to 1110 ºF), the thickness of this surfacecoating will increase to about 1 μm (40 μin.).Therefore, special fluxes have been formulatedfor use with aluminum alloys. These have to beparticularly effective in protecting the metal sur-face from oxidation before the solder melts andspreads.

Aluminum fluxes all contain halide com-pounds. These are highly corrosive especially inthe presence of moisture, including humid at-mospheres. Therefore, all flux residues must beremoved as completely as possible, as residuesleft behind after the cleaning procedures maycause corrosion in the vicinity of the joint.

The fluxes used for soldering of aluminum andits alloys are of two types [The Aluminium As-

sociation 1990]: organic fluxes and chloride-based fluxes.

Organic fluxes contain amines, fluoborates,and a heavy metal compound in an organic car-rier. They come in the form of viscous liquids orpowders. A typical example of this type of fluxhas the composition: 83% triethanolamine, 10%fluoboric acid, and 7% cadmium fluoroborate (aviscous liquid). Its operating range is 180 to 280°C (355 to 535 °F).

The fluxing action relies on disrupting the ox-ide, which cracks and crazes during the heatingoperation due to the differential thermal expan-sion between the metal and oxide. This enablesthe flux to come into direct contact with the alu-minum and deposit a film of the metal ion in theflux (in this instance, cadmium) onto the alumi-num surface via an exchange reaction. Organicfluxes must not be exposed to a torch or flame;otherwise they will char, and this will impedesolder flow. Aluminum alloys containing morethan about 1% Mg cannot be satisfactorily sol-dered using these fluxes, because magnesia ismore refractory than alumina and the flux is cor-respondingly less effective.

Chloride-based fluxes contain zinc or tinchlorides with ammonium chloride and fluorideand are generally applied as a water-based slurryor paste to precleaned component surfaces. Anexample of such a flux has the formulation: 88%tin chloride, 10% ammonium chloride, 2% so-dium fluoride (powder) with a working range of300 to 400 °C (570 to 750 °F). By substitutingthe tin chloride with zinc chloride, the tempera-

Table 3.6 Relative solderability of selected metals and alloys

Parent material Easy Intermediate Difficult Unsolderable

Aluminum . . . . . . X . . .Aluminum alloys . . . . . . . . . XBeryllium . . . X . . . . . .Brass X . . . . . . . . .Chromium . . . . . . . . . XCopper X . . . . . . . . .Copper-nickel . . . X . . . . . .Gold X . . . . . . . . .Invar/Kovar . . . . . . X . . .Lead X . . . . . . . . .Magnesium . . . . . . . . . XNichrome . . . . . . X . . .Nickel . . . X . . . . . .Palladium X . . . . . . . . .Platinum X . . . . . . . . .Silver X . . . . . . . . .Stainless steel . . . . . . X . . .Steel . . . X . . . . . .Tin X . . . . . . . . .Titanium . . . . . . . . . XZinc . . . X . . . . . .


ture of operation can be raised to 380 to 450 °C(715 to 840 °F).

The fluxingmechanism is essentially the sameas that for the organic fluxes—namely, one in-volving an exchange reaction whereby alumi-num on the surface is replaced by zinc or tin. Theeffectiveness of these fluxes is reduced by thepresence of silicon in the parent material, be-cause silicon is not as amenable to the exchangereaction as is aluminum.

Because both of these two fluxes operate bysubstituting for aluminum a metal that has rea-sonable oxidation resistance and that is morereadily wetted by the filler metal, they are fun-damentally different from conventional solder-ing fluxes. The latter act simply by cleaning andprotecting the original surfaces of the compo-nents. Because the replacement metal has ahigher density than aluminum, this type of flux-ing process is commonly referred to as “heavymetal deposition.”

The quantity of flux that needs to be appliedis a function of the humidity of the ambient at-mosphere. In moist atmospheres, a proportion ofthe flux is rendered ineffective through hydroliza-tion by reaction with water vapor. It has beenfound that the quantities of flux that need to beapplied can be reduced considerably by carryingout the soldering operation in a completely dryenvironment.

3.2.2.2 Stainless Steel Soldering Fluxes

Stainless steel is not an easy material to solderfor two reasons. Firstly, it is covered with anextremely stable (“self-repairing”) oxide layer,which indeed gives it its stainless characteristic.Secondly, for a metal it has an unusually lowthermal conductivity, one-thirtieth of that of cop-per, which causes complications when endeav-oring to heat joints quickly and to a uniformtemperature. Aggressive halogen-based fluxescan attack the oxide layer on stainless steel but,unless the residues are fully removed, they areprone to cause pitting corrosion. A solution ofzinc chloride in hydrochloric acid constitutes thechemically active ingredients of fluxes for stain-less steel [Mei and Morris 1992].

Phosphoric acid is also extremely effective asa flux on stainless steel, but unfortunately it pol-ymerizes at temperatures above 200 °C (392 °F),which precludes its use with the common tin-base solders. Multicomponent, lead-free, soldersbased on bismuth-tin have been devised for sol-dering to stainless steel using phosphoric acid

flux [Wakelin 1993]. The solder possesses suf-ficient ductility to survive bending several timesby hand and can be cold worked by rolling. De-tails of the solder composition have not beendisclosed, and it has not yet been developed asa commercial product.

3.2.2.3 Magnesium Soldering Flux

Magnesium is also deemed to be “unsolder-able” by conventional means. However, an in-teresting development is the discovery that mol-ten acetamid will directly dissolve the surfaceoxides on both aluminum and magnesium [Gol-ubtchik 1984]. Because this organic compoundmelts at 83 °C (181 °F), it offers a route wherebyindium-tin solder (melting point 120 °C, or 248°F) could be used to wet the exposed metal onthe substrate.An experimental trial conducted byone of the authors under very rudimentary con-ditions confirmed that the approach does havesome potential and possibly merits further in-vestigation, particularly as a method for tinningmagnesium prior to more conventional solder-ing.

3.2.3 High-Temperature Fluxes

Fluxes for soldering at high temperatures (inthe range 250 to 450 °C, or 480 to 840 °F) areformulated differently to fluxes designed to workwith lead-tin eutectic solder and alloys of similarmelting point. There are three reasons for this.

First, the higher process temperature meansthat there is more thermal activation available sothat the active constituents need not be as chemi-cally aggressive. Second, these fluxes are usuallyused with high-lead solders or gold-tin eutecticsolder. Thus, the oxide that impedes wetting andspreading on the surface of the solder is eitherlead oxide or tin oxide diluted by the noble metalso, again, decreased chemical activity is suffi-cient. The final consideration is that the flux needsto provide significantly more protection to thefiller metal and component surfaces against oxi-dation, because higher process temperatures usu-ally mean that the duration of the process cycleis longer.

For these reasons, high-temperature fluxestend only to be mildly active and are based pre-dominately on high-molecular-weight hydrocar-bons. As might be expected, the long-chain mol-ecules are very viscous at room temperature, andthese fluxes are very difficult to dispense, espe-cially on a cold morning. To overcome this prob-


lem, the flux is usually “thinned” with a lower-molecular-weight species. On heating, the lighterfraction of the mixture progressively volatilizesas the viscosity of the high-molecular-weight hy-drocarbon declines, so ensuring that the fluidityof the flux remains essentially constant through-out the working temperature range.A point to beaware of when using these fluxes, especially inconfined spaces such as flip-chip assembly op-erations, is that if the heating rate is too fast, thenon reaching about 150 °C (300 °F), the flux can“boil” and physically displace components. Thisis simply rapid volatilization of the lighter hy-drocarbon fraction. Incorporating a dwell in theheating cycle or a slow ramp through the criticaltemperature range are appropriate remedies.

In air, high-temperature fluxes produce aplume of smoke when heated to the solderingtemperature. This fume is flammable, so that sol-dering cannot be undertaken with a naked flame.In enclosed spaces the fume tends to recondenseon cooler surfaces, which requires that cleaningis undertaken on a planned basis. Higher-temperature fluxes are available in all of themod-ern designations including “water soluble” and“no-clean.”

3.3 Fluxless Soldering

The vast majority of commercial soldering op-erations involves the use of chemical fluxes. Theuse of flux makes the soldering operation toler-ant to an ambient air environment and maintainsthe solderability of surfaces through the heatingcycle, as explained in section 3.2. However, thereis a penalty to be paid when fluxes are used inthe residues that they leave behind. Flux residuesare never completely removed in normal clean-ing procedures, and the contamination can im-pair the product function, performance, and life.

High-performance electronics systems, par-ticularly those enclosed in hermetic packages,usually proscribe the use of all organic materialsin their assembly, including fluxes, as a means ofguaranteeing service life. A particularly criticalapplication is hybrid assembly for space elec-tronics, where even small traces of organic sub-stances present an unacceptable risk, owing tothe tendency of hydrocarbons to outgas. Becauseaerospace electronics are often contained withinhermetic packages, the presence of moisture inthe package environment can activate a litany ofsemiconductor failure mechanisms and is there-fore undesirable.

The problem is no less acute in relation tooptoelectronic and photonic modules. A typicalexample is a laser diode soldered to a heat sink(for example, a thermoelectric cooler) in order tostabilize the junction temperature. The joiningoperation must leave the adjacent light-emittingfacet in a pristine state. Contamination of thissurface with virtually any species will either de-grade the service life of the device or even havecatastrophic consequences for the operationalcharacteristics. Similar considerations pertain tothe assembly of tinymicroelectromechanical sys-tems (MEMS) [Kuhmann et al. 1998].

For these and other comparable applications,there is the need to use fluxless soldering pro-cesses, which will also furnish thin and well-filled joints to satisfy the thermal andmechanicalrequirements of the assembly. Fluxless solderingoffers other attractive features: it eliminates aprocessing material and more significantly interms of cost saving, processing steps includingcleaning after the soldering operation. Fluxlesssoldering also obviates a major source of voidsin joints, namely, that arising from trapped fluxproducts.

The crux of a successful fluxless solderingprocess is to eliminate all surface contaminationfrom the faying surfaces and to provide sufficientprotection of exposed surfaces from oxidationthrough the heating cycle. At the same time, del-eterious contaminantsmust be excluded from thejoining environment. A flux is designed toachieve these conditions, and if it is dispensedwith then other means have to be found to satisfythem.

As might be expected, there is direct corre-lation between oxide thickness and solderability.Similarly, there is a clear relationship betweenthe solderability of electronic component termi-nations and defect levels in volume manufac-turing of electronic circuit boards. Thus, it ispossible to establish a relationship between de-fect levels and oxide thickness on a given surfacefinish. Figure 3.11 presents results from boardtrials that were conducted in air using a nonac-tivated flux, involving copper lands. Clearly, thefirst 5 nm (0.2 μin) of oxide layer thickness hasa very significant impact on solderability and thenumber of defective joints that result.

While it is fairly straightforward to achieve anonoxidizing atmosphere of sufficient quality tolargely maintain the solderability of joint sur-faces during a rapid heating cycle, initial re-moval of surface contamination, and particu-larly thenativeoxides, ismore complex.Tackling


this problem first requires knowledge of thenature and thickness of the oxides involved.

3.3.1 Oxide Formation and RemovalDuring fluxless soldering the prime impedi-

ment to wetting and spreading is the presence ofoxide films on the surfaces of the parent andfiller metals. All base metals are covered with athin film of oxide through contact with air. Inthis respect, the noble metals, gold and plati-num, are exceptional. Surface coatings of thesemetals are exploited for their perpetual solder-ability, stemming from their inertness to oxy-gen.

The thickness of the native oxide on a basemetal surface depends very much on the historyof the particular component. The principal fac-tors governing the thickness of the oxide layerare time of exposure to air and temperature. Theoxidation of some common metals and metalli-zations at room temperature, after mechanicalcleaning of the surface, are indicated in Fig. 3.12.

This graph shows that the surface of these basemetals will be covered with more than 5 nm (0.2μin.) of oxide within 5 min of the cleaning op-eration, although the oxide layer could be sub-stantially thicker if the same metal were exposedto certain chemicals or heated in air. Obviously,this surface contamination must be removed ordisplaced by some means before wetting by thesolder can proceed.

The curves represented in Fig. 3.12 should betaken as being indicative only, because in prac-tice the growth rate is affected by a multitude offactors, including surface roughness, residualstresses, and humidity. For elevated tempera-ture, an approximate rule of thumb is that thenative oxide thickness on base metals eventu-ally doubles with every 200 °C (360 °F) in-crease in temperature. For this reason, it isdesirable to heat the joint rapidly and as soon aspossible after cleaning of the surfaces, evenwhen carrying out a soldering operation withthe benefit of flux and especially so in its ab-sence. A detailed theoretical treatment of nativeoxide growth on base metals is given by Martinand Fromm [1977].

Because oxide films grow so rapidly on mostbase metals, as shown in Fig. 3.12 and 3.13, inpractice clean and solderable component sur-faces can only be achieved if these are of goldor platinum. Accordingly, for fluxless soldering,a gold coating is generally applied to the com-ponent surfaces, which is of sufficient thicknessand adequately pore-free to ensure good solder-ability over its specified storage life. Furtherinformation on this subject is given in Chapter4, section 4.1.2.1.

An oxide film will likewise grow rapidly onany solder preform exposed to air. To a firstapproximation,metallic oxide growth has a para-bolic relationship with time, following a stan-dard diffusion-type equation:

X2 � Dt � Dot e(�Q/RT)

where X is the layer thickness, D is the diffu-sivity, t is the time in seconds, Do is the diffusioncoefficient (m2/s), T is the temperature in Kelvin,Q is the activation energy (J/mol), and R is theuniversal gas constant (8.314 J/mol • K).

For pure tin, Q � 33 kJ/mol and Do�3.7 �10�18 m2/s, while for eutectic lead-tin solder therespective values are 40 kJ/mol and 2.5 � 10–17

m2/s. Thus, the native oxide on solid solder pre-

Fig. 3.11 Effect of oxide thickness on copper lands on thedefect level of joints incurred during PCB assembly

Fig. 3.12 Oxide growth on four base metals at room tem-perature, as a function of time


forms will be typically a few nm thick shortlyafter cleaning, but will rapidly increase in thick-ness, as can be seen in Fig. 3.14, and especiallyas the heating cycle progresses, unless the at-mosphere is either inert or actively prevents re-oxidation.

The dominant species of oxide that forms onsolder is usually consistent with that predictedbased on classical thermodynamics, namely thestableoxideof theconstituent element thathas thelowest free energy of oxide formation. This me-tallic constituent will tend to oxidize preferen-tially at the surface of the alloy preform. Thus,even on the high lead 95Pb-5Sn solder, the oxideskin contains somewhere in the regionof 20 timesmore tin than lead. The free energies of oxide for-mation of typical elements used in solders arelistedinTable3.1.Metalhydroxidesare thermallyunstable with respect to the oxides, so native sur-face films are usually not hydrated to any appre-ciable extent. From Table 3.1 it can be seen thatonly the stability of indium and zinc oxides ex-

ceeds that of tin. Hence the majority of solder al-loys are predominantly coveredwith a layer of tinoxide and only indium- and zinc-containing al-loys have different surface oxides.This is also thereason why indium- and zinc-base solders eachhave their own unique flux formulations.

The tin oxide growth onmolten lead-tin solderfollows a square-root time dependence, whereasthe combination of oxides that form on In-48Snsolder follows a parabolic growth law, as shownin Fig. 3.15. Indiumoxide initially thickensmuchmore rapidly than tin oxide, particularly whenthe solder is molten and accounts for the relativedifficulty of obtaining satisfactory wetting andspreading using indium solders without flux. In-dium-tin solders oxidize to form a mixed oxideof In2O3 and SnO and so are somewhat easier touse than other indium solders, but still less sothan tin-base solders. Likewise, on gold-tin al-loys, the tin oxide is diluted through the incor-poration of gold. As gold does not oxidize, ex-cept in very unusual environments, for gold-tinsolder the rate of oxide growth is slightly slowerthan on pure tin.

3.3.2 Self-Dissolution of SolderOxides

Publishedwork indicates that if amoltenmetalor alloy is heated above some critical tempera-ture then dissolution of thin films of surface ox-ide into the bulk of the melt can occur, the extentdepending on the solubility of the oxide. If theheating cycle is conducted in a controlled atmo-sphere that prevents reoxidation of the free sur-face, then a fluxless soldering process may bepossible under these conditions.

Fig. 3.13 Oxide growth on three base metals as a functionof temperature. The time at temperature is of the

order of a fewminutes and has some correspondencewith typicalsoldering cycle times once component mass and heating rates aretaken into consideration

Fig. 3.14 Oxide thickness versus oxidation time for a range of solders held in air 140 °C (284 °F) above their melting point. Adaptedfrom Dong, Schwarz, and Roth [1977]


The superheats (temperatures above the melt-ing point) required to dissolve the surface oxideson seven common solder alloys, in nitrogen, havebeen determined experimentally and are given inTable 3.7 [Dong, Schwarz, andRoth 1977]. Inter-estingly, the high-tin-containing alloys performwell in this regard, comparedwith lead-tin eutec-tic solder, and this phenomenon might partly ex-plain the success being achieved in running lead-free soldering process at very low superheats,using Ag-Cu-Sn alloys with only mildly activefluxes. Solder-spreading tests suggest that thespreading of high-tin-containing solders has farhigher tolerance to theoxygen level in theprocessatmosphere (nitrogen) than other solders, asshown in Table 3.8 [Dong, Schwarz, and Roth1977], for the same reason.

The combined effect of the presence of surfaceoxide on both the substrate and solder can be seeninFig. 3.16,which shows the contact angle for In-48Sn solder on various substrates, after heatingfor 5min at 250 °C (480 °C), as a function of oxy-gen partial pressure.Because the quality of the at-mosphere used is nowhere near adequate to affectoxide reduction on the solder, the conclusion isthat spreading is being achieved through a com-bination of oxide dissolution and other physicalmechanisms that locally rupture the oxide filmsand facilitate sites of direct metal-to-metal con-tact from which spreading can proceed.

The ability of solders to dissolve their ownoxides has only a limited benefit for most flux-less soldering applications. While wetting andspreading characteristics may superficially ap-pear quite satisfactory, the degradation of thesolder resulting from oxide dissolution is mani-fested in increased viscosity and impaired me-chanical properties. As a process option it istherefore only really applicable to circumstances

where the joint contains a large volume of solder,a high ratio of the volume-to-surface area of thefiller metal, and large superheats can be used.

3.3.3 Reduction of SolderOxides by Hydrogen

For the reasons that are explained in section3.1.3.2, whether or not an oxide can theoreticallybe reduced by hydrogen depends on the stabilityof theoxide at theprocess temperature and thehy-drogen/water partial pressure ratio in the solder-ing atmosphere. Once these fundamental condi-tions are satisfied, one then needs to consider theoxide reduction rate. Even when an oxide is ther-modynamically unstable, the rate of transforma-tion of the surface oxides might be so slow as torender the process practically ineffective.

Inanormalsolderingprocess,onemightexpectthe initial layersofoxideon thesurfaceof thefillermetal tobeof theorderof, for example, 2nmthick(0.8 μin.) and the process cycle time available nomore than 1min. In this case, for reduction by hy-drogen to be effective, the dynamics of oxide re-movalhas toexceed2nm/min(0.8μin./min).Thismeans that either the quality of the atmospheremust exceed the thermodynamic minimum valueand/or the temperature must be increased abovethe base point. From measurement of the reduc-tion rate of solder oxides as a functionof tempera-ture in hydrogen gas containing about 10 ppmcombined oxygen and water vapor, it is possibleto define a threshold temperature abovewhich thereduction rate exceeds 2 nm/min (0.8 μin./min).As canbe seen fromFig. 3.17, for solders coveredwith tin oxide the threshold temperature is ap-proximately 430 °C (800 °F) and for solders cov-ered with indium oxide it is slightly higher at 470°C (880 °F). The initiation temperature for the re-duction of zinc oxides exceeds 500 °C (930 °F),at which point the volatility of zinc will start tocause other problems. Hence, hydrogen gas is

Fig. 3.15 Oxide growth on molten solders at an oxygen par-tial pressure of 1 Pa (1.5 � 10–4 psi). Adapted from

Kuhmann et al. [1998]

Table 3.7 Superheats required to dissolvenative oxides on selected solders

Superheat for dissolving oxides(a)

Solder °C °F

Sb-95Sn 10 18Cu-99Sn 18 32Ag-96Sn 19 34Pb-63Sn 77 139In-48Sn 93 167Sn-9Zn >300 >540

(a) Excess temperature above the melting point


not an effective substitute for flux unless the sol-dering process can be undertaken at very high ex-cess temperatures.

In an investigation of the oxidation and re-duction kinetics of selected lead-tin, gold-tin,and indium-tin eutectic solders, it was possibleto use hydrogen to achieve self-alignment ofsolder bumps during flip-chip interconnection[Kuhmann et al. 1998] (further information onflip-chip interconnection can be found in Chap-ter 5, section 5.2). However, in order to reachthis satisfactory result it was necessary to achievehydrogen/water ratios in excess of the limitingvalues derived from the Ellingham diagram.The investigators satisfied the requisite condi-tions by pumping down the soldering oven,which was an ultrahigh vacuum (UHV) system,to a pressure of 10–9 Pa (1.5 � 10–13 psi). Thispressure will have defined the residual partialpressure of water vapor present. Then, the intro-duction of sufficiently pure hydrogen at a pres-sure 1 Pa (1.5 � 10–4 psi) was able to satisfy thethermodynamic requirements for reduction ofoxides, including that of indium, at a suffi-ciently fast rate that the oxides were actuallyremoved within a few minutes at normal pro-cess temperatures. While the ability to achievethe stringent conditions required for the reduc-tion of solder oxides by hydrogen has beendemonstrated in a laboratory context, it is doubt-ful whether this would be realistic or economicin a manufacturing environment.

3.3.4 Reduction of SolderOxides by Atomic Hydrogen

Although molecular hydrogen is not practica-bly capable of reducing solder oxides at normal

soldering temperatures, the same does not applyto atomic hydrogen. Atomic hydrogen is farmore reactive than the molecular species. Inessence, the energy provided externally to splitthe diatomic hydrogen molecule into single at-oms then becomes available for accomplishingchemical work. Atomic hydrogen reacts withsurface oxides on solders and joint surfaces toform hydroxides, hydrogenated complexes, orwater, all of which are volatile at typical solder-ing temperatures and can therefore easily beremoved from the area of the joint. Molecularhydrogen can be dissociated into atomic hydro-gen using a variety of methods including aheated filament, photo dissociation, electricaldischarge, microwave radiation, and, probablythe most practical, low-frequency alternatingcurrent ionization. Commercial atomic hydro-gen soldering systems are available from atleast one vendor [ATV 2003].

The concentration of any monatomic hydro-gen produced decays rapidly with distance fromthe source so that it must be generated near theworkpiece. However, because atomic hydrogenis such a reactive species toward solder oxide, acommercial argon/hydrogen gas mixture can beused to feed the plasma generator and at a hy-drogen concentration that is sufficiently low toobviate the need for hydrogen-monitoring equip-ment and the associated safety features.

Because the species in a plasma are not inthermodynamic equilibrium, application ofchemical thermodynamics to the analysis of sucha system requires modifications. A quantitativeprocess description has been formulated that hasallowed a line corresponding to the followingchemical reaction to be added to the Ellinghamdiagram [Jacob, Chandran, and Mallya 2000]:

Table 3.8 Tolerance of molten solders to the oxygen level in the atmosphere, as indicated by theexcess temperature above the melting point necessary to achieve solder spreading

Superheat, °C above melting point

SolderO2 level10 ppm

O2 level100 ppm

O2 level1,000 ppm

O2 level10,000 ppm

Sb-95Sn 6 15 18 No spread<300 °C (<570 °F)

Cu-99Sn 3 7 18 No spread<300 °C (<570 °F)

Ag-96Sn 9 17 19 No spread<300 °C (<570 °F)

Pb-63Sn 22 24 87 No spread<300 °C (<570 °F)

In-48Sn 83 No spread<300 °C (<570 °F)

. . . . . .

Sn-9Zn No spread<300 °C (<570 °F)

. . . . . . . . .


4H � O2 � 2H2O

Under standard conditions, monatomic hydro-gen can reduce almost all metal oxides at tem-peratures below 500 °C (930 °F). Currently, it isnot possible to produce monatomic hydrogen atatmospheric pressure (Po � 105 Pa, or 15 psi),and the dashed line included in Fig. 3.18 repre-sents the reduction potential of atomic hydrogenat a base pressure of 1 kPa (0.15 psi). To exploitfully the reduction potential of monatomic hy-drogen, it is necessary to remove the products ofreaction and prevent back-reaction resulting fromthe formation of molecular hydrogen and oxide.This is achieved using a flow system where thehydrogen gas is passed through a dissociator andthen over theworkpiece before being transportedaway.

Using monatomic hydrogen, filler alloy andcomponent surfaces can be cleaned in situ andexcellent quality fluxless joints obtained at nor-mal soldering temperatures. The process cycletime can be less than 1 min, the time being lim-itedmore by themechanics associated with pass-ing the part through a heated zone containing themonatomic hydrogen rather than by the kineticsof the reduction process itself. Figure 3.19 is anexample of application of hydrogen plasma tolead-tin solder using a commercial solder reflowoven equippedwith a low-frequency plasma gen-erator. Similar findings are reported in the lit-erature [Hong and Kang 2001].

Reduction of metal oxide by hydrogen plasmahas been followed by direct measurement of theoxygen/silver ratio as a function of process time(see Fig. 3.20). The substrate was 40 nm (1.6μin.) thick silver, the base pressure 1 Pa (1.5 �10–4 psi), and the plasma intensity 1 eV. Thekinetic energy of the hydrogen ions was belowthe threshold for sputtering processes, verifyingthe chemical nature of the reductive mechanism[Bielmann et al. 2002].

3.3.5 Mechanical Removal ofOxides (Ultrasonic Soldering)

The native oxides on the surfaces of the solderand components can, of course, be removedphysically. Although an obvious form of me-chanical abrasion is to use abrasive paper or ascalpel, more refined and controlled methods ex-ist. For example, fast atom bombardment usinga neutral beam can mill surfaces at rates of tens

Fig. 3.16 Contact angle for In-48Sn solder on four metalsubstrates, after heating for 5 min at 250 °C (482

°F), as a function of oxygen partial pressure. Adapted from Preuss,Adolphi, and Werner [1994]

Fig. 3.17 Reduction rate of solder oxides in hydrogen, containing 10 ppm total oxygen and water vapor, for a range of solders asa function of process temperature. Adapted from Dong, Schwarz, and Roth [1977]


of nm per minute. The effectiveness of the pro-cess is demonstrated in Fig. 3.21, which refers toflip-chip joints made with high-lead Pb-3Sn sol-der. In this graph, joint strength is held to be anindicator of surface oxide thickness. The simi-

larity of form between Fig. 3.21 and Fig. 1.27(showing the effect of mechanical wiping) isnoteworthy as they are essentially the samephysical process. It is reported that an additionof 10% hydrogen into the sputtering plasma con-siderably boosts the oxide-removal rate. It islikely that the hydrogen ions contribute to theoxide removal through a chemical mechanism,as discussed in the preceding section [Hong,Kang, and Jung 2002].

The main drawback of all mechanical meth-ods of cleaning surfaces, which is illustrated inFig. 3.22, is that the oxide films promptly regrowif the parts are exposed to air or, indeed, left ina nitrogen atmosphere of even the highest qualityfor more than a few minutes. Therefore, a flux-less soldering process, not involving an atmo-sphere containing atomic hydrogen, must be ableto cope with a thin layer of native oxide on, atleast, the solder. Strategies for making a robustfluxless soldering process are given in the fol-lowing sections.

An automated version of physical abrasion toremove surface contamination is ultrasonic agi-tation, which is more correctly called ultrasonictinning. This technique is highly effective in re-moving even the most tenacious oxides and istherefore applicable to virtually all metals, but ismostwidely usedwith aluminum [Jones andTho-mas 1956]. It is not a new technology, and thereexists a patent reference to an ultrasonic solder-ing iron that predates World War II [Antonevich1976]. Ultrasonic fluxing normally involves ap-plying an ultrasonically activated soldering ironto tin the workpiece [Schaffer et al. 1962]. In avariant arrangement, a solder bath is ultrasoni-cally excited and the workpiece is dip coatedwith the solder. These ultrasonic systems typi-cally operate at 20 to 80 kHz and 0 to 300 W ofelectrical power.Most systems operate at a singlefrequency, but some provide for application ofhigh- and low-frequency acoustic energy simul-taneously. A more recent development is to in-corporate an ultrasonic transducer into wave-soldering equipment. The ultrasonic sonodes arean integral part of the solder wave and impartsonic action while the board travels through thewave [Swanson 1995]. In operation the sonodesremove the oxide from the component leads andPCB lands, thereby enhancing wetting of thesolder. The oxide removal is accomplished bythe cavitation action of the ultrasonic waves onthe oxidized surfaces. The system operates witha frequency in the 20 kHz range and an ampli-tude in the micron (sub-mil) range. These pa-

Fig. 3.18 Simplified Ellingham diagram for oxides with aline (dashed) on it representing the reduction po-

tential of monotomic hydrogen at a partial pressure of 1 kPa (POis defined as this pressure as it is not possible to generate mona-tomic hydrogen at atmospheric pressure by anymethod currentlyavailable.)

Fig. 3.19 Balls of lead-tin solder, approximately 100 μm di-ameter, reflowed using an atomic hydrogen

plasma. The solidified surface is exceptionally smooth and shinyand, at highmagnification the duplex phasemicrostructure soldercan be readily seen, attesting to the surface cleanliness. Source:Agilent Technologies


rameters are selected in order not to adverselyaffect the integrity of the components and circuitboard material.

Ultrasonic fluxing is a fluxing and tinning op-eration for applying a metal coating to a cleansurface of the workpiece. Once the metal coatinghas been applied, it protects the workpiece sur-face against reoxidation, and for that reason isakin to chemical fluxing, as described previ-ously. The coated part can subsequently be sol-dered (or brazed) using a conventional process.At present, the method is limited to tinning ofpassive electronic and optical devices. Activedevices generally do not respond well to expo-sure to high-intensity acoustic waves.

Vianco and Hoskin [1992] have studied thewetting of copper by molten tin when assisted bya point source of ultrasonic energy (20 kHz, 20–70 W). Some of their results are reproduced inFig. 3.23. Visual inspection of ultrasonicallytinned substrates revealed tin films on both thefront and back surfaces of the copper coupons.This implies wetting is not solely due to directmechanical erosion by the incident pressurewave, but also secondary propagation within thesubstrate and molten metal.

Ultrasonic fluxing is attractive as a combinedmethod of fluxing and tinning since it is fluxlessand residue-free. With the progressive improve-ments being made to the understanding of theprocess and the push toward “greener andcleaner” manufacturing, ultrasonic soldering isreentering more widespread use, including for

applying metallizations to refractory materialssuch as alumina [Naka and Okamoto 1989].

3.3.6 Reactive Gas Atmospheres forReduction of Oxides

Working within the definition of a “fluxless”process being “one that does not use a solid orliquid flux,” it is possible to include the optionof using reactive gases to chemically removesurface oxide films. A number of gases are suf-ficiently active to be capable of reducing evenindium oxide. Reactive atmospheres that havebeen examined and reported in the literature in-clude formic acid and acetic acid vapor, carbonmonoxide, and halogen gases (CF2Cl2, CF4, SF6,etc.) [Moskowitz, Yeh, and Ray 1986]. Formicand acetic acid are probably the most commer-cialized reactive gas for fluxless soldering. Car-bon monoxide and halogen gases require rela-tively high temperatures for the reduction of tinoxides in any reasonable timescale, for example,350 °C (660 °F) maintained for 3 to 5 min, whichprecludes their use in many soldering processes.In this regard, these gases are similar to hydrogengas.

Formic acid vapor is active at typical solder-ing temperatures, and the process implementa-tion is relatively simple. Its effectiveness can bejudged from Fig. 3.24, which illustrates the im-proved self-alignment of flip-chip assemblies asformic acid is introduced at progressively higherconcentrations, to the point of reducing surface

Fig. 3.20 Reduction of silver oxide by atomic hydrogen generated in proximity to the compound using a plasma. The energy ofthe plasma was deliberately chosen to be below the threshold at which oxide removal could be attributed to sputtering

processes.


oxide on the solder spheres [Lin and Lee 1999;Deshmukh et al. 1993].

The reaction between gaseous formic acid andmetal oxide can be described by:

MO � 2HCOOH � M(COOH)2 � H2O(where M is the metal)

This reaction occurs at a significant rate at tem-peratures above about 150 °C (300 °F) and istherefore effective at stripping surface oxide filmsprior to the solder melting. When the tempera-ture is higher than about 200 °C (400 °F), themetal formate decomposes to liberate carbon di-oxide and hydrogen:

M(COOH)2 � M � CO2 � H2

The concentration of formic acid vapor in theatmosphere needs to exceed 0.5% in order tohave a useful practical effect. Appropriate pre-cautions need to be exercised with this process,as many organic metal compounds, especially

those of lead, are toxic. Appropriately specifiedmaterials are required for furnacing equipmentcapable of utilizing formic acid vapor, owing toits corrosiveness, but the price premium appearsto be relatively small and off-the-shelf equip-ment designed to work with it can be readilypurchased.

Acetic acid vapor is effective at reducing tinoxide on solder. This is evident by the differencein the time to wetting for Pb-62Sn solder oncopper and gold-coated nickel (see Fig. 3.25).Short wetting times are always obtained on theclean gold/nickel metallization, but longer pro-cess times or higher concentrations of acid arenecessary to clean bare copper and obtain similarrapidity of wetting [Frear and Keicher 1992].

3.3.7 Surface Conditioning ProcessesAs an alternative to cleaning the faying sur-

faces on the components and the filler metal insitu, processes have been developed that convertthe surface layers of oxide to stable compoundsthat protect against reoxidation and that arereadily removed during the soldering cycle.

One such process exploits fluorine chemistry.This is the PADS (plasma-assisted dry soldering)process developed at The University of NorthCarolina. It uses a plasma-assisted reaction todissociate CF4 or SF6 within an appropriate car-rier gas to produce monatomic fluorine, which,it is claimed, converts tin oxide to tin oxyfluor-ide:

SnOx � yF � SnOxFy

The reaction time is a few minutes in a re-duced-pressure atmosphere, and the componentFig. 3.21 Effect of fast atom cleaning on the strength of high

melting point Pb-3Sn solder joints made withoutbreaking vacuumbetween cleaning and soldering. Adapted from:Kohono et al. [1996].

Fig. 3.22 Effect on joint strength of deliberate delay, in ahigh-vacuum environment, between mechanical

cleaning and assembly for high-lead solder (Pb-3Sn) flip-chipbonds

Fig. 3.23 Wetted area of copper coupons by molten tin at245 °C (473 °F) as a function of ultrasonic power

at 20 kHz and a process time of 30 s. The degree of wetting wasalso found to be directly proportional to the process time for theprocess temperature of 245 °C (473 °F), but the relationship wasless clear at other test temperatures. Adapted from Vianco andHoskin [1992]


temperature need not exceed 50 °C (122 °F).Treated parts can be stored in air and on subse-quent heating in a nonoxidizing atmosphere thetin fluoride effectively volatilizes, leaving cleanmetal surfaces that are readily soldered. The sol-derability shelf life forparts treated in thismanneris up to 1 week when stored in normal laboratoryair or 2 weeks in a nitrogen-purged desiccator[Marczi, Bandyopadhay, and Adams 1990].

Although pH-buffered chemical treatments areavailable commercially that are capable of modi-fying oxide films on copper, few are availablethat work well with solders or other base metals.One innovation that appears able to overcomethis limitation is the ROSA (reduced-oxide sol-dering activation) process (Rockwell Scientific).In this process, an acidic solution containing va-nadium ions and associated hydrogen is used toreduce oxides. The vanadium ions, which alsosupply hydrogen ions by virtue of their multiplevalency states, are regenerated electrochemi-cally in a closed loop, so the only effluent isoxygen gas. This process only removes surface

oxides, and therefore the treated componentsmust be used before the oxide substantially re-grows [Tench et al. 1995].

3.3.8 Fluxless SolderingProcesses Considerations

While the methods described in the precedingsections may be used to minimize the inhibitingeffect of surface oxides on soldering, even whenthey are applied in combination it is normally notpossible to make a fluxless soldered joint of thesame quality as a fluxed joint. The reason issimply that these agents are not as pervasive atreaching the critical interfaces in the assemblyand are not as active as a liquid chemical inrendering all metal surfaces clean at the solder-ing temperature and simultaneously protectingthe cleaned surfaces from reoxidation. Addi-tional means are therefore necessary to encour-age wetting and spreading of molten filler metalsduring fluxless processing. Generally, suchmeth-

Fig. 3.24 Effect of formic acid vapor concentration (0.35 to 1.7%) on the pull-in alignment of flip-chip components deliberatelymisaligned by 30 μm (1.2 mil). Adapted from Deshmukh et al. [1993]

Fig. 3.25 Fluxless wetting of Pb-62Sn solder on copper and gold-on-nickel metallizations, using acetic acid vapor of varyingconcentration as an active reductant in the atmosphere


ods involve utilizing nonoxidizable metalliza-tions on the parent materials, minimizing thesurface-area-to-volume ratio of the filler metal,applying mechanical means of enhancing solderflow and, sometimes, metallurgical modificationof the contact angle.

3.3.8.1 Solderable Component Surfaces

Because oxide films grow so rapidly on mostbase metals, as shown in Fig. 3.13, in practiceclean and solderable component surfaces canonly be achieved if they are of gold. Accord-ingly, for fluxless soldering, a gold coating isgenerally applied that will be of adequate thick-ness and sufficiently pore-free to ensure goodsolderability over its specified storage life. It isimportant to bear in mind the fact that thesolderable shelf life afforded by a gold coatingwill be a function of the roughness of the under-lying surface, the method of application of thecoating and its thickness (see Chapter 4, section4.1.2). Thus, a 0.5 μm (20 μin.) thick gold layerdeposited by sputtering can be relied on tomaintain excellent solderability for severalmonths, even where the coated surface is rela-tively rough. On the other hand, the solderabil-ity of a gold layer of the same thickness, butapplied by electroplating to a rough surface (Ra> 3 μm, or 120 μin.) may not offer adequateprotection to an underlying base metal fromatmospheric oxidation for more than a few days.

The use of gold-coated surfaces imposes con-straints on the solders that can be used. Inparticular, most tin-base solders, including lead-tin eutectic, are largely incompatible with thickgold metallizations, on account of the high solu-bility of tin in gold, which results in the forma-tion of AuSn4 and consequential embrittlementof the joints if this phase becomes dominant.This topic is discussed in further detail in Chap-ter 2, section 2.3.2. This restriction can only beovercome by applying high-quality gold coat-ings of less than a certain thickness to the jointsurfaces, so as to prevent the formation ofAuSn4 as the primary phase. Where thicker goldcoatings are used, fluxless soldering tends to beconfined to high-gold and indium-base solders,which do not form catastrophic embrittlingphases with gold.

The necessity for a noble metal surface onthe parent materials can be eliminated if thesolder itself is applied as the barrier coating to

previously cleaned component surfaces. Thisapproach is considered in the next section.

3.3.8.2 Preform Geometry

The form in which solder is admitted into ajoint gap can make a profound difference to thesuccess of fluxed and especially fluxless solder-ing processes. Notwithstanding the condition ofthe faying surfaces, it is a general rule that thegreater the solder volume is in relation to itsexposed surface area, then the more readily theprocess will work. This is simply because thereis proportionally less oxide to impede wettingand spreading. It is therefore perhaps not sur-prising that manufacturers of solder paste go toquite considerable lengths to ensure that thesolder balls used in its preparation are perfectlyspherical, have exceptionally high surfacesmoothness (low Ra), and that the most com-mon form of solder is round wire.

The most appropriate geometry of preformfor admitting solder into a joint gap depends onthe shape of the components being joined. Ide-ally, the solder preforms should be designed tohave not only a high volume-to-surface-arearatio but also to be orientated so that the ad-vancing front ofmolten solderwill sweep trappedgas out of the joint gap (see Chapter 4, section4.3.1.1).

In situations where it is desired to makejoints with very low aspect ratio—that is, thinin relation to the plan area—it is often notpossible to achieve sufficient solder spread toreliably obtain complete joint filling. This sce-nario is encountered commonly in the micro-electronics and photonics industries where thereis a need to join planar components but withextremely narrow joint gaps on account of therelatively poor thermal conductivity of mostsolder alloys compared with many other metals.The thinnest solder preforms that can be eco-nomically purchased are 15 μm (0.6 mil) thick.Handling these foils is almost an art form, andmechanically cleaning them is virtually impos-sible. The high surface-area-to-volume ratio ofsuch preforms also runs counter to the need tominimize native oxides.

A growing number of companies are nowoffering a solution to this problem in the formof substrates that are sold with the solder com-position of choice preapplied to surfaces. Thesolder is usually applied by either electroplating


or a vapor-phase technique, through a mask, sothat only the required areas of the substrate arecoated. The solder thicknesses available rangefrom 2 to 50 μm (0.08 to 2 mil), and almostevery common composition is available.

These solder coated substrates offer a numberof distinct advantages compared with a solderfoil or wire:

• Piece-part inventory and number of suppliersare reduced by dispensing with preforms.

• Jigging is likely also to be simpler.• The thickness of the solder joint is decreased

because the solder layer can be substantiallythinner than the minimum practicable thick-ness of approximately 25 μm (1 mil) requiredfor a solder preform.

• Soldering behavior is improved by eliminat-ing two joint surfaces with all the attendantproblems, including oxide layers, from thejoint gap.

• Solder spread is automatically confined to aprecise area, and the responsibility of sub-strate wettability is also passed on to the sub-strate producer.

The quality of substrates prepared in this man-ner has improved substantially, and, althoughthere is a price premium compared to traditionalfoil, they are now available to aerospace andtelecommunications qualified standards.

Because the success of a fluxless solderingprocess is critically reliant on the absence ofsurface oxide on the molten solder, it is goodpractice to test whether this condition is beingachieved. This may be readily accomplished ifthe furnace has a viewing port. On heating tothe process temperature, a source of clean sol-der will melt and its surface acquires a shiny,mirrorlike finish. This is often referred to as the“liquid lake” condition. Any defects in the liq-uid solder film, such as texture, a gray bloom,or brown spots indicate that some element ofthe process, usually either the solder itself orthe reflow atmosphere, is in some way deficientand must be corrected.

3.3.8.3 Mechanically Enhanced SolderFlow

No matter what precautions are taken, in anindustrial fluxless soldering process it will al-ways be the case that by the time the componentsand preform have been set in jigs, loaded into theprotective atmosphere, and heated to the processtemperature, then the solder will be totally en-

cased in a skin of oxide. If it is possible to extrudevirgin metal through fissures or other defects inthe solder, then there is an improved prospect ofachieving a sound joint. One method of doingthis is to apply a compressive force to the jointgap.

The effectiveness of this approach is illus-trated in Fig. 3.26. Clearly, the higher the com-pressive pressure, the more effective it is, withthe optimal load in the region of 10 g/mm2 (14psi), or more. A loading such as this is relativelyeasy to achieve with weights or spring-loadedjigs for all but the largest components. Precau-tions need to be taken to ensure that the load isapplied uniformly and parallel to the joint gap,and that the method of application does not posea thermal sink on the assembly that would giverise to adverse temperature gradients.

Figure 3.27 shows application of this ap-proach to fluxless soldering of a GaAs semicon-ductor die to a gold thick-film metallized alu-mina substrate. The back of the semiconductor ismetallized with gold, and the filler metal is apreformofAg-96Sn solder, 15 μm (0.6mil) thick.The process superheat was 30 °C (86 °F). X-radiography reveals the joint to be free of voids(the black circles are blind vias in the GaAs,which are included for functional reasons).

3.3.8.4 Metallugically Enhanced SolderFlow

Occasional reports appear in the published lit-erature of the significant difference that low con-centrations of other metals can make in promot-ing wetting and spreading of molten solders.

Fig. 3.26 Void level (as a percentage of the plan area of thejoint) versus the load applied to preforms of three

solders during fluxless processing. The soldering conditions usedfor the three solders were a superheat of 25 °C (or 77 °F) and drynitrogen as the joining atmosphere. The joints measured approxi-mately 10 mm � 10 mm � 25 μm (0.4 in. � 0.4 in. � 1 mil).


Some cases are cited in Chapter 2, section 2.2and Chapter 5, section 5.8. Virtually all fluxlessaluminum brazes make use of the fact that ppmlevels of bismuth and beryllium have a markedeffect on the wetting and spreading characteris-tics of the Al-12Si eutectic alloy. Further infor-mation on fluxless aluminum brazing is to befound in the planned companion volume Prin-ciples of Brazing.

The authors have investigated improving thefluidity of indium by minor additions of otherelements with a view to enhancing the use ofindium filler metals in fluxless processes. Figure3.28 shows the wetting angle as a function ofheating time for pellets of indium solders of con-trolled weight and geometry wetted on to silversubstrates at 200 °C (392 °F) in a vacuum of 1mPa (1.5 � 10–7 psi). The measurements weretaken from video stills of the substrate viewededge-on, with the timeline beginning at the onsetof observed melting. Of the additions, zinc hasan adverse effect. Antimony is initially neutral,but appears to impede any additional wetting.Bismuth and gold are beneficial additions thatimprove wettability, which requires an extendeddwell at the soldering temperature to take effect,while the rare earth metal cerium produces amarked and immediate reduction in contactangle. This could be a fruitful area for furtherresearch in developing new solder alloys (seeChapter 5, section 5.8).

3.3.9 Example of a Fluxless SolderingProcess Using In-48Sn Solder

The requirement was to improve the quality asoldered joint that was beingmade using a 25 μm(1mil) thick preform of the In-48Sn solder, with-

out resorting to the use of chemical fluxes. Thejoint measured approximately 10 by 5 mm (0.4by 0.2 in.), and the components were both met-allized with 3 to 5 μm (120 to 200 μin.) of nickeloverlaid with 3 μm (120 μin.) of gold. In theindustrial environment where this work was car-ried out, it was necessary to demonstrate thebeneficial effect of each change in the processbefore other alterations could be investigated.Shear strength was used to assess the quality ofthe soldered joints.

The first change to be considered was to applya load to the joint gap during the heating cycle.Application of a compressive stress helps ensurethat all free surfaces within the joint are in in-timate contact. This has the effect of minimizingtemperature gradients across the joints and, asexplained previously, promotes rupturing of theoxide skin on the solder when it melts.

The data shown in Fig. 3.29 (air) indicates therelationship between the applied load and result-ing joint strength, with the joint shear strength in-creasing by roughly 1 kg (2.2 lb) for every 100 g(0.2 lb) of appliedweight, for the range of appliedpressure shown (0 to 6 g/mm2, or 0 to 75 psi).

A more significant improvement to the pro-cess was made by carrying out the joining op-eration in a furnace maintained under dry nitro-gen, giving a combined oxygen and water vaporcontent of less than 10 ppm in the working at-mosphere. This measure provided some protec-tion against oxidation of the parts and solder anddelivered a stepwise improvement in jointstrength; see Fig. 3.29 (nitrogen). It was ob-

Fig. 3.27 Fluxless soldering of a GaAs monolithic micro-wave integrated circuit, approximately 3 � 5 mm

(0.12 � 0.12 in.) achieved by application of a compressive loadof 100 g/mm2 (140 psi) during the heating cycle. Source: BAESystems

Fig. 3.28 The effect of different doping additions on the flux-less wetting angle of indium on silver substrates, as

a function of time following the commencement of melting of thesolder


served that as a result of changing to the nitrogenatmosphere, considerably less solder was exudedfrom the joint gap, implying that the increase injoint strength is largely due to improved wettingof the component surfaces.

Hitherto, the foil preforms of solder had beenused as received from the supplier. These arecovered with a significant layer of native oxide.Mechanical abrasion of the preform surfaceswitha glass-fiber brush immediately prior to use ledto a further improvement in joint strength, as canbe seen in Fig. 3.29 (cleaned).

A thin foil solder preform is not a particularlyattractive method of admitting solder into a jointbecause:

• The oxide surface area-to-solder volume ra-tio is unfavorable.

• The foil is difficult to handle and clean if thin.• There is a high risk of trapping pockets of

vapor when the parts and foil are jigged to-gether (see Chapter 4, section 4.3.1.1).

Thin foils also command a cost premium, espe-cially when they have to be purchased to customdimensions.

Solder in the form of round wire avoids thesedrawbacks. The preform was prepared from twoequal lengths of wire, cold compression weldedso as to form a symmetrical cross, as shown inFig. 3.30. The surface area-to-volume ratio ofthis cruciform configuration is much lower thanfor a foil, and wiping the wire several times witha paper tissue soaked in solvent is an easymethodof striping the solder oxide (see Fig. 1.27). Thebeneficial effect of substantially removing theoxide film from the solder by these means isattested by the results presented in Fig. 3.29

(wire). The cruciform configuration also pro-vides formechanical stability during jigging, andthe direction of solder flow after melting helpssweep out any trapped gas from the joint gap (seeChapter 4, section 4.3.1.1).

Examination of the joint surfaces of disas-sembled parts revealed good wetting of the com-ponents and a continuous bead of shiny solderformed at the edge of the joint.

3.3.10 Fluxless Soldering ofAluminum

From the preceding discussion it might be as-sumed that to attempt fluxless soldering of alu-minum would be ineffectual. However, fluxlesssoldering of aluminum is practiced as a repairtechnique. It was developed duringWorldWar IIas a method of patching small bullet holes inairplane skins. The original process simply usedzinc as the joining material, but the joints wereweak and susceptible to corrosion. The modernvariation employs a filler alloy containing about90% Zn, 7% Al, with the balance being magne-sium, manganese, and other elements. The alloymelts at approximately 380 °C (716 °F), andtherefore it qualifies as a solder. The joiningmethod involves heating the aluminum part, of-ten with a standard blowtorch and puddling (rub-bing) the solder until a wetted surface film isdeveloped over the area to be joined. No flux isnecessary because the alumina is physically re-moved and the exposed aluminum overcoatedwith a more oxidation-resistant alloy. Both jointsurfaces are prepared in this manner, placed in

Fig. 3.29 Shear strength of joints approximately 10 � 5 mm(0.4 � 0.2 in.) in area made fluxless using In-48Sn

solder at process temperature of 150 °C (302 °F) to gold-metallized components, as a function of the applied compressiveload, showing also the effect of atmosphere quality and conditionof the solder (see text for details)

Fig. 3.30 Cross-shaped preform of In-48Sn solder preparedby coldwelding of two 300 μm (12mil) diamwires

at the common intersection. Source: BAE Systems


contact and the joint made by reheating withoutfurther addition of the fillermetal [Phillips 1994].Additional mechanical agitation helps meld thetwo molten liquid skins together. The aluminumin the solder helps to prevent corrosion of thejoint, the zinc addition lowers the melting point,while the minor elements assist wetting andspreading. It is reported that joints made in thismanner can have shear strengths around 90% ofthat of the parent metal.

If the layer of applied zinc-base solder can bemade extremely thin, it is possible to diffuse itcompletely into the base material if the processcycle is suitablyextended.This technique isprop-erly known as diffusion soldering, but it does en-able fluxless, effectively corrosion-resistantjoints to be made to aluminum, which are not asourceofmechanicalweakness[Ricksetal1989].The applicability of the approach is limited by thelongheating cycle times and theneed to either useprecisely machined parts or apply pressure toforce the abutting surfaces into intimate contact.This latter requirement arises from the very lim-itedvolumeofzinc in the joint andhence thesmallquantity of liquid filler metal available to fill thejoint gap.Adiscussionof diffusion soldering, is tobe found in Chapter 5, section 5.9 and diffusionbrazing in the planned companion volume Prin-ciples of Brazing.

Appendix A3.1:ThermodynamicEquilibrium and theBoundary Conditions forSpontaneous ChemicalReaction

The thermodynamic function that provides ameasure of the driving force of a chemical (in-cluding metallurgical) reaction is the Gibbs freeenergy, which is defined as:

G � E � PV � TS (Eq A3.1)

where E is the internal energy, S the entropy, Tthe absolute temperature, P the pressure, and Vthe volume of the materials system. The defini-

tion and physical meaning of internal energy andentropy are explained below. As is shown, animportant property of theGibbs free-energy func-tion is that it is always aminimum at equilibrium,and the extent of its departure from theminimumvalue provides a measure of the tendency of areaction to proceed spontaneously—that is, ofthe driving force for the reaction.

The First Law ofThermodynamics andInternal Energy

The subject of thermodynamics addresses en-ergy changes in systems. In thermodynamics, theterm “system” is used to describe a set of ma-terials that are capable of undergoing a change—as, for example, through a chemical reaction.

The First Law of Thermodynamics is a state-ment of the Principle of Conservation of Energy.The various statements of this law are bound upwith the differentiation of various types of en-ergy and, in particular, with the concept of in-ternal energy. The internal energy of a systemmay be considered as the aggregate of the kineticenergies and energies of interaction (i.e., poten-tial energies) of the atoms andmolecules ofwhichthe constituent materials are composed. Whenthe system is isolated from its surroundings, sothat no exchange of energy can take place, thenits internal energy remains fixed. However, ifmechanical work can be done on the system, butno heat is exchanged with its surroundings (i.e.,the system is adiabatically isolated)—for ex-ample, by an impellor stirring a liquid or gas inan insulated container—its internal energy, E,will change by an incremental amount equal tothe work, W, performed:

dE � �dW (adiabatic)

The minus sign denotes that work is done on thesystem to raise its internal energy. This expres-sion provides a thermodynamic definition of in-ternal energy.

The internal energy of a system depends onlyon the state of the system (defined in terms ofmacroscopic or thermodynamic properties, suchas the pressure and temperature of the system).For this reason, internal energy is termed a func-tion of state.


Where work is done in changing the volumeof a chemical system by an increment dV throughthe application of external pressure P, then:

dW � PdV

dE � �PdV

In practice, most systems are not totally insulatedfrom their surroundings so that thermal energymay be exchanged between them.

If an increment of work dW is done on thesystem and an increment of heat dQ is exchangedwith the surroundings, then the internal energydE will change by the amount:

dE � dQ � dW

This equation is a mathematical expression ofthe First Law of Thermodynamics which, for achemical system, may be written:

dE � dQ � PdV (Eq A3.2)

Entropy and the SecondLaw of Thermodynamics

Internal energy alone cannot determine theequilibrium state of a system. Although when asystem reaches a state of equilibrium the internalenergy achieves a fixed value, this may not be aminimum. For example, the internal energy willincrease when a solid melts at constant tempera-ture and pressure through the absorption of latentheat. For this reason, in addition to the internalenergy, it is necessary to stipulate the value ofanother state function of the system—namelyentropy—which, together with the internal en-ergy, measures the extent to which the system isremoved from equilibrium.

The concept of entropy arises in connectionwith the conversion of heat intomechanical workand vice versa. The Second Law of Thermody-namics defines the conditions under which thisconversion from one form of energy to anothercan occur. The Kelvin-Planck statement of thislaw relates to a device that can perform work byextracting heat from a particular source and per-forming an equivalent amount of work, withoutany other energy exchange with the surround-

ings. It follows that a reciprocating engine thatoperates by extracting heat from one sourcemustreject some of this heat to a sink at a lowertemperature. If the operating cycle of the engineis reversible, such that work can be performed topump heat from the sink back to the source, it ispossible to show that in accordance with theSecond Law, the integrated ratio dQ/T over onecomplete cycle is zero:

R� dQ

T� 0

The circle through the integral sign denotes thatthe integration is to be carried out over the com-plete cycle, and the letter R is a reminder that theequation applies only if the cycle is reversible.This result is known as Clausius’ theorem.

If the integration is carried out over only partof the cycle, say between two states 1 and 2, thenthe integrated ratio dQ/T is not zero, but equalsthe difference between the values of a thermo-dynamic function at the two states:

R�1

2 dQ

T� S1 � S2

This thermodynamic function of state is calledentropy. If the two states are infinitesimally close,then the relationship can be written:

(dQR

T)R

� dS (Eq A3.3)

Subscript R indicates that this equation onlyholds if the heat increment dQ is transferred re-versibly. This equation provides a mathematicalexpression of the Second Law of Thermody-namics.

A consequence of the fact that entropy is afunction only of state, a system that has changedfrom state 1 to state 2, always has entropy S2,which differs from that of the initial state S1 byS1,2 � S2 – S1, irrespective of the means used todrive the system. Thus, for example, the systemmay have been set in motion, and some of thekinetic energy converted into heat in overcomingfrictional forces, thereby raising its temperatureto a value that takes the system from state 1 to


state 2. In this irreversible process, the energywas not supplied to the system as heat so that:

I�1

2 dQ

T� 0

The letter I denotes that the process is irrevers-ible.

However, the entropy change S1,2 is still thesame as that obtained by a reversible changebetween states 1 and 2, because it depends onlyon these states and not on the process connectingthem; that is, here too S1,2 � S2 – S1. Thus, in allirreversible processes, the entropy change isgreater than (�dQ/T), where dQ is the heat ab-sorbed at each incremental step in the irrevers-ible change. This result can be generalized to thestatement that in a spontaneous irreversiblechange, the entropy of an isolated system willincrease, and when in equilibrium, it will remainconstant.

Considering the system and its surroundingstogether (i.e., the universe), any kind of processcan be represented in entropy terms by:

dS (universe) 0

Therefore, from a thermodynamic viewpoint,which is macroscopic, entropy can be under-stood as the propensity of a system to undergoa change, such as a chemical reaction. Aclearer physical picture of entropy can be ob-tained at the microscopic level, where a sys-tem may be regarded as an ensemble of atomsor molecules. On this basis, it can be shownthat entropy provides a measure of the degreeof atomic or molecular disorder that exists inthe system, and this will always tend to in-crease. This concept is consistent with the ob-servations that all metals are intersoluble, al-beit in some cases only to a small extent, andthat all liquid metals will wet the clean sur-faces of solid metals.

The Dependence ofGibbs Free Energy on Pressure

Having defined the thermodynamic func-tions internal energy, E, and entropy, S, and ex-plained their physical significance, it is pos-sible to demonstrate the significance of the

Gibbs free-energy function, G, to determine thetemperatures and pressures under which chemi-cal reactions are thermodynamically favorable,as well as the direction of the reactions.

In incremental form, Eq A3.1 can be written:

dG � dE � PdV � VdP � TdS � SdT

Substituting for dE and TdS from Eq A3.2 andA3.3 (all chemical/metallurgical processes beingreversible) gives:

dG � dQ � PdV � PdV � VdP � dQ � SdT

� VdP � SdT

For a reversible process at constant tempera-ture (isothermal) and constant pressure (iso-baric), that is, when the system is in equilibrium:

dG � 0

and G is constant and has a minimum value. Thisis an important result for metallurgical reactions,because these can be considered as taking placeusually at constant temperature and pressure.

More generally, at constant pressure, dP � 0,and then:

dG � �S dT

and at constant temperature, dT � 0, so that:

dG � V dP

If the system is an ideal gas, the Gas Law:

PV � nRT

applies, where n is the number of moles of gasand R is the gas constant. Then:

dG � nRTdP/P at constant temperature

so that the Gibbs free-energy change resultingfrom a change from state 1 to state 2 at constanttemperature is:


G2 � G1 � nRT ln P2/P1

The Gibbs free energy, like any other measureof energy, must have some reference point. Byconvention, a zero value of G is assigned to thestable form of elements at 25 °C (77 °K) and 1atm of pressure. Then the Gibbs free-energychange of a gas at constant temperature from itsvalue Go at atmospheric pressure, which is de-fined as its standard state value, is given by:

G � Go � nRT ln P (Eq A3.4)

where P is the pressure corresponding to thefree-energy state G, expressed in atmospheres.

Although Eq A3.4 is strictly valid for idealgases, it is also approximately applicable to realgases and can be used for them at pressures closeto normal atmospheric pressure (1 atm).

In the case of solids, the molar volumes aresmall compared with those of gases, so that thechange in the Gibbs free energy of solids result-ing from small pressure excursions, �P, suchthat �P � 1 atm (100 kPa) at constant tem-perature is small and, to a first approximation,may be neglected in reactions involving solidsand gases. It is also assumed that the solubilityof the gaseous species in the solid phases is neg-ligible at the temperatures of interest, as is largelythe case in practice.

It is now possible to determine the pressuredependence of the Gibbs free energy of the re-agents that participate in a chemical reaction.Consider a reaction involving four gases, A, B,C, and D and two solids, X and Y, all at constanttemperature T, as follows:

xX � aA � bB ↔ yY � cC � dD

where a, b, c, d, x, y are the number of moles ofeach of the reagents. The gaseous reagents areassumed to behave as though they are ideal gases.

The Gibbs free energies G(X) and G(Y) of thesolid constituents at moderate pressures are ap-proximately equal to their values at atmosphericpressure, as explained previously. Therefore:

Free energy of x moles of solid X � xG(X)

� xGo(X)

Free energy of y moles of solid Y � yG(Y)

� yGo(Y)

The Gibbs free energies of the gaseous constitu-ents are:

For a moles of gas A:aG(A) � aGo(A) � aRT ln P(A)

For b moles of gas B:bG(B) � bGo(B) � bRT ln P(B)

For c moles of gas C:cG(C) � cGo(C) � cRT ln P(C)

For d moles of gas D:dG(D) � dGo(D) � dRT ln P(D)

where G(A), G(B), etc. are the Gibbs free en-ergies of 1 mole of the reagents A, B, etc. atpressures P(A), P(B), etc., and Go(A), Go(B),etc. are the corresponding values at 1 atm.

The free-energy change for the reaction is,from Eq A3.4:

�G � G(products) � G(reactants)

� cGo(C) � dGo(D) � aGo(A) � bGo(B)

� RT ln�P(C)�c �P(D)�d

�P(A)�a �P(B)�b

� �Go � RT ln�P(C)�c �P(D)�d


The Gibbs free-energy changes of the solid re-agents can be neglected, for the reasons givenpreviously.

Under equilibriumconditions, temperature andthe respective pressures P(A), P(B), and so forth,are constant, and:

�G � 0


Hence, the Gibbs free-energy change when thegaseous reactantsAand B in their standard statesare transformed to the products C and D in theirstandard states may be expressed in terms of thepartial pressures of the respective reactants inequilibrium, thus:

�Go � �RT ln�P(C)�c �P(D)�d

�P(A)�a �P(B)�b(Eq A3.5)

Since the Gibbs free-energy change �Go, for aparticular reaction at a fixed temperature and atatmospheric pressure has a fixed value, so toodoes the argument of the logarithm. This con-stant is called the equilibrium constant KP, be-cause it can be used to determine the equilibriumstate that a reacting system will attain.

KP ��P(C)�c �P(D)�d


The subscript P denotes that the equilibrium con-stant is specified in terms of pressure.

Equation A3.5 becomes:

�Go � �RT ln KP

For an oxidizing reaction described by:

xM � �y/2�O2 ↔ MxOy

there is one gaseous constituent and two solids,so that the equilibrium constant is simply:

KP �1

�PO2

M� y/2

where PO2

M is the partial pressure of oxygen re-quired to effect the oxidation reaction, or thedissociation pressure of the oxide, and

�Go�RT ln PO2

M (Eq A3.6)

per mole of oxygen participating in the reaction.That is, the driving force needed to oxidize ametal, as expressed by the Gibbs free-energychange, is directly related to the oxygen partialpressure of the atmosphere according to EqA3.6.

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CHAPTER 4

The Role of Materials inDefining Process Constraints

THIS CHAPTER CONSIDERS the materialsand processing aspects of soldering and the man-ner in which these interrelate in the developmentof joining processes.

The starting point of any practical joiningprocess development is a need to fabricate aunitary assembly or product from a set of com-ponents. Often, the components are of differentmaterials in order to maximize the performanceof the product for a given cost. The productitself will have been designed to satisfy certainfunctional requirements, and, for some items,these can be very diverse. The joint propertiesmust also be consistent with the specified pur-pose and application of the product or system.

To identify and develop an appropriate join-ing process, it is essential to first give consider-ation to all possible aspects associated with theproduct. This will reveal an array of constraints,some of which are not immediately obvious,that govern the feasibility of the joining route.Aflow chart outlining the decision-making stepsand the constraints at each stage of designingand manufacturing assemblies containing sol-dered joints is given in Table 4.1. Those con-straints that are directly linked to the productitself include the cost tolerance of the product tothe joining process, the scale and throughput ofproduction that will have to be satisfied, and thestatutory regulations that apply. The operatingenvironment must then be considered. Here, thepeak temperature, stress condition, and the cor-rosion environment tend to be the critical pa-rameters. Finally, the materials used in the com-ponents, when taken individually, will imposean upper limit on the maximum joining tem-perature that can be used, on the grounds ofthermal degradation, and may restrict the atmo-

sphere in which the joining process can becarried out. When the overall assembly is con-sidered, any mismatch in thermal expansivity ofthe abutting components can force compro-mises with regard to the choice of materials andprocesses. All of these considerations will influ-ence the design of the assembly to a greater orlesser extent.

Having taken account of the more obviousconstraints in the design of an assembly, a se-lection of filler alloys can be made, each ofwhich will impose its own set of limiting con-ditions. Among the most important of these arethe minimum practicable joining temperature(i.e., liquidus temperature of the filler alloy,with the addition of a margin to allow for pos-sible temperature gradients across the joint), thegeometries that the joint can assume and intowhich the filler can be fabricated, and the per-missible joining atmosphere. From the “short-list” of filler alloys, it is usually possible toselect at least one whose window of usableconditions is compatible with the other steps ofmanufacture. Up to this point, the selectionprocedure is largely a paper exercise, but theviability of the proposed joining solution needsto be established by practical trials. This isbecause a multiplicity of other features alsoenters the equation, such as wetting, erosion ofthe parent materials, and intermetallic phaseformation within the joint. While any of thesephenomena can radically affect the integrity ofthe product in service, much of this type ofinformation is unavailable from the publishedliterature and cannot be correctly surmised.

If problems are identified at this stage of thejoining process development, it may be possibleto obtain remedies by a number of avenues.



Thus, in certain situations, catastrophic mis-match stresses may be overcome by modifyingthe stress distribution in the vicinity of the joint.Problems associated with the formation of del-eterious intermetallic phases in the joint, lack ofwetting, and, at the other extreme, excessiveerosion of one or more of the parent materialscan usually be circumvented by interposing alayer of a different metal between the filler and

the parent material and thereby altering themetallurgical constitution of the joints. Fluxesand active filler alloys can be used to improvewetting and reduce void levels. Some of thepossible remedies that may be brought to bearon these and other problems are detailed in thefollowing sections.

If, on the other hand, no joining solutionproves technically tractable, or if the solutions

Table 4.1 Materials systems approach to joining process developmentPhase of development Decision making Boundary conditions

Nature of the product andfunctional requirements

Scale of productionCost constraintsSize and weight limitsStatutory requirements

Definition of the product �

Service conditions Thermal environmentStress environmentChemical environment

�

Parent materials Joining temperatureJoining atmosphereMismatch stress

Materials and processselection �

Filler alloys Joining temperatureJoining atmosphereJoint geometryFiller geometry and

form�

Process assessment:identification of criticalmaterials problems

Determination ofprocess viability

Metallurgicalconstraints

� �

Mismatch Alloying

� � �

Wetting Erosion Phases

��

(Mechanical solutions) (Metallurgical solutions)Identification and

achievement of solutions �

Interlayers Surface processingMultilayers � Wettable coatingsGraded structures Barrier coatingsCompliant structures Active filler alloysDiffusion soldering/brazing Fluxes

Diffusion soldering/brazing

�

Prototyping and production Process specified Tolerance of processestablished

Adapted from Principles of Soldering and Brazing, 1993, p 112


are not economically justifiable, then changesearlier in the decision-making chain are re-quired. In extreme situations, this may requiredrastic revision, perhaps even to the extent thatother means of assembly have to be consideredor that the functional requirements of the prod-uct must be relaxed.

4.1 MetallurgicalConstraints and Solutions

In principle, most metals can be joined usingfiller alloys. However, when there is a require-ment to join two different parent materials to-gether, the available choice of fillers that arecompatible with both is narrowed somewhat,especially when the constraints on the solder,processingconditions, andpropertiesof the jointsmentioned previously are imposed. The prob-lem is often made more acute by the fact thatthe joining processes tend to be left to a laterstage of product design, which further reducesthe available options. The joints are alwaysconsidered as an integral part of the overalldesign of an assembly if manufacturing is to befacilitated.

Metallurgical incompatibility of materials andprocessingconditionswillmanifest itself throughpoor wetting, excessive erosion of the parent ma-terials, and/or the formation of undesirablephases. Means for eliminating or suppressingthese deleterious characteristics are described asfollows.

4.1.1 Wetting of Metals by SoldersRestrictions applied to the choice of joining

temperature and atmosphere, including the useof fluxes, can result in poor wetting of the com-ponent surfaces by the molten filler if the per-missible process conditions are inadequate to en-sure that the joint surfaces are sufficiently clean.Active solders (described in section 4.1.2.2 ofthis chapter) can often help to overcome thisproblem, but it is seldom possible to use solderscontaining active wetting agents. Alternativeremedies are then necessary.

Poor wetting is a particularly serious problemwhen the parent materials are refractory metals.Their reactivity with oxygen, and, in some cases,with other elements in the atmosphere, and thestability of the reaction products on the surfacesof these materials cause poor wetting. Nonme-

tallic phases present at the surface of materials,such as graphite inclusions in cast iron and non-metallic components in metal-matrix compos-ites, can also inhibit wetting.

These problems are not restricted to the parentmetals but can also encompass the filler metal. Aprime example is provided by zinc-base solderswhen these need to be used without an appro-priate fluxing agent. Foils and other preforms ofthese alloys tend to produce poor wetting andspreading over the joint surfaces. However, bymaking small additions of elements that lowerthe surface tension of the molten filler and thatalso help to destabilize the native surface oxidelayer, wetting can be considerably improved.This is illustrated for indium in Chapter 3, Fig.3.28, and other examples are outlined in Chapter2, section 2.2. The concentration of the indi-vidual additions must be restricted to a maxi-mum of approximately 1% in order to avoid per-ceptibly altering the bulk metallurgicalcharacteristics of the filler alloy.

If the joining environment, which includesfluxes when these are approved, cannot ad-equately clean and protect the components fromoxidation during the process cycle, a favoredsolution is to apply a coating of a more noblemetal to the precleaned joint surfaces. This coat-ing may be sacrificial in that it is subsequentlydissolved by the filler, which then wets the cleansurfaces of the parent material.At the same time,the composition and thickness of the coatingshould be such that when dissolved, it does notgive rise to brittle phases by reaction with thefiller and parent materials.

Copper, silver, and gold are the principal el-ements used as wettable metallizations becauseof their nobility, metallurgical compatibility withmost solder alloys, and ease of deposition. Tin isalso widely used in conjunction with solderingprocesses, because tin oxide is readily displacedby an advancingwave of themolten solder, whichitself often contains tin. The benefits of tin arethat it is relatively inexpensive and that it pre-sents minimal risk of unfavorably altering theconstitution of the solder.

In the semiconductor and optoelectronic in-dustries, platinum overlaid with a gold flash is afrequent choice of wettable metallization. Thin-film deposition equipment with a platinum sput-tering source is often readily available, and thematerials cost is insignificant compared to theprocess cost. Platinum has the benefit of excep-tionally low solubility in tin- and indium-basesolders and will not oxidize when exposed to air

Chapter 4: The Role of Materials in Defining Process Constraints / 147

(Table 4.2). The gold flash ensures good spread-ing and aids visual inspection of the faying sur-faces prior to use. The reaction product withtin-base solder has been identified as an excep-tionally thin interfacial layer of PtSn4 [Kuhmann1977]. This metallization scheme confers excel-lent wetting and spreading with precisely con-trolled reproducibility, which are useful at-tributes when joining high-value components.

Solderable metallizations are usually appliedby either wet plating or vapor deposition tech-niques. Tin coatings can be applied by dippingmethods. The choice of deposition method is de-terminedbyvarious factors, such as the sizeof thecomponent, its geometry, the required scale ofproduction, the capital and running costs of thedeposition equipment, and the thickness of themetallization.Additional details of themerits andlimitationsof eachof these coating techniques aregiven in Appendix A4.1 in this chapter.

Any metallization applied needs to be suffi-ciently thick for it to protect the underlying ma-terial from corrosion and to maintain wettabilityof the component over a reasonable storage pe-riod. The uniformity and density of sputter-deposited coatings of gold afford such protectionat a thickness of typically 0.3 μm (12 μin.),whereas electroplated coatings need to be morethan an order of magnitude thicker to provide thesame shelf life [Humpston and Jacobson 1990].This point is illustrated in Fig. 4.1, which showsthe wettability, in terms of the wetting forcemea-sured on a wetting balance, of chromium-metallized coupons coated with different thick-nesses of gold by sputtering and electroplating,as a function of the storage time in an open at-mosphere. The porosity of thin electroplated orevaporated coatings is frequently overlooked, es-pecially when a process is transferred from de-velopment to a less rigorous manufacturing en-vironment. Further proof of the ineffectivenessof noble metal coatings in preventing oxidationof an underlying base metal comes from the ben-eficial exploitation of their use to grow oxidefilms of controlled thickness in the manufacture

of solar cells. In particular, 5 μm (200 μin.) ofsilver, applied by thermal evaporation, to a cleancopper surface permits the growth of Cu2O at thecommon interface at a rate of roughly 0.1 μm/h(4 μin./h) at 500 °C (930 °F) [Rosenstock andRiess 2000].

The alloying behavior of the solder with thecoating metal may establish a maximum limit tothe thickness of the metallization beyond whichjoint embrittlement ensues, as explained in sec-tions 4.1.3 and 4.1.4 of this chapter. Where eco-nomic and size factors dictate that electroplatingis the preferred method for applying gold coat-ings, the components should be provided withtypically 3 to 4 μm (120 to 160 μin.) thick plat-ings of gold to confer an adequate solderableshelf life. Then, immediately prior to the sol-dering operation, the level of gold can be con-siderably reduced to suit, in particular, high-tinsolders by wicking this metal off the joint sur-faces. The removed gold can be reclaimed fromthe discarded solder.

Metallizations that promote wetting can alsobe used to advantage to confine themolten solderto specific areas on the surfaces of components.By selectively applying the coating to prescribedareas, the solder spread can be restricted accord-ingly.

If the wettable coating is not adherent as ap-plied, it often can be secured on the surface of ametal component by a subsequent heat treatmentin a nonoxidizing atmosphere. This process canpromote interdiffusion between themetal and the

Fig. 4.1 Solderability shelf life of gold-coated components.Thicker and denser coatings are more impervious to

oxygen and water vapor and therefore confer greater protectionto the underlying metal.

Table 4.2 Dissolution rate of platinum intin-base solder at selected temperatures

Temperature

°C °F Dissolution rate

80 180 (solid-state aging) 1.0 nm/day200 390 0.3 nm/s270 520 1.2 nm/s320 610 2.2 nm/s


coating, resulting in a graded interface akin tothat obtained using a carburizing (carbon-enriched) or sheradizing (zinc-enriched) surfacetreatment. For example, titanium is rendered sol-derable by coatingwith gold (as applied, the goldlayer will brush off if touched) and then heattreating at 750 °C (1380 °F) for 30 min in anitrogen atmosphere.

Copper lands on printed circuit boards andelectronic component leads are often coated soas to maintain solderability in storage and topermit extremely rapid wetting and spreading bythe molten solder at the appropriate time. Speedis important, because printed circuit boards passthrough the solder wave of a wave-soldering ma-chine in 2 s or less. The coatings must also en-train low materials and application costs. Thereare basically three main types: fusible coatings(tin, lead-tin, and other tin alloys); soluble coat-ings, that is, noble metals that are readily solublein molten solder (silver, gold); and organic coat-ings (1,2,3-benzotriazole, 1,3-benzodiazole, imi-dazole, and similar compounds). Often, thesecoating strategies are combined. For example,one commercial product system designed tomaximize the preservation of the solderability ofcopper involves application of a thin (0.07 to 0.1μm, or 3 to 4 μin.) layer of silver, which is itselfprotected from damaging environmental effectsby an organic layer, typically also 0.1 μm (4 μin.)thick. Silver is a good choice of overmetal be-cause it is more noble than copper, and unlikelead/tin coatings, it cannot form an intermetallicstructure with copper, which generally has poorsolderability. The organic coatings are designedto form a chemical bond to a clean substratesurface, and hence protect it during storage, andthen either dissolve in flux or decompose cleanlyjust below the soldering temperature (typically120 to 150 °C, or 250 to 300 °F). Nickel/gold andnickel/palladium finishes are more correctlytermed barrier coatings, because the nickel andpalladium protect against reaction between thesolder and the substrate. The gold overcoat pre-serves the solderability of the nickel or palla-dium. A comprehensive review of surface fin-ishes relevant to electronic circuit assembly onprinted circuit boards is provided by Vianco[1998].

4.1.2 Wetting ofNonmetals by Solders

Nonmetals—namely, ceramics, glasses, andplastics—are not wetted by most solders, even

when their surfaces are scrupulously clean. Thisis because they are chemically very stable, withtheir atoms strongly bound to one another. There-fore, these materials will not react with and bewetted bymolten solder unless the latter containsan active element, which can attach itself to theanionic species of the nonmetallic material. In acompound or complex, this is normally oxygen,carbon, nitrogen, or a halide element.

There are two solutions to soldering to ceram-ics.Oneis touseanactivefillermetal,asdiscussedin section 4.1.2.2 of this chapter. Active metaljoining is only effective if sufficiently high tem-peratures, typically above 800 °C (1470 °F), canbeused for the joiningoperation, so that the activeingredient is able to react with the nonmetal.Where activated fillers produce successful joints,it is observed that the active element concentratesat the interfacewith the nonmetallic basematerialand impartsmetallic characteristics to the surfaceof the ceramic.

Where this approach is incompatible with thejoining process, a very similar result can beachieved via a different route. This involves coat-ing the joint surfaces with a metal that will bondstrongly to the underlying nonmetal and that is atoncewettedbythefillerwhilenotentirelydissolv-ing in the process. Any erosion through to theoriginal component surface would result in de-wetting. This is exemplified by the dissolution ofchromium metallizations by bismuth-containingsolders observed at temperatures above 261 °C(502 °F). Bismuth and chromium react to form alow-melting-point eutectic alloy at this tempera-ture [Humpston and Jacobson 1990]. Further de-tails of solderable coatings for nonmetals aregiven in the following section.

4.1.2.1 SolderableCoatings on Nonmetals

Wettable coatings can be applied to the non-metallic components bymethods similar to thoseused on metals—namely, physical vapor depo-sition, chemical vapor deposition, and wet plat-ing. Also widely used are fired-on glass fritsloaded with particles of metal powder or flake,often referred to in the literature as thick-filmmetallization techniques.

The vapor phase deposition route tends to befavored wherever thin coatings of high qualityare required. It is versatile and permits a widerange of metal and alloy coatings to be applied.When metallizing a nonmetallic material, it isusual to deposit more than one layer onto the


joint surface. It is essential that the layer in directcontact with the nonmetal (henceforth referred toas the foundation layer) is an active metal, whichis usually nickel, chromium, or titanium. Thechoice is largely dictated by the solubility of thismetal in the filler alloy, which must be low butfinite, and the ability of these active elements tobond strongly to nonmetallic materials, providedthat the surfaces are clean.According to a simplethermodynamic model, it might be expected thatthe adhesion of a metal to, say, an oxide is givenby the Gibbs free energy of formation of theoxide of the metal: The greater the extent towhich the free energy is reduced when the metalreacts to form an oxide, the more active is thebonding. That thismodel is an oversimplificationof reality is clear from the observation that chro-mium forms an adherent bond to silica (quartz),even though the Gibbs free energy of formationof silica is higher in magnitude than that ofchrome oxide. Evidently, the Gibbs free energyof oxide formation is only one factor. This ex-ample merely highlights the complexity of thesubject of metal-to-nonmetal bonding and ac-counts for the considerable literature that it hasproduced [Peteves 1988]. Nevertheless, the de-sired end result is that the component surface isrendered sufficiently metallic to enable wettingby the filler metal.

A sputtering process usually incorporates acapability for cleaning surfaces prior to deposi-tion, by being operated in a reverse-bias mode,so that the atoms of the inert carrier gas in thedeposition chamber are made to bombard thesurface and physically remove films of contami-nant. A separate ion source is often included inevaporation systems to perform the same func-tion. An in situ preclean stage is essential toremove absorbed species and achieve good ad-hesion by the foundation metal.

Chromium, or a mixture of nickel with chro-mium (Nichrome), is frequently recommendedas the reactive metal for the foundation layer onnonmetals [Holloway 1980]. Nichrome has alower intrinsic stress than pure chromium, whichis beneficial for some strain-sensitive electronicand optical components. Chromium reacts withmany nonmetallic compounds to form complexchromates. These are not only strongly bondedto the nonmetal but also act as a barrier to furtherinteraction taking place by diffusion betweenchromium and the nonmetal [Mattox 1973].Hence, it is possible to use chromium as a high-integrity metallization on glass and intrinsicallystable ceramics such as alumina and quartz. Ten-

sile strengths are reported to be in the region of70 MPa (10,000 psi) [Vianco, Sifford, andRomero 1997]. The benefit of this type of chemi-cal bonding is evident when metallizing rela-tively unstable ceramics, such as zinc oxide. Zincoxide is widely used in electrical voltage surgesuppressors, known as varistors [Leite, Varela,and Longo 1992; Wersing 1992]. Heating a zincoxide ceramic metallized with titanium or zir-conium causes the metallization to blister andspall off due to the oxidizing reaction:

2ZnO � Ti � TiO2 � 2Zn

�G2ZNO�Ti (573 K) � �272 kJ (�65 kcal)

A by-product of this reaction is the volatil-ization of free zinc from beneath the coating assoon as the coated part is heated, which resultsin the detachment of the metallization. On theother hand, a chromium metallization is entirelybenign toward zinc oxide, even upon heating to400 °C (750 °F) for prolonged periods, due to thepresence of the intervening chromate barrierlayer. Chromium is also marginally easier to ap-ply than titanium, because it is slightly less re-active toward oxygen in the air, and hence, thequality of the vacuum chamber in which thedeposition is conducted can, accordingly, beslightly relaxed.

The high reactivity of titanium and other ac-tive metals toward oxygen and nitrogen in airmeans that they rapidly lose theirwettability evenafter a brief exposure to the atmosphere. To over-come this problem, a more noble metallizationthat offers good wettability to filler metals mustbe deposited over the active-metal foundationlayer immediately, while the deposition chamberis maintained under a protective (low-oxygen)atmosphere.Ametallization system that has beenfound to be effective for many soldering appli-cations is one that comprises a 0.1 μm (4 μin.)thick layer of a foundation metal and a 0.3 μm(12 μin.) overlay of awettablemetallization, suchas gold, both applied by sputtering. The chosenfoundation-metal thickness is specified to pro-vide good step coverage even after taking intoaccount parameter variations in an industrial pro-cess that may result in local thickness variationsup to �25%. The thicker gold layer provides areasonable shelf life and usually permits joints tobe made using tin-base solders without risk ofembrittlement (see Chapter 2, section 2.3.2).


An alternative method of coating a nonme-tallic material with an adherent metal coating isto use a two-stage joining process in which thefirst step is to allow an activated solder or braze,capable of providing a solderable surface, to wetand spread over the component surfaces. The“tinned” surface is then usually mechanicallydressed to present a flat metal surface for theactual joining step, which is performed at a tem-perature below the solidus point of the activatedfiller metal. Because of the number of elementsthat are likely to be present in the joint, the re-sulting alloy constitution can be somewhat com-plex unless the process is designed such that thealloy used for the coating and the solder alloyhave constituents in common.

A somewhat different approach is to fire on arelatively thick (typically, 1 to 10 μm, or 40 to400 μin.) metal coating. This type of metalliza-tion process tends to be used only on glass andceramic materials because of the high tempera-tures involved [Bever 1986]. In very simpleterms, a thick-film paste can be considered as amixture of metal and ceramic or glass particles.On firing, the glass or ceramic phase wets thecomponent surface,while themetal particles floatto the surface. Adhesion between the solidifiedglass or ceramic phase and the metal particles isby a combination of chemical and physicalmechanisms. Further details of the range of com-mercial processes that are available are given inAppendix A4.1 in this chapter.

Reference may sometimes be found in oldertechnical literature to the active hydride process.It involves applying a metal hydride, usually oftitanium or zirconium, in the form of a paste tothe component surface. On heating, the hydridethermally decomposes, liberating hydrogen toleave a metal film on the component surface[Pershall 1949]. This method has lost favor inrecent years to alternatives, because the qualityof the resulting metallization tends to be vari-able, as reflected in the mechanical properties ofthe joints [Mizuhara and Mally 1985]. The un-predictable nature of the metallization quality isdue to its high sensitivity to variations in theatmosphere during the firing stage. Oxygen andwater vapor contents, in particular, affect the ex-tent to which the highly reactive metallic con-stituent oxidizes after decomposition of the hy-dride and hence the ease of wetting by themoltenfiller.

A widely used method for metallizing oxideceramics, in particular, alumina, is the so-calledmoly-manganese process. It has the advantages

of being relatively straightforward to perform,highly reproducible, not requiring a vacuum en-vironment, and thus amenable for processinglarge parts. In this process, a slurry of powdersof molybdenum, manganese, and various glass-forming compounds is applied as a paint to alu-mina components. Then, the coated ceramic isfired in a wet hydrogen atmosphere (with a well-controlled dewpoint) at a temperature of 1450 to1500 °C (2640 to 2730 °F). The firing operationresults in chemical reaction between the glassyphase in the alumina and the manganese, whilethe molybdenum is established as a surface layeron the ceramic [Mattox and Smith 1985]. Be-cause molybdenum is not readily solderable, es-pecially when the process is limited to mildfluxes, it is common to complete this metalliza-tion scheme with a coating of electroless nickel,followed by gold.

An important point to be aware of is that met-allizations are often put down in a stressed con-dition. This is an inherent feature of most depo-sition processes and can be a source of criticalweakness. An example of the manifestation ofexcessive stress in a metallization layer is shownin Fig. 4.2. Here, a silver layer 20 μm (800 μin.)thick has cracked away from a ceramic materialthrough such stress. The stress concentration atthe interface between the component and themet-allization can be controlled by limiting the thick-ness of the metallization layers, modifying thedeposition parameters, and using conventional,postdeposition, stress-relief heat treatments.

Particularly in microelectronics and photonicsmanufacturing, there is an ongoing trend towardproduct miniaturization. This is demanding the

Fig. 4.2 A porous ceramic material metallized with a thicksilver electroplate. The residual stress in the metal-

lization has resulted in a peel failure through the near-surfacelayer of the ceramic.


development of ever-smaller joint geometries andhence the volume of solder in each joint. At thesame time, pricing pressures mean that any ap-plied metallizations must be as thin as possible,because their cost is usually a direct function ofmachine time. In theory, a metallization needonly be sufficiently thick that it ensures completecoverage of the faying surface and, if it is abarrier metal, will not be eroded back to theunderlying material by reaction with the solderduring the joining process or in service. While itis possible to give general “rules of thumb” re-garding metallization thicknesses, attempts arenow being made to derive more rigorous valuesby calculation. One such approach uses a com-bined thermodynamic and diffusion-kineticmodel, and it is sufficiently broad in its scope tobe able to deal with real systems [Ronka, VanLoo, and Kivilahti 1998]. No doubt, further ad-vanceswill bemade in this area over the next fewyears.

4.1.2.2 Active Solders

Reactive filler metals are mostly brazes, andfurther details of their design and function can befound in the planned companion volume Prin-ciples of Brazing. The basis of the approach is toincorporate into the filler metal small quantitiesof elements that are highly reactive. Providedthat at least one of the products of reaction withthe base material is metallic and remains as alayer on the surface of the nonmetal, then thefiller alloy can wet and form sound joints.

When selecting a reactive filler alloy, thechoice of active ingredient, its optimal concen-tration, and the appropriate processing condi-tions need to be considered in relation to thenonmetal of the component. One of the mostcommonly used active constituents is titanium.Less reactive elements, such as chromium, andmore reactive elements, such as hafnium, arealso used.

The addition of metals such as titanium tosolders will enable them to directly wet non-metals such as ceramics, provided the processtemperature is sufficiently high to provide thenecessary activation for the chemical bond toform with the nonmetal. Typically, this activa-tion takes effect above 750 °C (1380 °F), whichsomewhat negates the benefit of the low pro-cessing temperature characteristic of solders[Xian and Si 1992]. The process temperature isgoverned by the reactivity of the active ingre-dient, which increases with temperature. If the

temperature margin to the liquidus temperatureof the solder is substantial, the heating cycleshould be kept short in order to prevent extensiveerosion of the joint surfaces or the growth ofthick interfacial phases.

It has recently been reported that meldingsmall quantities (0.5 to 2%) of lutetium into sol-ders has a similar effect but enables a reductionin the process temperature to more normal val-ues. Rare earth elements have poor solubility insolder and therefore tend to agglomerate in smallislands within the matrix, where they are pro-tected from oxidation by the surrounding metal.However, in themolten state, sufficient rare earthmetal is conveyed to the joint surface, where itforms a chemical bond with the nonmetal. Theinterface layer is 1 to 5 nm (0.04 to 0.2 μin.) thickand appears to comprise lutetium oxide (Lu2O3)and lutetium-rich solder. By this means, tensilejoint strengths in the region of 800MPa (120,000psi) and shear strengths of the order of 10 MPa(1500 psi) can be achieved to silica-based ma-terials [Ramirez, Mavoori, and Jin 2002]. Fur-ther information on solders containing rare earthelements can be found in Chapter 5, section 5.8.

Hitherto, active fillers were found only to beeffective above threshold temperatures of ap-proximately 750 °C (1380 °F), and therefore, theyhave only been used for brazing. Recently, aroute for substantially reducing the temperaturethreshold has been identified, making active sol-ders a reality. This involves the addition of lowconcentrations of gallium and lanthanide ele-ments to a silver-tin eutectic alloy containing 4%Ti [Smith 1998]. Although the exact role of theminor additions is unclear, they appear to fulfilltwo functions. First, they modify the nature ofthe surface oxide that limits its growth. Second,the minor additions appear to increase the ac-tivity of the titanium, which enables the solderalloy to wet and bond directly to nonmetals suchas alumina, aluminum nitride, beryllia, siliconcarbide, silicon nitride, boron nitride, carbon/carbon composites, aluminum/silicon carbidecomposites, copper, silicon, diamond, titaniumand stainless steel, to name but a few, using apeak process temperature of only 250 °C (480°F). The titanium and other additions increasethe strength of the bulk solder and, interestingly,confer it with significantly enhanced resistanceto corrosion from aqueous media containingchlorides. The activated solder can be used in airwithout flux, but in order to facilitate wetting, itis necessary to use some form of mechanicalactivation, such as scrubbing. This locally rup-


tures the oxide skin and enables wetting andspreading to proceed from points of contact be-tween the nonmetal and molten solder.

4.1.3 Erosion of Parent MaterialsWhen a molten filler wets the surface of a

parent material, alloying occurs, leading to a de-gree of dissolution of the parent material (ormetallization), commencing at the joint inter-face.

Dissolution of the parent materials occurs be-cause the materials system encompassing thejoint is not in thermodynamic equilibrium. Thisprovides the driving force forwetting and spread-ing. The maximum solubility of the parent ma-terials in the molten filler can be predicted byreference to the appropriate phase diagram, asstated in Chapter 2, section 2.3. In summary,extensive erosion is likely where the liquidussurface on the phase diagram between the filler-metal composition and that of the parent materialhas a shallow slope and where the alloying de-presses the melting point of the filler in the joint.If the phase diagram exhibits either of these fea-tures, then it is only possible to limit erosion bylowering the process temperature, shortening theheating cycle, and/or restricting the volume ofmolten filler metal. Such changes must obvi-ously not compromise the integrity of the jointsby, for example, reducing the effective fluidity ofthe molten filler alloy, which in turn will impedewetting, spreading, and joint filling.

Erosion can be reduced somewhat if interme-tallic phases form along the joint interface so asto attenuate the rate at which solid material istransported from the components into the moltenfiller, or vice versa. This approach is only suc-cessful if the intermetallic phases formed arereasonably ductile, as exemplified by the case oftin-base solders used in conjunction with silvermetallizations that is described in Chapter 2, sec-tion 2.1.2; otherwise, joint embrittlement results.In other cases, it is possible to protect the com-ponents against erosion by interposing a metal-

lization that will act as a barrier. An example isthe application of a layer of nickel, typically 2 to5 μm (80 to 200 μin.) thick, on aluminum com-ponents to prevent reaction with lead-tin solder.Although nickel-tin compounds form at the com-mon interface, in normal soldering heatingcycles, the intermetallic layer is quite thin andtherefore does not adversely affect the mechani-cal properties of the joints.

The extent to which dissolution occurs in agiven heating cycle will also depend on the ki-netics of reaction. The rate of dissolution of com-mon engineering materials and metallizations ineutectic lead-tin solder is given in Chapter 2, Fig.2.53. These data show that decreasing the pro-cess cycle time and temperature can be used toreduce erosion, although not always to accept-able levels in all cases.

The propensity for dissolution (erosion) of theparent metal by a filler alloy can be reduced bypreloading the filler with this metal so that it isalready saturated and, on becoming molten, willnot dissolve any further quantity. This approachis used for soldering to gallium arsenide devicesusing the Au-20Sn solder modified by the addi-tion of 3.4% Ga. The addition reduces the solu-bility of gallium in the solder by half, withconsequential benefits to the process yield[Humpston and Jacobson 1989].

Metallurgical reaction between a solder andthe materials of the components being joined (ortheir metallizations) can substantially alter thecharacteristics of the filler, which may in turnhave unexpected consequences for the resultingjoints. By anticipating the effects of these reac-tions, they either can be exploited to advantageor at least forestall embarrassing problems. Twoexamples are described as follows.

Changes to the Melting Point of Soft Solders.Gold is widely used as the surface layer in met-allization schemes because of its chemical in-ertness and ease of visual inspection. The changein the melting point (solidus temperature) of se-lected binary tin-base solders, when these havedissolved gold, is given in Table 4.3. The effect

Table 4.3 The effect of gold on the solidus temperature of common, binary tin-base eutecticsolders

Solder compositionwt% gold to form

AuSn4 as the primary phasewt% gold that will dissolvein solder at 50 °C superheat

Change in melting point of solderby dissolving 1% Au, °C (°F)

In-48Sn AuIn2 formed <1 �1 (30)Bi-43Sn >1 4 �2 (28)Pb-63Sn >8 13 �6 (21)Sb-95Sn >10 30 �15 (5)Ag-96Sn >11 30 �15 (5)


is particularly marked in the case of the Ag-97.5Pb-1Sn solder, where the solidus tempera-ture is reduced from 309 to 217 °C (588 to 423°F) when 4% Au is added. For a layer of thesolder 50 μm (2 mil) thick, this radical depres-sion of the melting point corresponds to a mere1 μm (40 μin.) of gold coating being dissolvedfrom each side of the joint. An example of achange in the contrary direction is the dissolutionof gold in the In-18Pb-70Sn solder (melting range136 to 182 °C, or 277 to 360 °F), where a 5%Auaddition raises the solidus temperature by 14 °C(57 °F). Partial isopleths through these quater-nary systems as a function of gold concentrationare given in Fig. 4.3 and 4.4.

Changes to Rate of Erosion of Parent Mate-rials. Minor changes to the composition of thefiller alloy are capable of influencing the rate atwhich dissolution occurs. This is exemplified bythe effect of small additions of silver, typically2%, on the rate of dissolution of silver by lead-tinsolders, as can be seen from Chapter 2, Fig. 2.3.The extent of the reduction could not be pre-dicted from the Ag-Pb-Sn phase diagram andwas determined empirically. Changes of this sortdo not significantly alter other properties of thefiller, or those of the resulting joints, and can bereadily implemented.

4.1.4 Phase FormationAlloying between a molten solder and parent

materials more often than not results in the for-mation of intermetallic compounds. This is par-ticularly more true of solders, because the majorconstituentsofmostsoldersareelementswith lowcrystallographic symmetry, and hence, they donot readily form solid solutions with engineeringmaterials that tend to be based on simple body-centered cubic (bcc), face-centered cubic (fcc), orhexagonal close-packed (hcp) crystal structures.The distribution, morphology, and proportion ofthese phaseswill dependon several factors, as ex-plained in Chapter 2, section 2.3.

Because intermetallic compounds generallypossess a higher elastic modulus (i.e., they arestiffer) than many filler alloys themselves, well-dispersed intermetallic phases are often benefi-cial to the stress-bearing capability of the fillermetal. This effect is shown for gold additions tothe silver-tin eutectic solder in Chapter 2, Fig.2.52.

By and large, agglomerations of intermetalliccompounds are deleterious to the mechanical

properties of joints. This is particularly truewhere the compound has low fracture toughnessand forms as a continuous interfacial layer be-tween the component surfaces and the filler al-loy. It is sometimes possible to restrict the coars-ening of these phases by carrying out the joiningoperation under conditions that are unfavorablefor their growth—namely, restricting the heatingcycle duration and peak temperature. In some

Fig. 4.3 Section through the Ag-Au-Pb-Sn quaternary systemshowing the effect on the melting range of adding

gold to Ag-97.5Pb-1Sn solder. Adapted from Evans and Prince[1982]

Fig. 4.4 Section through the Au-In-Pb-Sn quaternary systemshowing the effect on the melting range of adding

gold to In-18Pb-70Sn solder. (This composition is not a ternaryeutectic, as is sometimes stated in older literature.)


cases, minor additions can be made to the filleralloy that will break up agglomerations of in-termetallic phases present in the joint. This is oneof the reasons for some minor additions made tozinc-base solders formulated for soldering to cop-per, which is described in Chapter 2, section2.1.3. Barrier metallizations applied to joint sur-faces can be used to prevent alloying with theparent materials and the consequential formationof undesirable phases.

Some filler alloys are themselves hard becausethey contain intermetallic phases, yet form jointsof high strength with certain parent materials,provided that the intermetallic phases are finelydivided within the joint microstructure. Gold-tinsolders fall into this category, with rapid solidi-fication of the solder being a prerequisite forrobust joints. Further details are given in Chapter2, section 2.1.4.

4.1.5 Filler-Metal PartitioningPartitioning of filler metals is a fabrication

method that is becoming more common. It is anapproach that can be beneficial if an alloy com-position is relatively brittle when prepared viaconventional casting of ingots, but selected com-binations of the constituents are ductile. Someexamples taken from the published literature arelisted in Table 4.4. The objective is either one ofattempting to improve the mechanical propertiesof the filler metal prior to melting or one of costreduction through decreasing the number of stockalloys required [Mackay and Levine 1986]. It isusual to have the lower-melting-point combina-tion in the cladding, with a higher-melting-pointmixture of ingredients in the core. The low-melting-point fraction thenmelts at the solderingtemperature and wets the joint interfaces.

The Au-20Sn solder, as produced in the formof wire or foil, is very costly, partly because ofits precious metal content but mostly because ofthe limited reduction in thickness that can beachieved on successive passes during hot rolling.This means that the fabrication costs are high.

Because gold and tin are both very ductile met-als, composite foil can be prepared by cold roll-ing. The noble nature of gold is exploited, and,by placing it on the outside of the sandwich, itis then able to prevent oxidation of the under-lying tin during storage and heating to the pro-cess temperature. A postfabrication heat treat-ment step is claimed to be beneficial inconsolidating the foil and initiating interdiffu-sion between the constituents prior to use[Tokuriki Hoten 1981].

An alternative approach is to modify the sur-face of a ductile foil of tin through electrodepo-sition of a layer of gold. In this case, it is thehigher-melting-temperature constituent that con-stitutes the surface layer. If the tin foil is 25 μm(1 mil) thick, and 12 μm (0.5 mil) of gold platingis applied on both sides the calculated compo-sition of the solder will be hypereutectic at 70wt% Au, and 30 wt% Sn. From the phase dia-gram for the gold-tin system, it may be deducedthat if this composition were prepared as a ho-mogeneous alloy, it would have a melting rangeof 280 to 370 °C (536 to 698 °F).

The melting behavior of this alloy has beencharacterized using differential scanning calo-rimetry. The results are given in Fig. 4.5 and 4.6.On initial heating (Fig. 4.5), there is an endo-thermic reaction as the solid-state alloy formedby the interdiffusion of gold and tin melts at thebinary eutectic temperature of 217 °C (423 °F).This is followed by a tin-rich transition reactionat 252 °C (486 °F) before onset of the gold-richeutectic transformation at 280 °C (536 °F).There-after, melting continues until all the solid is con-sumed, which occurs at approximately 380 °C(716 °F). Cooling of this sample back to 150 °C(302 °F) and repeating the thermal analysis cycle(Fig. 4.6) shows that the sample is now homo-geneous, with a solidus temperature of 280 °C(536 °F) and a liquidus temperature of approxi-mately 380 °C (698 °F). This result accords wellwith the predicted melting behavior deducedfrom the alloy phase diagram.

Partitioning of solders can sometimes also bebeneficial where the maximum permissible pro-cess temperature is only marginally above themelting point of the filler metal. Phased reflowsoldering, as this approach is described, has beenexploited in someof the lead-free solderpaste for-mulations based on Bi-Cu-Sn alloys. These re-placement solders for lead-tin eutectic have tofunction at very low superheats, because mostprinted circuit boards and components have rated

Table 4.4 Examples of partitioned filler metalsFiller metal Cladding Core

Silver-tin (mpt, 221 °C) Sn AgLead-antimony-tin (mr, 245–280 °C) Sb-Sn Pb-Sb-SnSilver-lead-tin (mr, 303–310 °C) Ag-Sn Pb-SnGold-tin (mpt, 280 °C) Au Sn

mpt, melting point; mr, melting range


160.00 190.00 220.00 250.00 280.00 310.00 340.00 370.00 400.00

20.00

10.00

0.00

mca

l/s

Temperature, ˚C

Fig. 4.5 Melting behavior of gold-plated tin foil having an effective composition of Au-30Sn

Fig. 4.6 Remelting behavior of gold-plated tin foil


maximum process temperatures of 235 °C (455°F). Copper-tin eutectic melts at 227 °C (441 °F),and bismuth-tin at 139 °C (282 °F). By partition-ing the ternary composition so that it includessome low-melting-point tin-bismuth solder, liq-uid phase sintering will commence on heatingabove 139 °C (282 °F). The presence of the liquidphase improves heat transfer to the unmelted par-ticles and helps minimize temperature gradientsthrough the solder paste, both of which improvethe reliability of filler-metal joining processesperformed at low homologous temperatures. Theprogressive melting of the solder is also reportedto be beneficial in reducing the incidence of“tombstoning” [Warwick 2002]. This is whereone land on a double-ended surface mount com-ponentwetsmoreeffectively thanat theother, andthe imbalanceof surface tension forces causes thecomponent to lift up vertically on one land. Sev-eral solder manufacturers have filed patents onphased reflow soldering.

There have also been attempts to effect sol-dering by liquid phase sintering, that is, to joinmetal components without melting the higher-melting-point constituent of a partitioned filler-metal paste. By this means, it has been foundpossible to produce joints with reasonable me-chanical integrity but nowhere near as sound asthat achieved by melting the solder [Palmer, Al-exander, and Nguyen 1999].

4.2 MechanicalConstraints and Solutions

In an assembly composed of heterogeneousmaterials, there is usually a thermal expansionmismatch between the abutting components. Thismanifests itself as stress on cooling from thesolidus temperature of the filler metal and is amaximum at the lowest temperature that the as-sembly experiences. Materials with a relativelylow elastic modulus can accommodate strain andwill tend to deform under the influence of thisstress, while brittlematerials, notably glasses andceramics, have a tendency to fracture, particu-larly if the stress distribution places the compo-nent in tension. Even if a heterogeneous assem-bly survives the joining operation, the stressesarising from the thermal expansion mismatchcan cause it to fail by fatigue during subsequentthermal cycling in service.

The stress in the region of the joint betweentwo isotropic materials, designated 1 and 2, with

differing thermal expansivities that develops oncooling from the freezing temperature of the fillermetal can be approximated by the followingequation [Timoshenko 1925]:

Stress �E1 E2

E1 � E2

(X1 � X2)(Tf � Ts )

where E is the modulus of elasticity of materials1 and 2, X is the coefficient of thermal expansionof 1 and 2, Tf is the freezing point (solidus tem-perature) of the filler alloy, and Ts is the tem-perature of the assembly corresponding to thestress. In the derivation of this equation, it isassumed that the materials are only deformedwithin their elastic limits and that the joint isinfinitely thin. Despite these simplifying assump-tions and the inaccuracies that they introduce,this expression is useful in providing an indica-tion of whether the stress due to thermal expan-sion mismatch is close to or exceeds the failurestress of either of the abutting materials, that is,whether failure of the assembly is likely to occur.Some worked examples are described by Haug,Schaefer, and Schamm [1989].

In an assembly with a planar joint betweentwo elastic but different materials 1 and 2, themagnitude of the bow distortion in one dimen-sion can be estimated from the physical prop-erties of the materials using a simplified model[Timoshenko 1925]. With reference to Fig. 4.7.

Bow distortion, B � (L2/ 8) / R

Radius of curvature �

(A1 � A2)[3(1 � M)2 � (1 � MN)(M 2 � 1/MN)]

6(X1 � X2)(Tf � Ts)(1 � M)2

(Eq 4.1)

where

M � A1/A2 and N � E1/E2, B is bow distortionL is the length of the jointR is the radius of curvatureA1 and A2 are the thicknesses of materials 1 and 2X1 and X2 are the coefficients of thermal expan-sion of materials 1 and 2E1 and E2 are the moduli of elasticity of materials1 and 2Tf is the freezing point (solidus temperature) ofthe filler alloy andTs is the temperature of the assembly correspond-ing to the bow distortion


Equation 4.1 assumes that the joint in the hetero-geneous assembly is infinitely thin and totally in-elastic.

From Eq 4.1, it can be seen that it is possibleto effect some reduction in expansion mismatchstress, that is, decreasing R, by applying one ormore of the following measures:

• Decrease the solidus temperature of the filleralloy, Tf

• Increase theminimum service temperature ofthe assembly, Ts

• Reduce the dimensions of the joint area, L• Change one or bothmaterials tominimize the

mismatch in thermal expansivity, � E1 � E2 �

Occasionally, it may be possible to implementone or more of these changes, but they are likelyto conflictwith other processing constraints if notwith the intended functional requirements of theassembly.Alternative solutionsmust therefore besought.

In practice, it is usually possible to obtain asmall reduction in the distortion of a bowed het-erogeneous assembly by heat treating it at a tem-perature below the solidus temperature of thefiller alloy to enable stress relaxation and creepto occur in the filler metal. However, there is alimit to the reduction in distortion that can beobtained by this means, typically 10% or less.This stems from the fact that stress-reducingmechanisms are diffusion related and becomemore effective as the temperature is raised to-ward the melting point of the joint, while mis-match stress increases as the temperature of theassembly is reduced below that at which stressrelief is effective. Thus, these two tendencies act

in opposition, and the optimal condition for re-ducing the distortion of bonded components is acompromise between the two. In the absence ofother indications, a good starting point is to usea temperature that is approximately 75% of themelting point of the filler metal, expressed indegrees Kelvin.

Some further improvement in the residual-stress level can be obtained by using a morecompliant filler, especially when the joint is rea-sonably wide (>25 μm, or 1 mil). Solders, almostwithout exception, have low elastic moduli,which is related to their low melting points, butthere are only minor differences between indi-vidual alloys. Hence, the only possibility for ob-taining a joint with improved compliance, shortof redesigning the assembly, is to replace a hardsolder with one that will creep. The benefit ofsuch a change might be offset by inferior di-mensional tolerancing, and the joint design andoperating environment should take this trade-offinto consideration.

The thermomechanical properties of most ofthe family of indium solder alloys (pure indium,silver-indium, indium-lead, indium-tin) aredominated by creep behavior at all normally ex-perienced temperatures and strain rates. Thus, asoldered joint made with one of these alloys willalways creep when subject to stress. The fact thatthey creep so readily, even at cryogenic tempera-tures,means that jointsmadewith indium soldersare substantially more resistant to failure by fa-tiguethanmightotherwisebeexpected(seeChap-ter 2, section 2.1.6).

Wide joint gaps (>500μm,or20mil) can some-timesbeused tominimize theeffectsofexpansionmismatch between two components. The soldermust have high viscosity in order to fill suchwidejoints. This is achieved by either using a fillermetal with a wide melting range and performingthe joining process below the liquidus tempera-ture, so that thealloyisnot fullymolten,orbymix-ing in metal powder with a higher melting point.Spacers are required to control the joint gap.Widejoints can also be achieved by inserting porousshims,asdescribed insection4.3.4of thischapter.One particular merit of wide joints to ceramiccomponents is that they are tolerant to variationsin thewidth of the joint gap, thereby obviating theneed tocloselymachine thematingsurfacesof thecomponents, which tends to be costly and canweaken the material by creating subsurfacecracks. Where the joint is wide, its mechanicalproperties are essentially those of the bulk solder(see section 4.3.3 of this chapter).Fig. 4.7 Bow distortion of a bimetallic strip


Avariety of mechanical schemes are availableto assist in overcoming the problem of thermalexpansion mismatch. Several approaches thathave proved successful are described follows.

4.2.1 Controlled Expansion MaterialsFabrication of most products requires the use

of many different materials in their assembly,each selected for a particular property or com-bination of properties it offers. Ashby and co-workers have devised a scheme of materials se-lection charts that pictorially represent materialsaccording to their properties, which is intendedto facilitate selection [Ashby 1994]. Because sili-con and most other semiconductor and opticalmaterials, as well as engineering ceramics, have

low thermal expansion coefficients, compared tometals, there is always interest in controlled ex-pansion alloys that can be used to bridge dimen-sional changes between thesematerials andmetalcomponents in the same assembly (Fig. 4.8).

Single-phase materials, which include manyengineering ceramics, such as alumina and alu-minum nitride, change dimension in a fairly lin-ear manner with temperature. This is certainlytrue over the temperature range of interest forconsumer products (�50 to �150 °C). For com-mon metals, their coefficients of thermal expan-sivity (CTE) at room temperature are directlyproportional to their melting points. The expan-sion effect is attributable to the atomic vibrationsof the crystal lattice. Raising the temperatureincreases the vibration, which means that each

Fig. 4.8 An Ashby materials selection chart. The linear expansion coefficient, �, plotted against the thermal conductivity, �. Thecontours show the thermal distortion parameter �/�.


atom occupies a greater space, and hence, thematerial grows in size. Because the maximumvibration is restricted by a material melting tem-perature, a low-melting-point material has asmaller expansion temperature range and willexhibit a higher CTE value than a metal with arelatively higher melting point, as demonstratedin Fig. 4.9, using the data in Table 4.5.

The principal low-expansion metals are tung-sten and molybdenum, which have CTE valuesat 20 °C of 4.5 and 5.1 10�6/K, respectively.These metals, especially tungsten, are hard andstiff, offering little compliance, and they are alsorelatively difficult to machine. Both are densematerials, so components fabricated from themwill be heavy. Their high melting point meansthat their atoms are strongly bonded, and, con-sequently, their machining requires considerableenergy. Consequently, the net cost of parts fab-ricated from these metals can be significant. Ti-tanium has a low CTE (5.6 10�6/K) for ametal, low density, but a poor thermal conduc-tivity of only 15 W/m · K. This rules it out forany application where substantial heat transferby conduction is an additional requirement. Alu-mina and aluminum nitride also suffer from en-ergy-intensive productionmethods and difficultyof machining. These ceramics have to be fabri-cated in near-net shape forms, so the tooling costof new components is quite high. Furthermore,their thermal conductivity is inferior to that ofmost metals.

The lackof single-phasematerials thatoffer thecombination of low thermal expansivity, highthermalconductivity, and,preferably, lowdensityhas led to the development of several families ofmultiphase materials with tailored properties.Theseincludeiron-nickelalloys,copper-tungstenand copper-molybdenum alloys, metal-metallaminates, metal-ceramic laminates, metal-ma-trix composites, and metal-metalloid alloys (inparticular, a family of aluminum-silicon alloys).

These materials and their key physical propertiesare listed in Table 4.6. Further information isgiven in the following section.

4.2.1.1 Iron-Nickel Alloys

Most readers will probably know iron-nickelalloys by trade names that include Invar, Kovar,Nilo-K, Nilo Alloy 42, or by the UNS numberK94610 or DINWNr 1.3981. Kovar and Nilo-K,for example, are essentially the same Co-Fe-Nialloy, of approximate composition 17Co-54Fe-29Ni, that has a CTE at 20 °C (68 °F) of 5.8 10�6/K. These alloys are readily available froma number of manufacturers in many shapes andforms and are competitively priced. Iron-nickelalloys are widely used in electronic packaging.Being fairly soft and ductile in the annealed con-dition, they are also used in shims for joininglow-expansion ceramics to higher-CTE metals,where they are capable of distributing and ab-sorbing expansion mismatch stresses.

Iron-nickel alloys offer abnormally low ther-mal expansion compared to their constituents.Indeed, over a limited range of temperature, it ispossible to design material that has zero expan-sivity. The unusual expansion characteristics ofthese alloys can be ascribed to the fact that theyare ferromagnetic.

At temperatures above the Curie point (ap-proximately 450 °C, or 840 °F), these alloys areferromagnetic and exhibit normal thermal ex-pansion characteristics. Below the Curie tem-perature, when they are in their ferromagneticdomain, the actual expansion is the sum of thenormal positive expansion due to lattice vibra-tion, counteracted by the negative expansion dueto the ferromagnetism. The latter is termed mag-netostriction. The result is reduced CTE belowthe Curie temperature that can be adjusted bycontrolling the state of cold work in the alloy. Atvery low temperatures, below �100 °C (�148°F), the expansion coefficient of iron-nickel al-loys reverts tomore normal values (1510�6/K).Expansion curves for some iron-nickel alloys aregiven in Fig. 4.10 and 4.11.

It is important to recognize that iron-nickelalloys only possess low and controlled expan-sion coefficients over a limited range of tem-perature. This is due to a combination of thelimited temperature interval over which mag-netostriction compensates for normal thermalexpansion, plus the necessity for the alloy tobe in a very particular state of cold work.Thus, further processing that involves either

Fig. 4.9 General relationship between coefficient of thermalexpansion, or CTE (between 273 and 373 K), and

melting point for metals, Tm. Adapted from Li and Krsulich [1996]


mechanical working or a temperature excur-sion will change the internal stress in the ma-terial and thereby its expansion coefficient. Fig-ure 4.12 illustrates the expansion coefficients ofFe-36Ni and Fe-42Ni alloys as a function oftemperature and state of anneal. As can beseen, the lower-nickel-content alloys can befabricated to have a smaller overall thermal ex-pansion, but it is stable over a more limitedrange of temperature.

These low-expansion alloys are notoriouslydifficult to machine, particularly in thin sec-tions to a high surface finish. Specialist metalworking companies have developed the neces-sary skills to deliver products of consistent andhigh quality. It is not possible to solder di-rectly to iron-nickel alloys using electronic-grade fluxes, and parts are normally plated withnickel and gold. Brittle intermetallic com-pounds form at the interface when cobalt-

containing mixtures are soldered directly withtin-base solders.

4.2.1.2 Copper-Molybdenum andCopper-Tungsten Alloys

Low-expansivity molybdenum (CTE � 5.1 10�6/K at 20 °C, or 68 °F) or tungsten (CTE �4.5 10�6/K at 20 °C, or 68 °F) is added tocopper (CTE � 17.6 10�6/K at 20 °C, or 68°F) to produce controlled-expansion alloys.Mol-ten copper is virtually insoluble in both molyb-denumand tungsten, and various techniques havebeen devised to enable the manufacture of 100%dense alloys of these materials. These techniquesinclude powder metallurgy as well as liquid in-filtration casting. Controlled-expansion alloyscan be produced with a continuum of propertiesthat range from essentially pure copper to purerefractory metal. By this means, families of ma-

Table 4.5 Metals and their properties used to prepare Fig. 4.9Melting point

Metal °C K CTE at 300 K, 10�6/K

Tungsten 3422 3660 4.5Molybdenum 2623 2888 5.1Palladium 1555 1817 11.0Gold 1063 1336 14.1Aluminum 660 933 23.5Cadmium 321 594 31.0Lithium 181 454 56.0Mercury �39 235 60.0

CTE, coefficient of thermal expansion

Table 4.6 Indicative physical properties for selected semiconductor and low-expansion materials at20 °C (68 °F). Exact values depend on the composition of the material, method of manufacture, testmethod, and test conditions. Reference should be made to suppliers’ data sheets for precise values.

MaterialThrough-thickness thermal

conductivity, W/m • KIn-plane thermal expansion

coefficient, 10�6/K Density, g/cm3

Gallium arsenide 42 6.5 5.3Silicon 84 2.5 2.3Alumina 20 6.7 3.9Aluminum nitride 165 4.5 3.3Beryllia 260 7.2 2.9Molybdenum 140 5.1 10.2Titanium 15 5.6 4.5Tungsten 174 4.5 19.3Copper-alumina-copper 26 7.3 4.1Copper-molybdenum-copper 166 5.5 10.0Copper-85% tungsten alloy 180 7.2 16.1Copper-85% molybdenum 160 6.7 10.0Invar (Fe-36Ni alloy) 14 2.2 8.1Kovar (Fe-29Ni-17Co alloy) 17 5.8 8.4Aluminum-50% silicon alloy(a) 150 11.0 2.5Aluminum-70% silicon alloy(a) 120 7.4 2.4Aluminum-68% silicon carbide composite 150 7.2 3.0Beryllium-30% beryllia composite (E20)(b) 210 8.7 2.1Beryllium-51% beryllia composite (E40)(b) 220 7.5 2.3

(a) As supplied by Osprey Metals Ltd. (b) As supplied by Brush Wellman Inc.


terials have been developed having a range ofcontrolled CTE values. A common ratio is 15%

Cu, which gives a reasonable boost to the ther-mal conductivity without greatly impairing thethermal expansivity of the base metal. More im-portantly, perhaps, is that the addition of a softcopper phase in the otherwise refractory metalmatrix greatly improves themachinability. Thesealloys still require considerable care to machine,because the copper is much softer than the re-fractory metal constituent, and it is easy to causedamage in the surface region.The principal draw-back of these alloys is their high density, al-though this is partly offset by their highmodulus,which means that thinner sections can some-times be used, depending on the functional re-quirements and design of the product.When usedin precision assemblies, these alloys need to bein the annealed condition to ensure dimensionalstability on thermal cycling.

4.2.1.3 Copper-Surface Laminates

Copper can be attached directly to alumina viathe copper/copper oxide eutectic reaction. Theseproducts are often marketed as direct-bondedcopper. The copper can be patterned so that it canalso fulfill the function of, admittedly, a verylow-density printed circuit board, and this hasmade it a very popular substrate for electronicpower modules.

A more recent development of this materialhas been its commercialization in the productionof arrays of through-thickness vias, which arefilled with copper. This arrangement does notchange the in-plane thermal expansivity of thealumina but significantly improves the averagethrough-thickness thermal conductivity. Obvi-ously, care must be taken when using this ap-proach with high-intensity heat sources, typifiedby optoelectronic and microwave devices, to en-sure that the pattern of heat studs matches thepoint sources in the semiconductors.

Nickel, %

0 10 20 30 40 50 60 70 80 90 100

18

16

14

12

10

8

6

4

2

Coe

ffici

ent o

f exp

ansi

on, 1

0–6 /

K

1.8

1.6

1.4

1.2

1.034 3735 36

Fig. 4.10 Expansion coefficient of iron-nickel alloys, at 20°C, as a function of composition in the an-

nealed state

Fig. 4.11 Total expansion of an Fe-36Ni alloy between�220and �250 °C

Fig. 4.12 Thermal expansion characteristics of Fe-36Ni and Fe-42Ni alloys as a function of temperature and state of anneal


Copper-molybdenum-copper and copper-Invar-copper laminates are also available but pro-vide a subtly different balance of properties. Theydo not offer the same stiffness and patterningabilities as the ceramic-cored alternatives andare less widely used.

4.2.1.4 Composite Materials

Some examples of composite material are in-cluded inTable4.6.Each is representativeofadif-ferent family: metal-metalloid, metal-ceramic,and metalloid-metalloid. These are all relativelynew materials, and many variations exist on themarket.Metalloid-metalloid composites are veryimmature products, and citing reliable propertyvaluesfromanassessmentofpublishedvalueshasproved problematic. Controlled-expansionmate-rials comprise two or more components: one, ametal or metalloid, and the other a lower-expan-sionmetal,metalloid,ornonmetal.Byvarying therelative proportions of these constituents, theCTE values can be adjusted over quite a widerange.

Alloys of copper-tungsten and copper-molyb-denum, made by solid-state sintering together ofpowders of the separate constituents and referredto previously, represent controlled-expansioncomposites of metal components. Metal-ceramiccomposites are typified by aluminum/silicon car-bide and beryllium-beryllia. These are of lowerdensity, but their principal limitation stems froman inability for machining at high speed and ob-taining an acceptable surface finish, because ofthe substantially different hardness of the con-stituents. Diamond tools need to be used on alu-minum/silicon carbide because of the presenceof extremely hard silicon carbide particles. Afterall, silicon carbide is itself a traditional cuttingtool material, and silicon carbide in a matrix ofaluminum constitutes a cemented carbide. In thecase of the Be-BeO materials, the particles ofBeO are poorly bonded to the beryllium metalmatrix, and machining tends to result in surface-opening cavities. There is also a potential healthhazard should beryllia dust be generated.

For these reasons, parts fabricated in thesemetal-ceramic composites tend to be producedusing near-net shaping methods. Care is takenduring the production to ensure that the part hasan outer skin of metal and that the finishing ma-chining does not expose ceramic particles. If itdoes, then there is a high chance that the particleswill either crack or pull out, and if neither ofthese occur, then the presence of an exposed

ceramic phase tends to cause local adhesion prob-lems if the part needs to be metallized.

Metalloid-metalloid composites are repre-sented by carbon-carbon fiber composites, whichwere originally developed for exceptionally high-temperature applications such as aircraft brakesand rocket motor nozzles. They can be metal-lized and also impregnated with copper. This isdesirable because it facilitates soldering and al-lows the in-plane thermal expansivity to be in-creased to more normal values. Care needs to betaken when selecting the carbon fibers and theweave in the component. Carbon fiber is a highlyisotropic material and hence, the properties ofparts fabricated by using it can vary accordingly.Likewise, there are many grades of carbon fiber,and, as onemight expect, it is themore expensiveones that possess the most desirable and stableproperties. Carbon-fiber-reinforced aluminumand copper composites are also available, butthese are highly anisotropic and tend to be di-mensionally unstable.

Metal-metalloid composites are perhaps bestrepresented by the series of silicon-aluminumalloys, marketed as controlled-expansion (CE)alloys and containing between 27 and 70wt% Si,to achieve a range of CTEs, as can be seen in Fig.4.13 [Jacobson 2000]. These are strictly alloys,not composites, produced from the melt by theOsprey spray-forming process. A relatively fine,two-phase, or duplex, microstructure of continu-ous, interwoven matrices of silicon and alumi-num forms naturally by eutectic solidificationduring the spray-forming process (Fig. 4.14). Thesilicon phase constrains the expansion of the alu-minum, while the latter fraction is largely re-sponsible for the transport of heat. Both siliconand aluminum are elements of low density, withsilicon being lighter than aluminum. The CE7alloy (Al-70Si) is closely expansion matched to

Fig. 4.13 Coefficient of thermal expansion (CTE) of Ospreycontrolled-expansion alloys (based on aluminum-

silicon) as a function of the proportion of silicon, inweight percent


alumina over a wide range of temperatures, mak-ing it suitable for packaging microwave/radiofrequency circuitry.

4.2.2 InterlayersOne route toward reducing themismatch stress

concentration that develops in soldered and, inparticular, in brazed assemblies involves a re-design of the joint to accommodate one or moreinterlayers. There are two basic configurationsthat are described in the literature.

In the first approach, a compliant interlayer isinserted that will yield when the joint is placedunder stress and thereby reduce the forces actingon the components. This approach is not par-ticularly effective with soldered joints for thefollowing reasons. First, the moduli of compliantmetals that are most effective in accommodatingstress (in particular, silver and copper) are tooclose to those of many solders to provide muchstress relief, so the solder will tend to yield inpreference to the interlayer. Second, solders tendto form hard, interfacial phases with most engi-neering metals and alloys, which will confer ahigh modulus to the adjacent interlayer and ac-tually exacerbate the situation. The approach istherefore usually only relevant where “hard”gold-tin solder is mandated. Filler metals withintrinsically low and controlled thermal expan-

sion coefficients are discussed in section 4.3.3.2of this chapter.

An alternative approach for reducing mis-match stress concentrations is to redistribute thestresses across a much wider zone, so that thestress is within tolerable levels everywhere in theassembly. A graduated redistribution of stressmay be accomplished by inserting into the jointone or more thick shims or plates that have CTEsthat are intermediate between those of the abut-ting components. The plates must be sufficientlythick so that they are not significantly distortedby the imposed stresses and therefore are notusually less than a few hundred microns (severalmils) thick.An assembly containing a single platewith an intermediate thermal expansivity isshown in Fig. 4.15.

This approach is particularly suitable wherethere is a need to joinmetals to ceramics andotherceramic-like nonmetals. If the intermediate plateis selected tohavea thermal expansioncoefficientthat is close to that of the nonmetal, then it is pos-sible to transfer themajor proportion of the stressto the more robust metallic part of the assembly.Where the two components have greatly differentthermal expansivities, it may be necessary to useagraduatedseriesofplates toreducethemismatchstresses in each joint to an acceptable level.

A monolithic plate of graded compositionand thermal expansivity can be used in place ofa series of discrete homogeneous plates. Typi-

Fig. 4.14 Photographs of the microstructure (micrographs) of two controlled-expansion (CE) alloys produced by spray forming,showing their uniform phase distribution. The lighter and darker phases are primary aluminum and silicon, respectively.

(a) CE17 (Al-27wt%Si. (b) CE7 (Al-70wt%Si). Courtesy of Osprey Metals


cally, these may be prepared from powder com-pacts. Copper/tungsten components of this typeare made by infiltrating a loose compact oftungsten powder with molten copper. By adjust-ing the packing density of the powder, the rela-tive proportions of copper and tungsten willvary, and the properties of the component canvary from those of essentially pure copper toapproximately 95% W. This enables one side ofthe component to be made tungsten-rich, with alow expansion coefficient, and the other sidecopper-rich, with a much higher expansion co-efficient. Because there are no abrupt interfacesin such a component, it can survive thermalcycling over wide ranges of temperature almostindefinitely without suffering distortion throughcreep or fatigue fracture. A graded copper-tung-sten plate is shown in Fig. 4.16.

If one or both of the components is highlybrittle and vulnerable to fracture under a tensileor shear stress, it is often the practice to providereinforcement by attaching it to a more mechani-cally robust metal plate of similar expansivity.This subassembly can then be joined via the re-inforcing plate to further components of differentexpansivities, using approaches detailed previ-ously. Such a configuration involving a rein-forced subassembly is used formounting of semi-conductor edge-emitting lasers, because thewavelength of the light produced is a strong func-tion of the strain in the semiconductor material.

The following disadvantages are associatedwith the use of graduated joint structures basedon the use of intermediate plates to accommo-date mismatch stresses:

• An increase in the thickness and often in theweight of the assembly, whichmay be signifi-cant.Thismodificationwill also introduce ad-ditional materials and fabrication costs.

• At least two soldered joints are used in placeof a single joint. Because further materialsare introduced to the assembly, alternativefiller alloys and joining processes may needto be developed and qualified.

• The thermal and electrical conductance be-tween the joined components is likely to bedegraded. This is a consequence of the in-crease in the overall thickness of and numberof interfaces in the assembly. Furthermore,materials with low expansion coefficientstend to be poor conductors. An exception isdiamond, and a thin shim (300 μm, or 12mils) will function both as a buffer againststrain and as a heat spreader for optical andelectronic components.

• The method is difficult to apply to joints thatdo not have simple planar geometries.

4.2.3 Compliant StructuresEquation 4.1, given previously for calculating

the bow distortion of a bimetallic assembly, im-plies that the mismatch stress is a sensitive func-tion of joint dimension or, more precisely, jointarea. Although the overall size of the assemblyis likely to be fixed by the functional require-ments of the product, it may be possible to re-place one of the monolithic components with afilamentary, brushlike structure. Then, the di-mensions of each individual bimetallic joint canbe made as small as necessary, thereby effec-tively eliminating the mismatch stress from this

Fig. 4.15 Use of a plate of intermediate thermal expansivityto reduce the stress due to thermal expansion mis-

match in an assembly between an aluminum alloy mount and thebody of a solid-state laser

Fig. 4.16 Monolithic plates of graded composition, varyingfromessentially pure copper to approximately 95%

W. The thermal expansivities of the two surfaces differ by ap-proximately 16 10�6/°C (29 10�6/°F).


source, while the high aspect ratio of the fila-ments confers a degree of lateral compliance thatcan accommodate the mismatch strain. Ex-amples of these highly compliant structures areillustrated in Fig. 4.17 and 4.18, and others aredescribed in the scientific and technical literature[Huchisuka 1986].

The so-called flip-chip bonding process usedin semiconductor assembly also results in com-pliant contacts between the components, al-though in this case these are provided by theactual solder joints [Yung andTurlik 1991]. Flip-

chip bonding involves electroplating or vapordepositing solder bumps on the contact pads ofan electronic component.These bumps are heatedto reform the solder as hemispherical balls, whichare typically 0.1 mm (0.004 in.) high. In a sub-sequent stage, the components are joined to cir-cuit boards by reflowing the solder bumps. Theentire sequence is shown schematically in Fig.4.19. The substrate is prepared in such a mannerthat solder wetting is laterally confined, in orderto maximize the height of the solder pillars thatconstitute the joints. The Pb-5Sn solder is widely

Fig. 4.17 Examples of compliant structures for mitigating mismatch expansivity (�) of the abutting components

Fig. 4.18 (a) Longitudinal and (b) transverse sections through a compliant structure that is capable of accommodating a thermalexpansivity difference between joined components. (a) 99. (b) 450


used for flip-chip bonding because it has a highcompliance, although its fatigue resistance isrelatively low for a metal. The flip-chip processis described in detail in Chapter 5, section 5.2.

The use of compliant structures of the formsshown obviously incurs a cost penalty, due to thegreater complexity of manufacture. The conduc-tance between the components via the filamen-tary member will also be impaired. Even withfilaments having a hexagonal cross section toproduce a close-packed structure, it is difficult toobtain a compliant structure that will work ef-fectively with a packing density of greater thanapproximately 85% [Glascock and Webster1983]. Furthermore, the ability to simulta-neously make large numbers of small-area jointsis by no means a trivial exercise but one thatdemands stringent control of tolerances andhighly specified joining processes.

4.2.4 The Role of Fillets

Wherever practicable, it is good practice todesign a joint so as to encourage the formationof a fillet. Then, even if the joint contains voids,for whatever reason, the fillet will serve to sealthe joint because there is a higher probability thatwhere fillet formation is promoted, these tend tobe continuous and void-free. Fillets also have abeneficial effect on the mechanical properties ofjoints. Well-formed fillets of filler metal can en-hance the measured tensile, shear, and peel

strengths by as much as an order of magnitude,the value depending on the geometry of the jointand the mode of stressing. This is illustrated forresistance to peel initiation in Fig. 4.20 to 4.22.The improvement in mechanical properties maybe attributed to the gradual transition in geom-etry that the fillet provides in minimizing stressconcentration at the joint periphery and therebyjoint failure through crack and peel initiation atthe surface. The stress concentration can be cal-culated and is presented as a function of contact

Fig. 4.19 Schematic illustration of the flip-chip joiningprocess

Fig. 4.20 Typical peel force (P) profile of original geometryjoints of a flat-pack module on a circuit board. The

peaks in peel strength are associated with the fillets at each endof the joint.

Fig. 4.21 Experimentally derived relationship between thestress required to initiate peel fracture and the

height of the solder fillet


angle in Fig. 4.23 [Eley 1961]. Provided the sol-der wets to form a fillet with a contact anglebelow 30°, there is no appreciable stress con-centration at a change in geometric profile. Thesubject of stress concentration is discussed insection 4.3.3 of this chapter.

Because it is difficult to form fillets of exactlyreproducible geometry, testpieces for mechani-cal testing are often designed to exclude fillets,either by preventing their formation through theuse of nonwettable surfaces outside the edge ofthe joint or by removing any that happen to form.Although this practice makes the measurementsmore readily reproducible, it modifies jointstrengths to an extent whereby they may not berepresentative of most practical situations.

Examination of edge fillets can provide anindication of the filler/substrate contact angle atthe onset of solidification. This is usually taken

to be a good indication of the overall quality ofthe joint. While this is generally true, visual in-spection of edge fillets can be misleading as tothe quality of the interior of the joint, as dis-cussed in Appendix A4.2 in this chapter. Infor-mation on internal integrity can only be obtainedby radiography, scanning acoustic microscopy,transient thermography, or destructive methods.

4.3 Constraints Imposed by theComponents and Solutions

A large-area soldered joint may be defined asone where the total joint area exceeds approxi-mately 20 mm2 (0.3 in.2) and the length in anydirection is greater than approximately 5 mm(0.2 in.). This definition is based on the followingpractical criteria:

• The significant distortion of assemblies thathave joints between materials of CTE dif-fering by as little as 5 10�6/°C (9 10�6/°F). Distortion is related to the solidus tem-perature of the filler alloy, and this problemis therefore most acute for high-melting-point filler alloys.

• The incorporation of significant void levels(above 10%) in joints. This problem tends tobe more pronounced in soldered than inbrazed joints and is associated with the lowerjoining temperatures that are used for these.

Fig. 4.22 Typical peel force (P) profiles of joints modified by increasing the solder volume and providing for a fillet at both ends.The joint strength is unchanged, but the resistance to peel initiation is greatly improved.

Fig. 4.23 Roleof fillets in reducing stress concentrationat thechanges in section between abutting components


Distortion of assemblies can arise from a num-ber of causes, some of which were discussed inChapter 1, section 1.3.2. Apart from CTE mis-match, the most common sources of warping orbowing are uneven heating, which leads to tem-perature gradients in the components, and re-sidual stress from earlier stages of fabrication,which is relieved in the heating cycle. Distortionfrom these causes can be avoided by performingthe joining operations under carefully controlledconditions. Notably, heating should only be car-ried out in furnaces that provide highly uniformtemperature zones, with the rate of heating tai-lored to the thermal mass of the components, andafter appropriate stress-relief routines have beenperformed.

4.3.1 Joint AreaThe strength per unit area of a joint between

components of the same material tends to reducein proportion to the area above some lowerthreshold, often approximately 20 mm2 (0.3 in.2)for a soldered joint. This effect can be seen in Fig.4.24, which shows the percentage of the joint byplan area that comprises voids as a function ofthe (square) component size. There is clearly anincreasing tendency for voids to accumulate inthe joint as its area is increased, with a thresholdjoint length of approximately 1 mm (40 mil).

The voids have two causes:

• Gas trapped or generated within the joint• Solidification shrinkage of the molten filler

metal

Each of these is considered in further detail in thefollowing sections.

4.3.1.1 Trapped Gas

By far, the largest source of voids in solderedjoints is trapped gas, even for joints made in highvacuum using both components and filler alloysthat have been given a vacuum bakeout prior tothe joining cycle. Void levels in soldered jointsgreater than approximately 100 mm2 (1.5 in.2) inarea and that are no narrower than 5 mm (0.2 in.)can typically reach 50% of the joint volume,which is consistent with the entrapment of air oran evolved gas. Figure 4.25 shows a radiographof a joint between a silicon chip and a metallizedceramic package, made in high vacuum using asolder preform, that graphically demonstrates theproblem. Figure 4.26 is a scanning acoustic im-age of the same component and convenientlyillustrates the correspondence that can be ob-tained between the two analytical techniques.

Air becomes trapped when the componentsare assembled, with the mating surfaces formingan effective seal. As the temperature of the as-sembly is raised to the peak process temperature,the trapped air is augmented by gas (usuallymoisture) evolved from the joint surfaces. Thetotal gas volume will increase as the temperatureis raised and the ambient pressure is reduced, inaccordance with the gas law (pressure volume� constant absolute temperature).

If the path length between a gas bubble and thejoint periphery is small, the gas pressure cannormally exceed the hydrostatic force exerted bythemolten fillermetal, allowing the gas to escape

Fig. 4.24 Void content versus joint length for a range of rep-resentative solders. The substrates were square

coupons of polished alumina metallized with thin-film titanium/gold, applied by sputtering. The joints were made at a superheatof 25 °C, and the void level was assessed by quantitative x-radiography.

Fig. 4.25 Radiograph of a silicon chip (10 10 mm) sol-dered into a metallized ceramic package. Voids in

the joint gap are evident as the light areas.


to the surrounding atmosphere. However, thelimit to the path length for this to occur is of theorder of 1 mm (40 mil) for soldered joints (Fig.4.24), but it is significantly longer for brazedjoints, because the process temperature, andhence the pressure developed by trapped gas orvapor, is higher. Adsorbed water is particularlydetrimental in this regard, because it expandsrapidly as it vaporizes. One method of allowingthe trapped gas to escape is to momentarily splitthe joint apart during the reflow process. Whilethis has been demonstrated as being highly ef-fective in laboratory trials, it is not readily ap-plicable to volume manufacture [Xie, Chan, andLai 1996].

When undertaking soldering using preforms,a common mistake made when admitting thesolder into the joint gap is to use a foil preformof similar dimensions to the plan area of the joint.This approach typically results in a high level ofvoids because of the large surface area exposedto the atmosphere.

An effective method of removing trapped airis to design the joint in such a manner that themolten solder is made to flow from the center ofthe joint out toward the periphery or through thejoint from one edge, as tends to occur when feed-ing the filler into the joint from a rod or a wirepreform. Suitable arrangements for achieving thistype of flow are illustrated schematically in Fig.4.27. The advancing front of molten solder isthen able to displace the vapor and air ahead ofit as it flows into the joint gap. However, neitherapproach is entirely satisfactory. A preform ofincreased thickness and reduced area should notexceed 2 mm (80 mil) in plan, which can makejigging of the components difficult. Moreover,the solution of introducing the solder from oneside of the joint is only effective with filler alloysthat do not react strongly with the substrate ma-terials to stifle flow of the molten alloy in its paththrough the joint (see section 4.1 of this chapter).

There are a number of solutions to this prob-lem.Afirstmodification is to introduce the solderfrom more than one side. The preferred configu-ration is to place two preform discs at either sideof the joint, as illustrated in Fig. 4.28. Disc pre-forms are readily available in a variety of sizes,and, by using a pair, the necessity to have ahighly fluid solder is diminished. The preformscan be placed so that the voids that form will be

Fig. 4.26 Scanning acoustic microscope image of the sol-dered joint shown in Fig. 4.25. Voids in the joint

gap correspond to the light areas.

Fig. 4.27 Two configurations showing flow by a molten filler designed to sweep trapped gas out of a joint


in prescribed locations and their presence can beallowed for in the design of the assembly [Weston1974].

An improvement on this technique uses asingle foil preform in the shape of a cross, withthe arms orientated along the longest diago-nals of the component. By doing so, the jig-ging requirements are greatly simplified [So-colowski 1987]. However, this method requiresa custom-made preform for each application,and foil preforms are difficult to clean me-chanically (to remove the native oxide) by me-chanical abrasion prior to use (see Chapter 3,section 3.3.5). Admitting the solder in the formof round wire preform overcomes three of thefundamental deficiencies associated with usingfoil preforms:

• The surface-area-to-volume ratio is consid-erably reduced. Correspondingly, the detri-mental effects of surface oxide on the sur-face of the solder are greatly diminished.Furthermore, the solder oxide can be easilyremoved by wiping the wire several timeswith a paper tissue soaked in solvent (seeChapter 1, Fig. 1.27 and Chapter 3, section3.3.5). If this operation is conducted imme-diately prior to the heating cycle, then voidsstemming from solder oxide can be effec-tively eliminated.

• A round wire has only one small area of con-tact with each of the faying surfaces. Thisprecludes gas pockets being trapped duringjigging, and the advancing solder front is al-ways in the optimal location to sweep out airfrom the joint gap. As a side benefit, solderwire is readily available at a low price pre-mium over the metal content, compared tofoil preforms, and a few stock sizes can beused for a wide variety of joining applica-tions.

• Ductile solder wires can be cold compressionwelded together and shaped as a cross (Fig.3.30). This configuration provides amechani-cally stable platform for jigging as well as theoptimal fluid flow pattern for the molten sol-der to sweep out air and fill the joint gap, asillustrated schematically in Fig. 4.29. Crosseswith short arms and thick solder wire arepreferable, from the considerations of opti-mal fluid flow and minimizing the surface-area-to-volume ratio of the filler metal.

The solder fill of 100 mm2 (0.155 in.2) jointscould be consistently maintained at over 95%using this approach, comparedwith only 45% forthe joints made using a single flat preform thesame size as the joint area [Lodge, Humpston,and Vincent 2001]. A test structure made usingthis method, with the difficult process constraintsof an indium solder, fluxless, and at only 10 °C(18 °F) superheat, is shown in Fig. 4.30.

An aid to reducing the volume of trappedgas is to reduce the number of surfaces in thejoint. Solders can be applied as vapor depos-ited, electroplated, or “tinned” coatings to thecomponents, thereby eliminating two free sur-faces. However, considerable care must betaken to ensure that the coated layers do notthemselves contain significant volumes of gasesor other volatile constituents. Particularly in theelectronics and photonics industries where thepiece-part cost is relatively high, the selectiveapplication of solders by vapor-phase tech-niques, which result in predeposited coatings ofhigh purity and density, can often be justifiedon economic grounds if it results in improvedyields.

Fig. 4.28 Dual discs of preformused to reduce the incidenceof voiding in large-area joints

Original preform Solder flow

Fig. 4.29 Schematic illustrationof theoutwardflowbyacen-tral crossoffillermetal, aconfiguration thathelps to

prevent entrapment of vapor in pockets in a large-area joint


Substrates with solder already applied in de-fined areas are now available commercially froma number of suppliers. These have the advan-tages of:

• Piece-part inventory and number of suppliersare reduced.

• Jigging is likely to be simpler.• Thickness of the solder joint is decreased,

because the solder layer can be substantiallythinner than the minimum practicable thick-ness of circa 25 μm (1 mil) required for asolder preform.

• Soldering behavior is improved by eliminat-ing two joint surfaces, with all the attendantproblems from the joint gap.

• Solder spread is controlled.• Soldering operation is reproducible.

For some large components, it may be possibleto incorporate vents through the components toprovide a passage for trapped gas to escape fromwithin the joint. By careful design, this vent canbe further exploited to increase the effectivenessof the pressure variation process described sub-sequently in section 4.3.2 of this chapter [Hump-ston et al. 2001].

Evolved vapor or gas can originate from sev-eral sources, in particular:

• Organic residues and adsorbed water vaporon the surfaces to be joined. These speciesvolatilize as the temperature of the compo-nents is raised. Residues can be minimizedby carefully precleaning the surfaces.Abake-out in vacuum immediately prior to joining isusually effective in removingwater vapor andorganic residues, provided that the tempera-ture used exceeds approximately 150 °C (300°F). Reactive ion etching, using a hydrogenor halogen plasma, and oxygen plasma ash-ing are component cleaning methods thathave recently gained in popularity because oftheir effectiveness at dealing with organiccontamination on surfaces and the greateravailability of off-the-shelf equipment. Theseprocesses have the further benefit of usuallyinvolving a combination of elevated tem-perature and reduced pressure, coupled witha chemically active ingredient that helps re-move volatile species.

• Volatile materials within the bulk of the com-ponents. Problems with volatile materials aremost pronounced when the components areporous or polymeric materials and when thecomponents (including any metallizations

present), the solder, or the flux contain con-stituents that volatilize during the heatingcycle. A vacuum bakeout immediately priorto the joining cycle can help prevent subse-quent outgassing from porous materials.

Gas evolution frompolymericmaterials is usu-ally caused by thermal decomposition. The onlypractical remedy in this instance is to use eithera lower-temperature process or to change thepolymeric material to one that has superior ther-mal stability. Some materials used for printedcircuit boards are less stable than others, often inrelation to their cost. Similar considerations ap-ply to metallic components and filler metals,where these contain volatile elements such aszinc, magnesium, cadmium, and, to a lesser ex-tent, manganese, and also to fluxes and pastes,which generally contain volatile constituents.

One common method of introducing a solderinto a joint gap is in the form of a fluxed paste.If not correctly formulated and strict process con-trols observed, jointsmadewith fluxed paste tendto contain a higher proportion of voids than jointsmade with a fluxed preform. This relationshiphas been studied. The voids are primarily causedby the volatilization of entrapped flux. The pro-pensity of voids to form increases with decreas-ing wettability of the faying surfaces, decreasing

Fig. 4.30 Example of a large area, 10 10 mm (0.4. 0.4in.), made fluxless, at 10 °C (18 °F) superheat,

using In-15Pb-5Ag solder introduced in the form of a wire cross,shown in Fig. 3.30. The joint fill is revealed by x-radiography. Aline of residual voids marks the location of the original wire crossthat the surface oxide regrew in the interval between cleaningand melting. Courtesy of BAE SYSTEMS


flux activity, increasing flux volume, and in-creasing plan area of the joint. The extent ofvoiding is also influenced by the paste design, inthat the sooner coalescence of the metal occurs,relative to the ability of the flux to deal with theoxide present on the faying surfaces, then thehigher will be the resulting void content. Like-wise, the use of flux media with high boilingpoints relative to the coalescence temperature ofthe solder paste tends to exacerbate the forma-tion of voids [Hance and Lee 1992].

4.3.1.2 Solidification Shrinkage

In any soldered joint, a fraction of the residualvoiding does not derive from trapped air, mois-ture, or gas. These residual voids are extremelydifficult, if not impossible, to remove becausethey are intrinsic to the filler metal, being causedby the shrinkage when it solidifies and furthercools. Table 4.7 lists values for the shrinkagevolume contraction of elements common tomanysolders. The reservoir of solder represented bythe edge spillage fraction is seldom able to feedlarge-area joints and compensate in part for thecontraction, because the outer extremities ofjoints usually solidify first through radiative heatlosses to the surroundings.

The magnitude of solidification shrinkage, asgiven in Table 4.7, accounts for the fact that it isdifficult to make joints of large area that containless than approximately 3 to 5% voids. Shrink-age voids tend not to occur in small or narrowjoints (<2 mm, or 0.08 in., in one of the joint areadimensions). This is because the thermal gradi-ents that usually develop along a joint when theassembly is cooled from the joining temperatureare large in relation to the joint dimensions, andthis causes the filler to directionally solidify fromone edge to the other, thereby preventing voidsfrom forming within the joint.

It is possible to achieve the same effect inlarge-area and wide joints by imposing a tem-perature gradient on the assembly, from eithercenter-to-edge or edge-to-edge, such that someperiphery of the joint is always the last portionto solidify. However, this is not always easy toachieve, especially when large numbers of com-ponents are involved. Furthermore, the imposi-tion of a temperature gradient on a large assem-bly may produce stress gradients and therebydimensional distortion of the components, whichbecomes fixed when the solder solidifies.

Where the parent materials and solder are ofsimilar composition, maintaining the assembly

at elevated temperature but below the solidustemperature of the filler can result in a gradualreduction in void levels arising from solidifica-tion shrinkage by vacancy diffusion. This is themechanism by which dry joint interfaces are re-moved in diffusion bonding (see Chapter 1, sec-tion 1.1.7.2). The process times then becomerelatively long and not less than 1 h.

Bismuth and, to a lesser extent, antimony areexceptional amongmetals in having a volume ex-pansionrather thanavolumecontractiononfreez-ing, as shown inTable 4.7.Therefore, by combin-ingbismuthand/orantimonywithotherelements,it is possible to produce solders with essentiallyzero volume change on solidification. The Bi-43Sn solder (melting point of 139 °C, or 282 °F)is an example of one such alloy that has been rec-ommended for applications in the electronics in-dustry,where joints of guaranteedhermeticity arerequired [Dogra 1985]. However, the volumechange that occurs is not instantaneous but oftentakes place over several hours after the solder hassolidified [Manko 2002]. The forces accompany-ing the volume expansion can be significant andmust be allowed for in the joint design.

4.3.2 Void-Free SolderingRegardless of the origin of gaseous voids, a

very successful method has been devised to com-press the trapped gas so that it occupies a smallervolume fraction of the joint. Because this pro-cedure works while the solder is molten, it alsohelps reduce the volume of voids arising fromthe liquid-to-solid phase change through solidi-fication shrinkage.The pressure variationmethodwas developed specifically to reduce void levelsarising from trapped gas in adhesively bondedjoints [Bascom and Bitner 1975]. It can be usedto make large-area joints using solder preforms

Table 4.7 Solidification shrinkage of selectedelements common to widely used solders

Element

Volumecontraction

on solidification,% of solid

Solidexpansivity,linear(a)10�6/K

Liquidexpansivity,

cubic, 10�6/K

Zinc 6.9 31 167Gold 5.2 14 86Silver 5.0 19 97Copper 4.8 17 100Lead 3.6 29 123Tin 2.6 23 87Indium 2.5 25 96Antimony �0.9 10 87Bismuth �3.3 13 132

(a) CTEs of the solids are the average values over the range 0–100 °C (0–212°F), while the liquid CTEs are just above their melting points.


that have void levels consistently below 5% [Mi-zuishi, Tokuda, and Fujita 1988]. Most of theresidual porosity is due to solidification shrink-age, as discussed in the preceding section. Theprinciple of the pressure variation method is touse the pressure of an external atmosphere tocompress the gas trapped in the joint. The pro-cedure, as applied to vacuum joining, is as fol-lows (Fig. 4.31):

• The components to be joined are located ina jig and placed in the bonding enclosure attemperature T1, which is then pumped to areduced pressure, P1.

• The temperature of the assembly is raised toT2 tomelt the filler metal, while keeping pres-sure P1 constant.

• The pressure in the enclosure is raised fromP1 to a value of P2, which is several orders ofmagnitude higher.

• The assembly is allowed to cool to T1 so thatthe filler solidifies while the pressure is main-tained at P2.

The voids corresponding to trapped gas arereduced in volume roughly according to the gaslaw (with suitable corrections applied to takeaccount of the nonideality of the particular gasused). In the simplest case, corresponding to ide-

ality, if the initial void volume at pressure P1 isV1, then, at constant temperature, the volume atpressure P2 is:

V2 �V1 P1

P2

Hence, the greater is P2 in relation to P1, themoreeffective is the method. This condition is alsofound to apply qualitatively to practical situa-tions.

The experimentally derived relationship be-tween pressure variation (P2/P1) and the volumeof voids, V2, obtained using this process is shownin Fig. 4.32. The nonlinearity of the relationshipat large P2/P1 ratios is due to departure from ide-ality of the gas and the hydrostatic friction at theinterface between the component surface and theliquid solder. Solder reflow ovens are availablecommercially thatautomaticallyperformthenec-essary pump/pressure cycles to achieve jointswith fewandcontrolledvoids.Thepressurevaria-tionmethodforminimizingvoids isobviouslynotsuitable for situations where vapor is continuallybeing evolved from volatile constituents.

A general approach that has been found to beeffective in producing well-filled and hermeticjoints is one in which strong metallurgical re-actions occur across the joint during the heatingcycle while a compressive force is applied (seeChapter 3, section3.3, for a discussiononfluxless

Fig. 4.31 Pressure variation method for reducing void levelsdue to trapped gas. T, temperature; P, pressure

Fig. 4.32 Experimentally derived relationship between pres-sure variation and void level obtained in large-area

soldered joints using the pressure variationmethod.Adapted fromMizuishi, Tokuda, and Fujita [1988]


soldering).Thevoid-free jointsobtainedusingthediffusion soldering and diffusion brazing pro-cesses are associatedwith such reactions.Thedif-fusionsolderingprocess isdescribed inChapter5,section 5.9, and diffusion brazing in the plannedcompanion volume Principles of Brazing.

4.3.3 Joints to Strong MaterialsNew materials that have enhanced strengths

are continually coming onto the market. Recentexamples are composite materials such as metal-matrix composites (MMCs) and precipitation-strengthened and dispersion-stabilized alloys.Both of the latter types of strengthening havebeen exploited in high-carat gold, suitable foruse in jewelry and bond wire for electronics[Humpston and Jacobson 1992; Jacobson, Har-rison, and Sangha 1996; du Toit et al. 2002].While there is a desire to exploit these materialsin a range of applications, widespread adoptionis contingent on being able to use the favorablebulk strength levels in joined assemblies. In gen-eral, the strength of a joint, even when preparedby welding, is inferior to that of the materials inmonolithic form. Moreover, the heating cycleused in the joining process can itself degrade theproperties of these materials. For example, alu-minum/SiC MMCs are susceptible to degrada-tion when heated above approximately 500 °C(930 °F) due to reaction between the constitu-ents, which results in the formation of a brittleinterfacial layer of Al3C4 [Iseki, Kameda, andMaruyama 1984].Although solders aremechani-cally inferior to welds and brazes, they are nev-ertheless sometimes used to make joints that arerequired to sustain moderate forces. The primaryexample is joints in copper water pipes, wherethe pressure can easily be 0.7MPa (100 psi) witha very modest pump.

In most circumstances, application of stress toa joint does not result in all regions of the jointsharing an equal proportion of the load. The un-evenness of the stress distribution is referred toas the stress concentration, K. Mathematically,this is a dimensionless number that simply de-scribes the magnification factor of the actualstress at one location compared to the uniformstress that would prevail in the absence of anystress concentrations. Expressed as an equation:

K (stress concentration) �

Local stress at a specified location

Applied force/Joint area

The key to making high-strength joints is to pre-vent the development of stress concentrationsand, at the same time, the strength of the filleralloy must be maximized. These aspects are con-sidered in turn.

4.3.3.1 Joint Design to MinimizeConcentration of Stresses

Stress concentrations can be reduced by usingjoint configurations that distribute the load awayfrom the joint. Further details of this subject aregiven in the planned companion volume Prin-ciples of Brazing, because brazed joints tend tobe used more often than ones made with solderfor load-bearing applications.

In order to understand the origin and magni-tude of stress concentrations that can arise, ref-erence is made to a single-lap joint loaded intension. Stress concentrations arise from twosources: namely, the differential straining of thecomponents and filler, and the eccentricity of theloading path. In lap joints, the shear strength ofthe joint per unit area (length) actually decreaseswith increase in the joint length. This apparentanomaly can be explained by the fact that theshear stress is highest toward the ends of thejoint, so that if the length of the joint is increasedbeyond a certain limiting value, the filler in thecentral portion of the joint will carry little or nostress, with the applied stress concentrating atboth ends, as depicted in Fig. 4.33. This explainswhy simply increasing the length of the overlapdoes not improve the strength of this type of jointbeyond a certain level. The stress concentrationin the joint is proportional to the length of over-lap, up to a limiting value, the thickness of themembers, and to the thickness of the joint. There-fore, the stress concentration is least in thin jointsof short overlap.

Far more relevant to the strength of single-lap joints are the tensile or “peeling” stressesthat act normal to the ends of the joint andoriginate from the eccentricity of the loadingof the assembly. The elastic analysis is rela-tively complex, but the result obtained is thatlongitudinal loading of a single-lap joint effec-tively applies a perpendicular tensile stress ofapproximately four times that amount to theends of the overlaps [Harris and Adams 1984].These perpendicular tensile forces initiate fail-ure of the joint by peel. With the continued ap-plication of stress, the sample rotates in an at-tempt to correct for the axial misalignment, and


the fracture continues to propagate due to peel-type debonding. The stress concentrations in asimple lap joint and their influence on its re-sulting failure mode are illustrated in Fig. 4.34.

Fillets at the edges of a joint act to reduce thestress concentration in that region, as indicatedin Fig. 4.23. They do this by coupling some ofthe applied stress into the ends of the laps, therebyreducing the differential straining between the

components and the filler in the joint and alsoshifting the position of the maximum perpen-dicular tensile stress (originating from the ec-centricity of loading) to outside the joint. Themagnitude of these effects depends on the radiusof the fillets, R, the step height, H, and the elasticproperties of the filler. To be effective, the radiusof the fillets must exceed the step height, that is,R > H, and hence, it is desirable for solderedjoints to have large and well-rounded fillets attheir periphery (Fig. 4.35).

The ideal joint is one in which, under all prac-tical loadingconditions, thefillermetal is stressedin the orientation in which it best resists failure.The complexity of the joint should also take intoaccount the load intensity to be sustained and anyaesthetic considerations. In general, simple, low-cost joint designs work well with unobtrusivejoints and low-level loads, while conspicuousjointswithhigher andmorecomplex loading situ-ationsdemandmoreelaborateandexpensivecon-figurations. Strategies for some of the more com-mon joint geometries are presented as follows.

Lap joints are probably themost common con-figuration because of their use in electronics (sur-face mount) and plumbing (pipe joints). Increas-ing the length of overlap will improve the abilityof the joint to resist load along its length, but,following the law of diminishing return, this is

Fig. 4.33 Schematic illustration of the stress distribution in the filler metal of lap joints of short and long overlap. When stressedin shear, the central portion of a long lap joint carries little or no load.

Fig. 4.34 Failure in a simple lap joint loaded in tension. (a)Stress concentrations. (b) Initiation of failure. Edge-

opening crack (free arrow) formed and propagated by the highnormal stress concentration. (c) Progression of joint rotation tofracture. Plastic bending of the joint region results in the majorityof the failure being due to peel-type debonding. Adapted fromDunford and Partridge [1990]

Fig. 4.35 A lap joint showing step height, H, fillet radius, R,and contact angle,


because the center of the joint carries no effectiveload. Further improvement can be made by ta-pering the ends of the overlap, which is easilyachieved if the components are thin and a filletis encouraged to form in this region. Lap-jointstyles for different stress regimes are illustratedin Fig. 4.36(a).

Butt Joints. For a butt joint between two cir-cular rods subject to tension, there is no stressconcentration. The strength of such joints istherefore proportional to area. However, the plainbutt joint is only suitable for the least demandingof applications. The main reason for this is thatthe joint has very low resistance to bendingforces. The scarf butt joint has the merit of onlyrequiring simple machining to prepare the fayingsurfaces yet is highly efficient at resisting de-formation under load. Scarfing results in differ-ential strains and hence the stress concentrationsat the ends of the joint being considerably re-duced, while the landed or step joint relies on thestep sizes being small to achieve the same effect.Both configurations are symmetrical, and there-fore, axial stresseswill be balanced over the joint.They are illustrated in Fig. 4.36(b).

By making the scarf angle sufficiently small,the joint strength can be made to approach thatof the parent materials; that is, when the scarfangle is 90°, the joint is a butt joint, whereas ifthe scarf angle is 0°, there is no joint, just two

parallel pieces of parent material. From a theo-retical perspective, a radially symmetric tongue-and-groove joint should be the best able to resistloads, but unless the components being joinedare particularly large in diameter, achieving ad-equate filling of the joint could be problematic,and the component preparation costs could bequite high.

Strap joints are often used as cheap alterna-tives to butt joints, because the component piecesare generally simpler, and less precision machin-ing is required, although the thickness andweightof the assembly are increased, and its aero/fluiddynamic performance is often impaired. Somerecommendations for different stress regimes areshown in Fig. 4.36(c). The main problem withthis style of joint, as with lap joints, is that anyasymmetry in thickness or material propertiesresults in stress concentrations that cause the par-entmaterial to fail prematurely just outside of thejoint region (this can erroneously be taken as anindication that the joint is stronger than the par-ent material). The stress concentration can bereduced by adding taper to the straps by ma-chining and, ideally, allowing generous and con-tinuous fillets to form.

In the preceding discussion, the assemblieswere considered to be loaded solely in uniaxialtension. The location of any stress concentrationand its magnitude will change as the stressing

Fig. 4.36 Recommended designs of (a) lap, (b) butt, and (c) strap joints for different stress environments


mode is altered, and hence, the optimal style ofjoint varies depending on the stress environmentin which the component is required to operate.

4.3.3.2 Strengthened Solders to EnhanceJoint Strength

One of the limiting parameters of jointstrength, especially of thick joints, is that of thesolder itself. Solders can be strengthened bymet-allurgicalmechanisms involving elements placedin solid solution, microscopic second-phase par-ticles (of either intermetallic precipitates or adispersed refractory phase), and refinement ofthe grains of the filler. These novel solders offersignificantly improved mechanical properties,particularly at room temperature. However,hardly any have yet been developed to the pointof commercialization. An overview of the re-search in this area is presented in Chapter 5,sections 5.5 and 5.8.

Another approach is to load the solder with auniform distribution of coarse particles or fibers(typically, 100 μm to 1 mm, or 4 to 40 mil, insize) of a refractory or nonmetallic material. Thedimensions of the reinforcement dictate that thejoint gap must be relatively wide. In laboratorytests using chopped carbon fibers as the rein-forcement, substantial enhancement of the shearand tensile strength of the joints with respect tothe unmodified solders has been demonstratedand, more particularly, a significant reduction inthe thermal expansivity of the solder [Ho andChung 1990; Cao and Chung 1992]. Further de-tails of these investigations can be found inChap-ter 5, section 5.6.

4.3.4 Thick- andThin-Joint Gap Soldering

Under normal circumstances, a solder jointwill naturally tend to be a few tens of microns(approximately 1 mil) thick. Sometimes, it isnecessary to create joints that are either signifi-cantly thinner (<10 μm, or 0.4 mil) or thicker(>50 μm, or 2 mil). Solders do not have particu-larly good thermal conductivity, so that if a jointis required to transport heat through its thickness,then thin joints are obviously desirable. Thickjoints tend to be encountered where themechani-cal tolerance of the components does not allowfor joints to be consistently made narrower, orwhere creep is desirable in order to relieve me-chanical stress.

It is perfectly practicable to make solderedjoints that are as thin as 2 μm (80 μin.), even involume manufacturing. Joints of this thinnessrequire that the solder is preapplied to one of thejoint surfaces as a high-quality film. Ion-assistedvapor phase deposition and sputtering are theonly reliable methods of achieving this. Elec-troplated and thermally evaporated films are notsufficiently dense, and the residual porosity re-sults in the proportion of oxide being largeenough to interfere with wetting and spreading.It is generally not possible to make thin solderedjoints by simply using a narrow joint gap andhoping that the solder from an adjacent reservoirarea will run in and fill it. This is, again, a func-tion of the very small volume of filler metalwhose composition will change on alloying withthe faying surfaces. Where the joint is thin, thecomposition of the advancing solder front rap-idly becomes uniformly alloyed with materialfrom the joint surfaces, and isothermal solidifi-cation ensues. In thin joints, liquid fluxes inter-fere with wetting and spreading, because thevolatile species have trouble escaping from anarrow gap, as can be seen from Fig. 4.37. Nar-row joints are therefore best made using a gas-eous flux or fluxless (see Chapter 3, sections3.3.6 and 3.3.8, respectively).Anothermethod ofmaking a thin joint is to use a more conventionalquantity of solder, appropriate to a wider joint,

Fig. 4.37 Shear strength of soldered joints in brass testpiecesas a function of joint thickness. Narrow joint gaps

are progressively more difficult to fill, thus decreasing the mea-sured shear strength of thin joints. A gaseous flux is better able topenetrate narrower joint gaps than a liquid flux; consequently,thinner joints can be made before the joint-filling problem ap-pears. Adapted from Manko [1992]


and, once the joint surfaces have been wetted bythe molten filler, apply sufficient compressivestress to overcome the hydrostatic pressure of thesolder to extrude surplus material from the jointgap. Physical stops can be used to control thefinal joint gap. If the lower component is largerthan the upper one, lands can be provided tocatch the overspill in a controlled manner. Thestress required to do this reliably is of the orderof 100 g/mm2 (1 Pa, or 0.2 lb/ft2), which maydamage some components, but it does usuallyenable the soldering process to be fluxless (seeChapter 3, section 3.3.8.3).

A frequently overlooked consideration whenattempting to make thin joints is the cleanlinessof the components and, particularly, the envi-ronment in which the assembly joining is con-ducted. When the desired joint gap is just a fewmicrons wide (typically, 100 μin.), there is nopoint jigging the components in a room wherethe airborne particles are larger! For this reason,the soldering process must be undertaken in asemiconductor-grade clean room, and close at-tention must be paid to the particulate content ofall process gases, cleaning chemicals, tools, andso on. Table 4.8 shows the correlation betweenthe various classes of clean room and their par-ticle size distributions. Clearly, if the require-ment for a joint gap is below 5 μm (200 μin.),then a class M4 (class 100) or better clean roomis required.

When endeavoring to make particularly thicksolder joints, the problem encountered is how toretain the solder in the joint gap, particularly ifthere is compressive stress acting on the moltenfiller. The traditionalmethod of solving this prob-lem has been practiced for generations in the

plumbing industry. The approach is to select asolder that has a wide melting range and to con-duct the joining operation below the liquidustemperature, when the filler alloy is in a pastystate (a mixture of solid and liquid). The pres-ence of the solid phase drastically modifies theviscosity of the alloy and prevents it from flow-ing out of a wide joint gap. The same result canbe achieved by loading a solder with solid par-ticles. Inserting thin parallel shims, for example,of copper, into the joint effectively partitions thejoint gap into a series of much thinner joints andenables conventional joining methods to be em-ployed. Brazing of wide joint gaps is regularpractice as a crack repair technique in the aero-space industry and involves inserting a honey-comb into the joint gap, again to partition thejoint into a number of cells of more conventionaldimensions. Further details can be found in theplanned companion volume Principles of Braz-ing. The successful use of this approach withsolders has not been documented.

Flip-chip interconnects sometimes make useof joints that are thick in relation to their planarea. For example, the solder interconnects onball-grid array integrated circuits (BGAICs) caneasily be 1mmhigh. Here, surface tension forcesare exploited so that the solder interconnectadopts the shape of a truncated sphere and cantherefore be tall in relation to its diameter. Thisapproach to thick joint gaps only works becausethe packaged IC is light in relation to the totaljoint area (i.e., the sum of all the individual sol-der balls). Hence, the total surface tension forceis sufficient to levitate the IC and hence achievea thick joint gap. Flip-chip technology is dis-cussed further in Chapter 5, section 5.2.

Table 4.8 Relationship between clean room class designation and airborne particle size distribution

Federal standard 209F airborne particulate cleanliness classes

Class limits

0.1 μm 0.2 μm 0.3 μm 0.5 μm 5 μmClass name Volume units Volume units Volume units Volume units Volume units

Sl English m3 ft3 m3 ft3 m3 ft3 m3 ft3 m3 ft3

M1 . . . 350 9.91 75.7 2.14 30.9 0.875 10.0 0.283 . . . . . .M1.5 1 1,240 35.0 265 7.50 106 3.00 35.3 1.00 . . . . . .M2 . . . 3,500 99.1 757 21.4 309 8.75 100 2.83 . . . . . .M2.5 10 12,400 350 2,650 75.0 1,060 30.0 353 10.0 . . . . . .M3 . . . 35,000 991 7,570 214 3,090 87.5 1,000 28.3 . . . . . .M3.5 100 . . . . . . 26,500 750 10,600 300 3,530 100 . . . . . .M4 . . . . . . . . . 75,700 2,140 30,900 875 10,000 283 . . . . . .M4.5 1000 . . . . . . . . . . . . . . . . . . 35,300 1,000 247 7.00M5 . . . . . . . . . . . . . . . . . . . . . 100,000 2,830 618 17.5M5.5 10,000 . . . . . . . . . . . . . . . . . . 353,000 10,000 2,470 70.0M6 . . . . . . . . . . . . . . . . . . . . . 1,000,000 28,300 6,180 175M6.5 100,000 . . . . . . . . . . . . . . . . . . 3,530,000 100,000 24,700 700M7 . . . . . . . . . . . . . . . . . . . . . 10,000,000 283,000 61,800 1750


Appendix A4.1: A BriefSurvey of the MainMetallization Techniques

The four main techniques that are used forapplying metal coatings to metallic and nonme-tallic materials are as follows.

Physical vapor deposition (PVD) embracesall methods where the coating material is physi-cally converted into a vapor and then made tocondense onto the surface of the substrate with-out undergoing any fundamental chemicalchange in the process. The various methods aredistinguished by the means used to generate anddeposit the vapor of the coating material.

Vacuum evaporation covers those methodswhere the source material is thermally vapor-ized. This is commonly accomplished by resis-tance heating or by electron beam bombardment.

In sputtering, by contrast, material on the sur-face of a solid target is vaporized by bombardingit with inert gas ions, accelerated by a potential of500 to 5000V.Aglowdischarge in a low-pressureatmosphere of the inert gas—either self-sus-tained, as in cathodic sputtering, or supportedthermionically, as in triode sputtering—is nor-mallysetupfor thispurpose.Therateofsputteringmay be increased by magnetically intensifyingthe glow discharge, as in magnetron sputtering.Reverse bias sputtering or fast atom bombard-ment is normally available as abuilt-in facility forcleaning of the substrate surfaces immediatelyprior to the sputtering operation. This can consid-erably enhance the adhesion of coatings.

Where the deposition process takes advantageof the ionized fraction of the condensing vapor,the process is described as ion-aided [Martin1986]. Ion plating and ionized-cluster beamdeposition exemplify two techniques based onthis principle. Instead of generating a dischargearound a target, energetic ion beams may beaimed directly at the surface of the substratewheneither a surface coatingwill be obtained, as in ionbeam deposition, or embedded within the sur-face, as in ion implantation, at higher incidentenergies. Because ion plating can develop thickdeposits of high density and purity, it is gainingpreference over evaporation as the vapor phaseprocess of choice for applying solders to sub-strates.

Chemical Vapor Deposition. The depositionof a coating by means of a chemical reaction oc-

curring from a gaseous phase on or immediatelyadjacent to the surface of a substrate is known aschemical vapor deposition (CVD). The substrateis usually heated to generate the reaction.

Chemical vapor deposition may be classifiedaccording to the type of chemical reaction in-volved. In a decomposition reaction, a gaseouscompound AB may be decomposed into a solidcondensate A and a gaseous product B whenplaced in contact with a colder substrate. If thecompound AB instead dissociates into a solidphase A and a gas phase AB2, say, then the CVDprocess is referred to as one involving a dispro-portion reaction. Oxidation and reduction of ha-lides constitute the two other types of reactionthat are widely employed.

Wet plating of metallic layers encompassesprocesses where coatings are deposited on a sub-strate through immersion of the substrate in aliquid, usually aqueous, containing the appro-priate metallic ions. The deposition often func-tions by ionic discharge, with the metal depos-ited onto an electronegatively charged,conductive substrate (cathode). The plating pro-cess can introduce organic compounds into themetal coatings, although these can often bemini-mized by judicious choice of the bath formula-tion.

Thin coatings can be grown autocatalytically(i.e., without an applied electric field) through areduction of metal ions in the plating bath by theimmersed substrate. This process is known aschemical displacement and also as immersionplating. Another autocatalytic method, com-monly referred to as electroless plating, involvesthe deposition of metal from a plating bath con-taining the metal ions together with a reductant.This process differs from chemical displacementin that no significant reaction occurs within thevolume of the liquid, and the depositing metalcatalyzes further deposition, so that thicker filmscan be grown. It is usually necessary to activatenonmetallic substrates by chemical treatment forthem to generate the catalytic reaction. Nonme-tallic elements, principally, phosphorus and bo-ron, tend to be incorporated into the metalliccoating from the reductant.

Thick-film formulations usually comprise aslurry, containing the metals or metal com-pounds and sometimes a glass in an organic car-rier, which is intended to be applied by paintingor screen printing onto the desired areas. Sub-sequent firing drives off the organic fraction andstabilizes the metallization by producing a dif-fused interface with the nonmetal substrate. It is


usual practice to apply and fire each thick-filmmetallization separately, although processes havebeen developed whereby at least two thick-filmlayers are fired together. Common thick-filmmet-allizations are discussed as follows.

Systems based on reactive metals (zirconium,tungsten, titanium, manganese, molybdenum).These formulations are fired at approximately1600 °C (2900 °F) in a reducing atmosphere. Be-cause of the relatively refractory nature of the re-sulting metal surface, either a strongly reducingenvironment is required to effect subsequentwet-ting by solder, or a wettable metallization shouldbe applied on top.Alternatively, the wettable sur-face layer may be applied over the reactive metallayer and the two layers fired together. An ex-ample is a tungsten-loaded frit overcoated with anickelpaste,whichiscofiredat1300°C(2370°F).Even at this temperature, the interdiffusion be-tween the two metals is slight, so that a discretelayer of nickel forms on the surface after the heat-ing cycle [Kon-ya et al. 1990].

Systems based on noble metals (copper, silver,gold, palladium, platinum). These materials arefired between 850 and 950 °C (1560 and 1740°F). The silver, gold, and platinummetallizationscan be fired in air, while a reducing atmosphereis generally required for the less noble metals.

Metal-loaded glass frits are fired on the sur-faces of components above 400 °C (750 °F) toform a glaze that is strongly adherent to the non-

metal substrate. The concentration of the glass atthe interface with the component means that theouter layer of the coating is sufficiently metallicin character for it to be electroplated or directlysoldered.

Thick-filmmetallizations are supplied as com-plex proprietary formulations and are availablein different physical forms, each tailored for alimited range of substrate materials. It is advis-able to consult the supplier on their conditions ofuse and likely properties.

As might be expected, there are advantagesand disadvantages associated with the differentmetallization techniques. Vapor deposition isgenerally superior to wet plating in offering bet-ter control of impurities and reduced porosity inthin coatings. Wet plating, by comparison, tendsto be faster and cheaper and can provide thickercoatings. Thick-film metallizations may be morereadily applied to selective areas and are moretolerant of substrate topology. The technique thatwill normally be chosen will be the one bestsuited to the particular application on the groundsof its fitness for purpose and cost. A brief com-parison of the characteristics of the principalmethods used for applying metallizations, to-gether with those of the coatings that they arecapable of producing, is presented in Tables 4.9to 4.12. It must be pointed out that the entries inthe tables represent the general situation; par-ticular casesmight be out of the ranges indicated.

Table 4.9 Techniques for applying metallizations: characteristic features

ProcessMetallic materials

capable of depositionSuitable

substratesThrowingpower

Filmthicknessachievable

Filethicknesscontrol

Throughputof process

Vacuum evaporation Elemental metals andsome alloys

Mostnonvolatilematerials

Line-of-sightprocess

nm-μm Good Low, batch

Sputtering Wide range ofelemental metalsand alloys

Mostnonvolatilematerials

Moderate(function oftarget size,gas pressure,and target-substratedistance)

nm-μm Excellent Low, batch

Chemical vapordeposition

Elemental metals Materials thatcan withstandthe hightemperaturesrequired

Good μm-mm Good, butneed tostringentlycontrolseveralprocessvariablessimultaneously

High, manyitems at atime; batchorcontinuous

Electroplating Elemental metalsand some binaryand ternary alloys

Electricalconductors

Moderate μm-mm Generallyless precisethan forvapordeposition

Very high,can becontinuous

Electroless plating Elemental metals anda few binary alloys

Wide rangeof materials

Good μm Generallypoor

High, can becontinuous

Thick film Wide range ofelemental metalsand alloys

Materials thatcan withstandthe firingtemperatures

Physical accessto surfaces isrequired

μm-mm Moderate High, batchor conveyorbelt


Further information on these metallizationtechniques can be found in the literature, whichis extensive. A comprehensive review that cov-ers the chemical bonding, chemical reaction,

interface structure and properties of metal/ceramic interfaces, property measurement, andtheir fracture behavior is given by Howe[1993]. Other useful sources of information are

Table 4.10 Metallization techniques; relative meritsProcess Advantages Disadvantages

Vacuum evaporation Relatively simple equipment required forresistance heating evaporation, which issuitable for coatings of mostelemental metals

Not suitable for alloys that have constituents withgreatly differing vapor pressures. Meticuloussubstrate cleaning prior to deposition is required.

Sputtering Possible to coat a wide range of compositionsDense coatings and good adhesion obtainable

Requires sophisticated equipment. Lowthroughput. Heating of substrates and lowdeposition rates in conventional diode or triodesputtering


High-quality coatings are obtainable. Outputis generally high.

Equipment is sophisticated and is usually specificto particular coatings.

Electroplating High throughput. Large areas can be coatedwith uniform thickness. Limited onlyby the size of the plating bath. Relativelyeasy to control

Chemical handling, vapor, and effluent problems.Film impurities and imperfections can also presentproblems. Can only apply coatings to electricallyconductive materials. Thorough cleaning andchemical activation of substrates are requiredprior to plating.

Electroless plating Large areas can be coated with uniformthickness.

Good throwing power. Only very basic equipmentis required.

As above, except that nonconducting materials canbe plated. Range of available coatings isrestricted.

Thick film Requires simple equipment. Lends itself to highvolume production using screen printing andfiring in belt furnaces

Relies on manufacturers’ proprietary formulations.Relatively high process temperatures are used.Only thick films can be applied by this technique.

Table 4.11 Metallization techniques: important process parameters

ProcessRate of deposition,

μm/minPressure in deposition

chamber, mPaSubstrate temperatureduring coating process

Vacuum evaporation 0.001–5 0.01–10 Substrate is often heated to 200 °C (390 °F)to promote adhesion

Sputtering 0.005–1 100–10,000 Mostly below 100 °C (212 °F)Chemical vapor deposition 5–100 10,000–100,000 200–2000 °C (390–3630 °F), but usually

400–800 °C (750–1470 °F)Electoplating 0.1–100 Ambient 10–100 °C (50–212 °F)Electroless plating 0.1–1 Ambient 10–100 °C (50–212 °F)Thick film 1,000–10,000 (does not

include firing times)Ambient 400–1800 °C (750–3270 °F)

Table 4.12 Metallization techniques: coating quality

ProcessCoating thickness

uniformityCoating

continuityCoatingpurity

Coating adhesion tosubstrate

Vacuumevaporation

Variable; determined bysource-substrategeometry

Moderate to lowporosity

Purity limited by sourcematerials anddeposition atmosphere

Fair

Sputtering Higher uniformitypossible than forvacuum evaporation

Low porosity Purity limited by sourcematerials anddeposition atmosphere

Generally excellent


Good uniformitypossible; depends ondesign of thedeposition chamber

Dense and essentiallypore-free

Purity is that of thestarting materials oreven better

Variable; dependenton materials andprocessing conditions

Electroplating Good uniformity onflats, nonuniform atedges

Susceptible to porosityand blistering

May incorporate saltsand gaseous inclusions

Variable; often excellent

Electroless plating Fair uniformity Susceptible to porosityand blistering

May incorporate saltsand gaseous inclusions

Variable; often excellent

Thick film Variable Dense coatings areachievable

Often contain glass andpossibly organicresidues

Variable; dependent onmaterials andprocessing conditions


given in the selected bibliography appended tothe preface.

Appendix A4.2: Critique ofVoid-Free SolderingStandards

In the attachment of semiconductor die to sub-strates by soldering, it is frequently a require-ment that the resulting joint be free of voids. Itis generally true that voids in the joint betweenan active device and its heat sink will result inhot spots and the premature failure of the com-ponent. Target criteria for the inspection of jointsare defined by standards.

Current inspection standards for void-freejoints are specified in Military Standard (MIL-STD) 883D,Methods 2010.10 and 2012.7. Thesestandards are echoed in British Standard (BS)9450 and European Space Agency (ESA) speci-fications 20400, 2045000, 4045000, and2099000. Taken together, they require the fol-lowing to be met:

• >75% of the perimeter to exhibit a fillet• >75% of each side to exhibit a fillet• Two or more sides to exhibit complete fillets• <50% voids in the joint• No single void to traverse either the length or

width of the joint, or to be >10% of the jointarea

Although these standards provided well-defined criteria for judging an assembly process,they are not well suited for the purpose of as-sessing void-free joints. In particular, there is noexternal indication of dewetting in the joint gap,and 49%voids is arguably a somewhat high levelfor a joint that is judged to be void-free.

The aforementioned inspection criteria re-late mostly to visual inspection of the joint andtherefore its external appearance. To properlyassess whether or not a joint is free of voidsrequires inspection of the joint interior. This isoften possible using scanning acoustic or x-raytechniques (see Chapter 5, section 5.10). Stan-dards exist that define what constitutes a de-fect when voids are detected using these in-spection methods. For example, ESA x-ray

Standard (2099000) specifies an unacceptablejoint as one containing:

• Voids in excess of one-half of the total planarea

• A single void equal to the length of the joint• A single void that traverses the width of the

joint

Again, these criteria permit a joint to containlarge voids that potentially can be responsible forcatastrophic failure yet is able to pass inspection.

For application where void-free joints are es-sential to the functionality and reliability of theproduct, proposed inspection criteria are as fol-lows:

Visual Inspection. A sensible inspection cri-terion for void-free joints that relies only on vi-sual examination is the presence of a completeand uninterrupted fillet around the joint periph-ery. Although this condition does not guaranteezero voids within the joint, it does at least pro-vide assurance that the external surfaces achievedand maintained wettability during the processcycle. It also indicates that the joint is leaktight.

According to the same rationale, it is desirablethat the fillets are smooth, shiny, and have lowcontact angles. A smooth and shiny solder sur-face is an indicator of the absence of an oxideskin. Any blemishes, such as nonuniform hue,discoloration, bumps, cracks, craters, or foreignmaterial, give cause for concern as to the integ-rity of the hidden volume of the joint. Low filletcontact angles are desirable as an indicator ofgood wetting, but this inspection criterion is notusually available because solder spread often isconfined by lands, causing the contact angle tobe higher than it would be if the solder were ableto spread freely.

X-Ray or Scanning Acoustic Microscope(SAM) Inspection. The most demanding targetthat can sensibly be set for joints larger than afew millimeters per side is <5% voids by planarea. A truly void-free joint is not seen to be anachievable target in volume production, not leastbecause even state-of-the-art assessment tech-niques can lead to ambiguity in identifying avoid, especially when using thin, predepositedsolder and thick components.

For each application, it is should also be pos-sible to define a maximum acceptable void size(either by area or linear dimension) in variousareas of the joint. This figure can be obtainedfrom thermal and physical modeling. Void frac-tion, maximum void size, and their impingementinto areas where solder is essential can be gen-


erated automatically by modern x-ray and SAMequipment in real-time.

The bottom line in devising an inspection strat-egy is to define a set of criteria that can offerassurance of product integrity and thereby add toits value and reputation for reliability.

Appendix A4.3: Drynessand Hermeticity of SealedEnclosures

A frequent requirement of electronics, optics,and microelectromechanical systems packagingis to provide an enclosure that is dry and resistantto ingress of moisture during the lifetime of theproduct. Attainment of such enclosures is gen-erally referred to as hermetic packaging.

The first step toward meeting this objective isto choose a suitable material for the packageconstruction. No material is truly hermetic tomoisture, although glasses, ceramics, and metalsare obviously superior to any plastic material.Figure 4.38 provides an indication of the relativeranking of different classes of materials tothrough-thickness moisture penetration, whichaccounts for the choice of metal and ceramicpackages for the most demanding applications,and polymeric packaging solutions for less du-rable and less cost-tolerant consumer products.

The contents of the package need to be fixedin place, using materials that will not later outgas

and contaminate the package atmosphere. Sol-der, particularly when used fluxless, is ideal forthis task. Gold-silicon solder, which is frequentlyused for semiconductor die attach applications,acts as a desiccant after the package is sealed,because the silicon in the alloy reacts with watervapor. Tests show that preforms of this materialreduced themoisture content in the package from15,000 ppm for control samples to 125 ppm [Car-ley, Nearhoff, and Dennin 1984]. The hydrogencontent in the package was found to increasefrom 1800 to 7000 ppm, implying that the likelyreaction taking place is:

Si � 2H2O → SiO2 � 2H2

Adhesives have been formulated that do notgreatly outgas. However, caremust be takenwiththeir use in sealed enclosures, because they areonly stable in this regard provided they are notheated beyond the curing temperature. For ex-ample, if a semiconductor die is attached usingan epoxy adhesive, but the package is then sealedusing a high-temperature solder, the adhesivewill liberatemoisture as the sealed package coolsfrom the lid-seal process temperature.

Next, the package needs to be dried before itis sealed.Water adsorbed ontometals needs to beheated to 115 °C (239 °F) in order to force it todesorb. Onewidespreadmisconception is the be-lief that a low moisture reading on a sealing boxhygrometer will ensure a dry ambient inside thepackage. The fallacy arises because the majorsource ofmoisture is that adsorbed onto the pack-age walls, and thus, the dry box reading has nocorrelation with the resulting package ambientatmosphere. Removing the water of hydrationalone is not sufficient. Although this moisturewill desorb at 115 °C (239 °F), it is not until ametal surface is heated above approximately 160°C (320 °F) that the absorbed hydroxide specieswill convert to metal oxide, to the accompani-ment of further evolution of water [Swartz et al.1983], according to the reaction:

M(OH)2 → MO � H2O

where M represents the metal.Figure 4.39 shows the moisture content in a

metal package as a function of the bakeout timeand temperature. Low residual moisture contentis obviously favored by high bakeout tempera-tures sustained for long periods.

Fig. 4.38 Predicted time for moisture to permeate variouspackaging materials in one geometry. Adapted

from Traeger [1976]


A problem frequently encountered is that thepackage contents or assembly method do notpermit the use of elevated temperature for thebakeout, while manufacturing economics are notcompatible with extended bakeout times at a re-duced temperature. One solution to this problemis to place the package in a sealed chamber and, atthe highest permissible temperature, repeatedlysaturate theatmospherewithdrynitrogengas,andthen pump it out. Nitrogen gas obtained from acryogenic source and conveyed by stainless steelpipework will have a moisture content of 2 ppmor less. The moisture on the package walls willequilibrate with that in the atmosphere, so thatwhen the nitrogen gas is removed, it will takemoisture with it. Twenty such cycles, conductedover a period of a fewhours,will reduce themois-ture content of the package to a few parts per mil-lion. This approach has the additional advantageof reducing the time that the package is exposedto vacuum, which can result in hydrocarbon con-tamination of surfaces arising from backstream-

ing of vacuum pump oil vapors. Note that thepackage must be sealed while still in the bakeoutchamber because, if it is exposed to air, the sur-faces will instantly become saturated with wateragain, even if they are kept hot.

Having hermetically sealed the package, itmight be expected that the contents would nowbe permanently protected from the atmosphereoutside. This is not true because, in practice,hermeticity can only be specified in terms of afinite leak rate. Small gas molecules will enterthe sealed cavity by diffusion or permeation un-til, ultimately, equilibrium with the atmosphereoutside is reestablished. Package leak rates aremeasured using helium as a tracer gas, and theacceptable leak rate is defined for the specificapplication. For example, a package having aninternal volume of 100 mm3 and a leak rate of 5 107 Pa · m3/s (4 10�8 ft · lb/s) will have atheoretical water buildup rate of roughly 104 ppmper day, indicating that the package atmospherewill be in equilibrium with the air in a matter ofdays. At a leak rate of 10�9 Pa · m3/s (8 10�11

ft · lb/s), equilibrium will be achieved within ayear. Figure 4.40 shows predicted moisture pen-etration rates for a helium leak rate of 10�8 Pa ·m3/s (8 10�10 ft · lb/s) for ambient air at 70%relative humidity [Stroehle 1997].

Ahelium leak rate of 10�9 Pa · m3/s (8 10�11

ft · lb/s) is generally considered to be a minimumacceptable level of hermeticity, and 10�11 Pa ·m3/s (8 10�13 ft · lb/s) to provide adequateprotection against moisture ingress for all but themost sensitive components. The latter leak rateis also the lower detection limit of low-cost he-lium leak check equipment, while 10�14 Pa · m3/s(8 10�16 ft · lb/s) is about the limit of spe-cialized laboratory systems.

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Fig. 4.40 Calculated ingress of water vapor into a hermeti-cally sealed package having a leak rate of 10�8 Pa

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CHAPTER 5

Advances in Soldering Technology

To many readers, this chapter is something ofa disappointment, because it does not containany sparklingly new solders, fluxes, metalliza-tions, processes, or diagnostic tools. The reasonfor this is quite simply that, as far as the authorsare aware, there have been no significant com-mercial developments in the 10 years since thefirst edition of Principles of Soldering and Braz-ing went to press. Even the “hot topic” of the1990s—lead-free solders—is based on alloys thatwere known and largely characterized previ-ously. Soldering is a relatively mature technol-ogy, so that while there has been progress andmany undoubted improvements made in recentyears, these have tended to be evolutionary ratherthan revolutionary in nature. Composite anddoped solders are, perhaps, some of the few ex-amples to fall into the latter category, but they arestill a long way from becoming a commercialreality. Indeed, one might even argue that withmodern health and environmental awareness,there has actually been a contraction in the num-ber of soldering materials in general industrialuse in recent years. With the proscription of ma-terials recognized to pose health hazards, cad-mium- and lead-containing solders along withnatural rosin-based fluxes, volatile organic com-pounds, and cleaning agents containing chloro-fluorocarbons have either disappeared or are inthe process of being removed from manufactur-ers’ catalogs.

This chapter endeavors to present a number ofmaterials and processes, none of which are en-tirely new but by virtue of changing industrialneed or technical innovation are likely to berather different from the processes the readermay have been aware of some years ago. Theprimary example is, of course, the knowledgebased on lead-free solders. The others are a rep-resentative selectionmade by the authors and feltto be useful or having appeal as laboratory curios

awaiting industrial application but, inevitably,may not accord with some readers’ preferences.To some extent, the newmaterial also is intendedto fill gaps in the previous edition of Principlesof Soldering and Brazing or to provide supple-mentary detail to the preceding chapters. This isthe reason for the inclusion of sections devotedto flip-chip processes, diffusion soldering, andmodeling. Scanning acoustic miscroscopy(SAM) and fine-focus x-ray techniques are in-cluded because of the considerable technical ad-vance that has been made with these diagnostictools in recent years, coupled with the dramaticreduction in price of the equipment. It is now byno means uncommon to see ranks of high-volume production lines, each with a dedicatedSAM and microfocus x-ray system at the end,performing 100%product inspectionwith totallyautomated defect-recognition software. Alliedwith these facilities are image-recognition sys-tems that read individual chip, resistor, and ca-pacitor values and check that printed circuitboards (PCBs) have been populated with thespecified components in the desired orientation.The cumulation of these advances is industriessuch as the mobile phone market, where totalhandset production now exceeds 2 billion units,achieved in less than 20 years.

5.1 Lead-Free Solders

The literature on lead-free solders is almostoverwhelming. A computer search on the termlead-free quickly reveals tens of thousands oftechnical papers and references to conferencepresentations on the subject. This depository ofknowledge is augmented by entire issues of jour-nals, chapters in several books, and entire books[e.g., Hwang 2001]. Thus, there ismuch valuable



technical information on the metallurgy, physi-cal and chemical characteristics, and reliabilityof joints made using lead-free solders. Extensiveas it is, these collective works often do not ad-equately describe the process window for a par-ticular joining problem of interest nor predict thereliability of the resulting joints. This is, perhaps,not surprising, given such key information is stillnot immediately available for lead-tin solder, de-spite many years of endeavor. Many of the lead-free alternatives are technically more complexmaterials.

The principles of lead-free soldering are notfundamentally different from those of lead-tinsoldering. Lead-free soldering processes are nowa commercial reality, and within very few years,it is likely that the vast majority of electronicsand optoelectronics products will be manufac-tured using this new technology.

5.1.1 The Drive forLead-Free Soldering

The drive for lead-free soldering for mass-market electronics assembly first came to promi-nence in the late 1980s with Senator Read’s billin theUnited States Senate [U.S. Senate bill S3911990; U.S. Senate bill S729 1993]. This billsought to address the concern for human healthposed by the rapidly increasing quantities of dis-carded electronics equipment accumulating inlandfill waste disposal sites. Although batterymanufacture accounts for the vast majority of the5.5 million tonnes (6.06 million tons) of leadconsumed each year, battery recycling is almost100% effective. However, virtually none of the60,000 tonnes (66,000 tons) of lead used by theelectronics and lighting industries is recycled.The waste in landfill sites is subject to chemicalattack by rainwater, fromwhere leached-out con-stituents, including lead, eventually find theirway into drinking water supplies [Smith andSwanger 1999].

In 1998, the EuropeanUnion introduced a draftdirective, which is a precursor to law, entitled“Waste Electrical and Electronic Equipment”(WEEE) [The European Commission 2000; TheEuropean Commission 2001]. This followedSenator Read’s initiative and called for a ban onlead in all electronics, except for automotive use,by 1 January 2004. The WEEE directive wasintended to ban the selling, importing, and ex-porting of electrical/electronic equipment con-taining lead. The extent to which this directivewas politically motivated may be the subject of

informed debate, but it is noteworthy that in thepreceding January, the Japanese Electronic In-dustry Development Association (JEIDA) andthe Japanese Institute of Electronic Packaging(JIEP) presented a roadmap to totally lead-freetechnology by April 2001.

The combined effects of these proposals wasto put pressure on the rest of the world to followsuit and generated a large investment in solder-ing process, equipment, and materials develop-ment. Many tens of thousands of hours werecommitted to research and development of lead-free solders during the 1990s. Although the ef-forts were somewhat uncoordinated, the conclu-sions of these studies conducted across the globewere remarkably similar: Lead-free solderingwas seen to be a technical possibility, and indeed,furtherwork has resulted in the commercial avail-ability of lead-free solders and lead-free elec-tronics products [IDEALS 1999].

The threat of legislation has now receded.Senator Read’s bill was withdrawn, and the Eu-ropean Union has greatly expanded the categoryof exceptions and pushed back the implementa-tion date to 2006 or 2008 (the date is not firmlyfixed). Nevertheless, most responsible compa-nies now have accepted a commitment to cleanand “green” manufacturing, whereby no newelectronic products may contain lead. The au-thors are of the opinion that there will be a pro-gressive transition to lead-freemanufacturing forelectronics products over the next decade or so,driven largely by companies wishing to promotean environmentally aware image.

The majority of the work on lead-fee soldershas concentrated on the need to find a replace-ment for lead-tin near-eutectic alloys. While thisendeavor has been successful, no alternatives tothe high-melting-point, lead-rich family of al-loys with melting points of approximately 300°C (570 °F) have been identified. In recognitionof this fact, the pending legislation exempts high-melting-point solders. Section 5.1.8 of this chap-ter outlines some of the reported attempts to findreplacements for these alloys.

One positive benefit of the switch to lead-freesolders is elimination of the need to sourcelow-alpha-emission solder for electronics appli-cations. Lead, being a heavy metal, contains aproportion of radioactive isotopes, and the al-pha particles emitted by spontaneous decompo-sition can cause electronic circuits to generatespurious digital signals. Although lead is nomi-nally cheap, being only an eighth of the cost oftin, certified low-alpha lead is much more ex-


pensive. The price of low-alpha lead roughlydoubles with every tenfold reduction in thecounts per hour of emitted particles, so its re-moval from premium-quality solders representsa potential cost saving.

5.1.2 Compatibility withLead-Tin Solder

Lead-tin eutectic solder has many desirablecharacteristics, namely, excellent wetting andspreading behavior and a relatively advanta-geous range of physical properties. Moreover,because hitherto the entire electronics and light-ing industries had developed around this alloysystem, all associatedmaterials, components, andsubsequent processes were developed to be com-patible with its properties and limitations. Theinitial goal of the work on lead-free solders wasto find a drop-in replacement for lead-tin thatcould be used at the same process temperature,on the same equipment, with the same substratematerials and fluxes, and could deliver compa-rable performance and reliability. For the reasonsthat are elucidated, it is now recognized that thisis not technically possible. Instead, there nowexist several families of lead-free solders, eachsuited for a different group of applications andcompatible with only a limited subset of metal-lizations and processes. A further consideration,of which any practitioner of soldering needs tobe aware, is that many of the lead-free solders areincompatible with lead-tin solder. This willpresent difficulties for many years ahead, be-cause an attempt to rework or repair a lead-freejoint with lead-tin solder, or vice versa, can resultin a joint with severely compromised function-ality and reliability. This problem will be furtherexacerbated, because it is not easy to distinguisha lead-tin joint from one made with lead-freesolder without resorting to analytical techniques.

5.1.3 Alternatives to Lead-Tin Solder

In the mid-1990s, it was estimated that theelectronics industry uses approximately 60,000tonnes (66,000 tons) of lead-tin solder each yearand the lighting industry somewhat more. This isused to make approximately 1013 soldered jointsper annum [Vincent and Humpston 1994]. Forany alloy to be a worthwhile solder for the elec-tronics industry, it must meet specific qualitiesunder the following criteria:

• Melting range: An alternative solder musthave an upper process window that is suffi-ciently low in order that existing componentsand boards are not damaged by the solderingprocess. Inpractice, thismeans apeakprocesstemperature of 260 °C (500 °F), so that the liq-uidus of the alloymust therefore be somewhatlower.On theother hand, the soldermust havea solidus that is sufficiently high that the jointsare robust in service, and its usemust be com-patible with existing step-soldering opera-tions. This means that the solidus, after alloy-ing with any metallizations present on theboardandcomponent leads,shouldnotbesub-stantially below 170 °C (340 °F).

• Physical and chemical characteristics: Thenew lead-free solder must wet and spread onthe common engineering metals and metal-lizations, namely, gold, silver, platinum, pal-ladium, nickel, tin, copper, and iron. Ideally,the solder should also be compatible withexisting flux technology and certainly not re-quire a more aggressive or environmentallyadverse material. The solidified solder alsomust be sufficiently inert to resist the corro-sive environments associated with electronicequipment (for example, the electronic con-trol boards inside domestic washing ma-chines and dishwashers are subject to incred-ibly harsh chemical, physical, and thermalenvironments).

• Environmental, health, and safety aspects:The alloy and its components must be non-toxic.This automatically rules out alloys con-taining cadmium, thallium,mercury, and alsopossibly nickel. The intrinsic toxicity of thesemetals means that they cannot be consideredas constituents of replacement solders (Table5.1).Similar considerations apply to the flux for-mulations for use with these new solders andthechemicalsused to remove theflux residuesfrom the assembly after fabrication.A furtherconsideration is that the pollution and envi-ronmental damage arising from the increasedmining and extraction of the alloy ingredientsin response tonewdemandmustbeatminimalcost and environmental damage.

• Economics and availability: For any alloy tobe considered as a potential replacement forlead-tin solder, its constituents must be suf-ficiently abundant so that the needs of themarket for the new solders can be met fairlyreadily. This condition on availability is nec-essary to ensure that the alloy or alloys are

Chapter 5: Advances in Soldering Technology / 191

not subject to supply or price constraints thatwould prevent their widespread adoption.Based on the current world consumption oflead-tin solder, each 1% substitution for lead-tin solder byanother element represents anewannual consumption for that elementof theor-der of 500 to 600 tonnes. Thismeans that sup-ply levels should be higher by severalfold.There are, for example, some excellent alloysthat contain indium, but only 100 tonnes of in-dium are produced each year, andmost of thatquantity is recovered as a by-product of zincextraction.World production of elements po-tentially suitable as constituents of lead-freesolder is listed in Table 5.2.

Taking into account availability, cost, meltingpoint, and environmental issues leaves tin as theonly element on which new lead-free solders forthe mass market can be developed. Because tinmelts at 232 °C (450 °F), alloying additions arenecessary to depress the liquidus temperatureand, preferably, to reduce the cost, because tin isrelatively expensive compared to lead. Gallium

and indium are simply too scarce and too ex-pensive to use for other than niche applications.The situation with regard to bismuth, silver, cop-per, antimony, and zinc is less restrictive, andvirtually all of the lead-free solders availablecommercially are based on tin with selected ad-ditions of these elements.

Almost without exception, lead-free soldersare based on the silver-copper-tin eutectic withsmalloptionaladditionsofbismuth,antimony, in-dium, and zinc. Many of the most promising al-loys have been assigned patent protection. Someof the key compositions are listed in Table 5.3.

Bismuth is often recommended as an additionto lead-free solders because of its beneficial ef-fect on melting point and mechanical properties,as shown in Fig. 5.1 and 5.2.

Users of bismuth-containing lead-free soldersneed to be aware of a potential incompatibilitywhen undertaking repair work or using it to sol-der to lead-tin component or board finishes. Thepresence of lead in the alloy mixture results infinal solidification of the alloy at temperaturesthat can be below 100 °C (212 °F), depending onthe composition and composition gradient in thesoldered joint. Obviously, this could have a cata-strophic effect on the reliability of an assembly,because it would normally be expected that thejoint would remain solid up to at least 150 °C(300 °F).

Lead-free solders are more expensive thanlead-tin solder, because lead is substantiallycheaper than all of the other constituents, andonly zinc comes close to it in price. While thisis important for wave-soldering processes wherean industrial machine will require tens of kilo-grams of solder to fill the solder bath, the pricepremium of lead-free solders over lead-tin solderis minimal, when in the form of dispensable

Table 5.1 Key characteristics of thelower-melting-point metals

ElementMelting

point, °C Comments

Mercury �38.9 ToxicCesium 28.5 Highly reactiveGallium 29.8 Extremely rareRubidium 38.9 Highly reactivePotassium 63.7 Highly reactiveSodium 97.8 Highly reactiveIndium 156.9 Low availability and high cost

(more expensive than silver)Lithium 179.0 Highly reactiveTin 231.9 Nontoxic. Ample supplyBismuth 271.3 Obtained as a minor

by-product of leadThallium 303.5 ToxicCadmium 320.8 Toxic vaporLead 327.5 Toxic

Table 5.2 Availability of potential alloyingelements in lead-free solderMetal World production in 2000(a), tonnes

Silver 17,700Bismuth 5,880Copper 13,200,000Gallium 100Indium 335Antimony 118,000Tin 238,000Zinc 8,730,000

(a) Source: U.S. Geological Survey data. To convert to U.S. (short) ton, multiplyby 1.10

Table 5.3 Composition ranges (wt%) claimedfor some lead-free solder alloys

Tin Silver Copper

Bismuthor

antimony Indium Zinc

88–99.4 0.05–3.0 0.5–6.0 0.1–3.0 . . . . . .90–93.5 2.0–5.0 0.3–2.0 0.5–7.0 . . . . . .92–99 0.05–3.0 0.7–6.0 . . . . . . . . .Bal 3.5–7.7 1.0–4.0 0–10 . . . 0–1Bal 1.0–3.0 0.5–2.0 1–10 . . . . . .Bal 0.1–20 . . . 0.1–25 0.1–20 . . .Bal 0.5–3.5 0.5–2.0 . . . . . . . . .Bal 3.0–5.0 0.5–3.0 0–5 . . . . . .

Note: The obvious overlap in composition ranges makes the legal fortitude ofmany of the patents questionable. However, cross-licensing of nearly all thesilver-copper-tin family of lead-free solder alloys now appears complete, mean-ing that most formulations are readily available from the majority of suppliers.


pastes. The economics of paste production aresuch that they are dominated by the costs ofpreparing clean, spheroidized solder powder inprecise size distributions and those of the spe-cialist organic chemicals that comprise the binderand flux. Thus, while silver-copper-tin eutecticalloy costs approximately five times that of lead-tin eutectic, when based on metal prices, thepremium paid for paste is typically only of theorder of 1.1 to 1.

5.1.4 Silver-Copper-TinTernary Phase Equilibria

Because of the importance of silver-copper-tinalloys to lead-free soldering technology, it is im-perative to have a reliable determination of theconstitution of the silver-copper-tin phase dia-gram. With the aid of modern analytical tech-niques, the phase relationships in this systemhave recently been revisited and precisely de-termined. An excellent publication on this topicis one by Loomans and Fine [1999], which iswell worth consulting as a model on how todetermine a phase diagram using thermal analy-sis and metallography, particularly for a systemwhere the composition of invariant reactions is

cooling-rate dependent. It is now generally con-sidered that the composition for the ternary eu-tectic under equilibrium conditions is 3.5Ag-0.9Cu-95.6Sn, with a eutectic point at 217.2 °C(423.0 °F).A complete liquidus projection of theternary system is given in Fig. 2.14. Substitutionof 0.9% Sn in the binary Ag-96.5Sn alloy bycopper reduces the eutectic temperature by al-most 4 °C (7 °F). Even this small temperaturereduction can be significant for electronic as-semblies with components tailored for use withthe lower-melting-point lead-tin solder, whichcan be detrimentally affected by higher solderingtemperatures.

5.1.5 Metallurgical, Physical, andChemical Properties ofLead-Free Solders

The metallurgy, physical, and chemical prop-erties of lead-free solders are not markedly dif-ferent from other solders. They wet, spread, andalloy with substrate metals in a similar fashion,and all other properties are broadly similar. How-ever, there are small differences and various nu-ances that have been identified that may or maynot be critical for certain processes. Some of thekey parameters are presented in the followingsection, which draws extensively on the reviewby Glazer [1995]. As explained previously, theterm lead-free solders is generally taken to meanalloys based on the silver-copper-tin ternary eu-tectic, although the data used for comparisonalso include other low-melting-temperature lead-free alloys.

In attempting any comparison of solder prop-erties, it must be remembered that for commonsolder alloys, room temperature represents closeproximity to their melting points, in terms ofchemical thermodynamics. Thus, in any test, it isvirtually impossible to standardize factors thatare governed by diffusion. This makes it ex-tremely difficult to compare work by differentinvestigators. While the data presented here mayappear to be quantitative, it is best interpreted assemiquantitative and as an indicator of trends.

5.1.5.1 Surface Tension

The surface tension of tin is either raised orlowered by the addition of other elements in lowproportions. Lead, bismuth, and antimony causea reduction, while silver and copper increase sur-

Fig. 5.1 Reduction in the liquidus and solidus temperature ofan off-eutectic silver-tin alloy as a function of the

bismuth addition

Fig. 5.2 Effect of additions of bismuth on the tensile strengthof bulk samples of silver-tin solder


face energy. Values for surface tension are givenin Table 5.4. Surface tension values of soldersexposed to air are typically lower than in an inertatmosphere, because oxidation lowers the sur-face energy of the molten solder.

5.1.5.2 Other Physical Properties

Data for the density, electrical resistivity, ther-mal conductivity, and coefficient of thermal ex-pansion (CTE) of a few solders are given inTable5.5. Density and thermal conductivity largely fol-low the rule of mixtures. On the other hand,resistivities and CTEs are all closely similar forthe different solders, with the exception of Bi-43Sn. Bismuth is a semimetal, rather than a nor-mal metal, that is present in near-equal propor-tion by volume as the tin phases in themicrostructure of the Bi-43Sn solder, which ac-counts for its higher resistivity and lower thermalexpansivity.

5.1.5.3 Mechanical Properties

Because solders creep rapidly at room tem-perature, measurement of mechanical propertiesis extremely sensitive to strain rate. An addi-tional complication arises because mechanicalproperties are also very sensitive to grain size,

joint thickness, and alloying with the parent ma-terials.

For solder alloys that are composed primarilyof a phase mixture of two pure metals, the elasticmodulus can be estimated from the rule of mix-tures, based on the volume fraction of each phasepresent. Of the common lead-free solder alloys,only the elastic modulus of indium-tin alloyscannot be deduced from data on the respectivemetals but must be measured, because the in-dium-tin eutectic is formed between two inter-metallic compounds of indium with tin.

The yield strength of solders is a strong func-tion of temperature, because thermal activationassists dislocation movement, while testing at ahigh homologous temperature renders the resultsparticularly sensitive to strain rate. However, be-cause solder generally has very lowyield strengthand ismostly used in a temperature regimewherecreep is important, this parameter is often of littlepractical value. More relevant is ultimate tensilestrength, although measurement of this propertyis equally unreliable for the same reasons. Shearstrength normally correlates closely with tensilestrength, except where plastic flow plays a sig-nificant role, as it does in most solders at roomtemperature. Therefore, for all three of theseproperties—yield strength, tensile strength, andshear strength—it is probably fair to deduce thatthe differences between the solder alloys aresmall, with the exception of indium-tin alloys,which are noticeably weaker.

Elongation is an important property, becausehigh ductility enhances the low cycle fatigue re-sistance of a material. Also, reasonable elonga-tion to failure at high strain rates is beneficial inpreventing catastrophic failures if a joint is in-advertently overstressed. The Bi-43Sn and In-49Sn eutectic solders are greatly superior to Pb-62Sn and Ag-96Sn in this respect.

Comparative mechanical properties for se-lected solders are given in Table 5.6.

Table 5.4 Measured values of surface tensionfor binary solder alloys in air and nitrogen

AlloySurface tension, mN/m

Air Nitrogen

Bi-43Sn 319 349Sn-9Zn 518 487Pb-62Sn 417 464Ag-96Sn 431 493Cu-99Sn 491 461Sb-95Sn 468 495

Test conditions: test temperature, 250 °C (482 °F); polytetrafluoroethylene sub-strate; SNMA, synthetic, mildly activated. (RMA-type, rosin, mildly activated)flux

Table 5.5 Physical properties of selectedsolders at room temperature

Solderalloy

Density,kg/m3

Resistivity,μohm · cm

Thermalconductivity,

W/m · K

Coefficientof

thermalexpansion,

10�6/K

Pb-62Sn 8400 14 33 25Ag-96Sn 7360 11 50 26Bi-43Sn 8700 32 21 15In-49Sn 7300 15 34 20

Table 5.6 Comparative mechanical propertiesfor selected solders at room temperature

Solderalloy

Elasticmodulus,

GPa

Ultimatetensile

strength(a),MPa

Shearstrength(a),

MPaElongation(a),

%

Pb-62Sn 39 50 35 50Ag-96Sn 50 60 30 70Bi-43Sn 42 70 30 10In-49Sn 24 20 10 100

(a) Measured at a strain rate of 0.001 s�1


The resistance of metal to creep can be mea-sured in many ways. Creep is a complicated pro-cess, and most solders exhibit at least three typesof creep, denoted as primary, secondary, and ter-tiary. To provide a very rough guide to the rela-tive ranking of different solders in their resis-tance to creep, stress-rupture life is a usefulparameter, because there is a linear relationshipbetween the applied stress and the life of thetestpiecewhen plotted on logarithmic scales. Rel-evant data for some solders at room temperatureare given in Fig. 5.3. Silver-tin solder is the mostresistant to failure by creep, whereas indium-tinsolder is the least resistant among the four sol-ders represented.

The fatigue resistance of different solders isone of the most relevant parameters, from thepoint of view of their application, but also themost difficult to determine reliably. The difficultystems from the fact that the results are extremelysensitive to the specimen configuration and testconditions, and these factors can greatly affecttheir relative ranking. However, there is a gen-eral consensus that fatigue resistance of the foursolders may be ranked as follows: In-Sn < Bi-Sn< Pb-Sn < Ag-Sn. Trials have shown that, forsurface-mount assembly of components on FR4printed circuit board, both silver-tin and copper-tin lead-free solders are superior to lead-tin eu-tectic under most conditions. However, if theapplication involves operation above approxi-mately 150 °C (300 °F), then, surprisingly, thelower-melting-point lead-tin solder actually ex-hibits the superior fatigue performance. This isbecause the almost equally divided and fairly

fine two-phase lamella structure of lead-tin sol-der provides a modicum of strength right up tothe melting point, whereas above 150 °C (300°F), the large crystals of the dominant tin fractionin silver-tin and copper-tin eutectics are less re-silient to fatigue.

5.1.5.4 Corrosion Resistance

Lead-free solders generally do not suffer fromcorrosion in normal environments. The excep-tion is alloys that contain zinc, because of thelarge difference in electrochemical potential be-tween this metal and the other constituents. It isfor this reason that zinc-containing alloys are notconsidered as suitable replacements for lead-tinsolder for most applications. A further disincen-tive to the use of zinc was that, until recently, thewetting and spreading on copper surface wasextremely poor, compared with lead-tin. How-ever, recent work using fluxes containing tin2-ethylhexanoate shows considerable promise.This compound decomposes at the soldering tem-perature to leave metallic tin on the faying sur-face just ahead of the advancing solder front. Bythis means, contact angles as low as 20° can beachieved, which is low enough to be widely ac-ceptable [Vaynman and Fine 2000].

5.1.5.5 Susceptibility toTin Pest and Tin Whiskers

Tin undergoes an allotropic transformationfrom its normal metallic white form, which has

Fig. 5.3 Stress-rupture life of jointsmadewith low-melting-point solders, tested at room temperature. Silver-tin solder ismore resilientthan lead-tin eutectic, while indium-tin alloys are less able to resist creep.


a body-centered tetragonal crystal structure, to apowdery gray material, with a diamond cubiccrystal structure, at, nominally, 13 °C (55 °F)(see Chapter 2, section 2.2). However, in orderto initiate the transformation, it is generally nec-essary for white tin to be subject to strain, and themaximum rate of conversion occurs at approxi-mately �40 °C (�40 °F). Because the formationof gray tin from the white allotrope is accom-panied by a 26% expansion in volume of themetal, the individual grains separate, and the sol-der disintegrates into fine powder. The transfor-mation is termed tin pest. Clearly, with mostlead-free solders having a higher proportion oftin than the lead-tin eutectic alloy they are re-placing, there exists the possibility that tin-richlead-free solders will be vulnerable to failure bythis mechanism. Lead-tin solders are carefullyformulated with controlled impurity additions toprevent the transformation from taking place.This is explained further in Chapter 2, section2.2.

The susceptibility of three pure tin-base sol-ders to tin pest has been studied [Kariya, Gagg,and Plumbridge 2000]. It was found that on stor-age at �18 °C (�0.4 °F), a proportion of thetin-rich phase in silver-tin, silver-copper, and tin-zinc solders transformed to gray tin. The incu-bation period was shortest for the zinc-contain-ing alloy and longest for the silver-containingsolder. Further trials are necessary to establishwhether this allotropic transformation can be sup-pressed by controlled doping, as it can with lead-tin solder (see Chapter 2, section 2.2).

Pure tin, under certain conditions, has a ten-dency to grow long crystals, or whiskers, that cancause electrical shorts between adjacent conduc-tors on PCBs. The elimination of lead from sol-ders applies equally to the solderable finishesapplied to component leads and lands on PCBs.Some of the replacement finishes are tin-base.Therefore, both lead-free solders and some PCBfinishes are potentially susceptible to tin whiskergrowth. Work is believed to be underway to as-sess this vulnerability, and the results are awaitedwith interest.

5.1.6 Process Window forLead-Free Solders

Lead-free solders intended to be as close aspossible form-fit-function replacements for lead-tin solder generally have a melting range of ap-proximately 210 to 230 °C (410 to 446 °F) [Bra-dley, Handwerker, and Sohn 2003]. This is

approximately 30 to 50 °C (54 to 90 °F) higherthan for lead-tin solder. However, the upper tem-perature for soldering of electronics and opto-electronics assembly is fixed at 260 °C (500 °F)by the design of existing parts. The peak tem-perature of 260 °C (500 °F), historically, gave anadequate margin for use with lead-tin solder.Lead-free solders are therefore required to wetand spread with considerably less superheat thanwas previously considered good practice for lead-tin solder. This has two consequences:

• In lead-free soldering processes,much greaterattention must be paid to the thermal cycle.In particular, the preheat cycle needs to becarefully optimized to ensure that all parts ofthe assembly where joining is to be effectedare ultimately raised above the liquidus tem-perature of the solder. This can require con-siderable attention to processing conditionsfor such products as double-sided PCBspopulated with tiny surface-mount compo-nents and massive programmable logic arraychips in ceramic packages.

• Care should be taken to ensure that the com-ponent metallizations and solder are compat-ible. As discussed in the following section,different lead-free solders wet standard met-als and metallizations (copper, tin, nickel,gold, etc.) at different rates, so what workswith one alloy from one manufacturer mightnot be appropriate for an ostensibly similaralloy composition from another supplier.

The reduced superheat also has consequencesfor the choice of flux. At the same time as themove toward lead-free solders, concern was alsogrowing over the damage to the ozone layerabove the Earth, with the principal culprit iden-tified as the family of chlorofluorocarbons. Theprogressive withdrawal of these standard de-fluxing agents led to the reassessment of clean-ing procedures and prompted the development ofno-clean fluxes (see Chapter 3, section 3.2.1.2).A further environmental concern arose over theuse of volatile organic compounds (VOCs). Hith-erto, these were a major ingredient of solderfluxes, and emissions ofVOCs from an industrialwave-soldering machine can be many kilogramsper hour. The VOCs play a major role in thegeneration of photochemical smog.

No-clean and low-VOC fluxes have inher-ently low chemical activity that is further di-minished by the low superheat of lead-free sol-dering processes. This combination results in


processeswith unacceptably narrowwindows, towhich the response has been the widespreadadoption of nitrogen-inerted soldering. The ni-trogen atmosphere retards oxidation of the solderand substrate before the flux becomes active andthereby reduces the work that the flux has to doin maintaining oxide-free interfaces ahead of theadvancing solder front. If large superheats can beapplied, the benefits of a protective atmosphereare reduced, because solder is able to spreadfurther and faster (Fig. 1.14 and 5.4) [Buckley2000].

In recent years, much effort has gone into thedesign of nitrogen-inerted soldering equipment,and remarkably good-quality atmospheres cannow routinely be achieved on open-access sys-tems for surprisingly modest rates of gas con-sumption. Equipment with specifications of 10ppm oxygen, or below, in the working zone isavailable commercially.

5.1.7 Wetting and SpreadingCharacteristics of Lead-FreeSolders

The wetting and spreading characteristics oflead-free solders are not straightforward. Therehave been many studies on this topic, and muchdata exist in the published literature. Under cer-tain regimes, it is possible for lead-free soldersto wet and spread better than lead-tin solder, butfor many it is not. The root of this complexity isthat the wetting and spreading of lead-free sol-ders is influenced by the alloy composition, theprocess temperature, the process atmosphere, theflux, and the substrate metallurgy. This difficultyis exemplified by the data presented in Fig. 5.5and 5.6, which show the time required for thewetting force to reach the acceptance value (2⁄3

of the theoretical maximum) in wetting balancetests for a range of process and materials com-binations [Lau and Ricky Lee 2001]. It is im-possible to discern any consistent trends in thedata.

Fortunately, the manufacturers of the variouslead-free solder alloys have undertaken sufficientdevelopment to know what materials and pro-cess conditions are necessary to obtain wettingand spreading behavior that is at least as good asfor the lead-tin process being substituted. There-fore, when making the change to lead-free sol-dering, it is advisable to adopt one alloy, flux,process window, and recommended componentfinish as a complete and integrated package. It isthe authors’experience that a surface-mount PCBassembly line can be switched from using lead-tin to a lead-free solder in less than 20 min. Thisline was used for the mass production of circuitboards, and the manufacturing yield and in-service reliability obtained from lead-tin andlead-free solderwere indistinguishable. The tran-sition time was essentially that required to swapthe solder paste print cartridges, allow the reflowovens to stabilize with a different temperatureprofile, and change some of the reels on the com-ponent placement machines, because some partscome supplied in a choice of lead-free or lead-tinplating.

5.1.8 High-Melting-PointLead-Free Solders

High-lead solders (>85% Pb) are currentlygranted an exception, subject to review, to theban on lead in solders [European Union Direc-tive 2001]. This dispensation has been grantedon the grounds that there are no technically andeconomically viable alternatives. Lead-rich sol-ders are used in large quantities by the lightingindustry, and most discarded light bulbs are dis-posed of in landfill sites, which provides an in-teresting perspective on the merits of the ban onlead-tin eutectic solder.

Extensive phase diagram modeling suggeststhat the prospects for finding a new multicom-ponent solder with a melting point in the regionof 300 °C (570 °F) are low [Lalena, Weiser, andDean 2002].

Shimizu et al. [1999] have proposed the qua-ternary alloy Zn-4Al-3Mg-3.2Ga as one possi-bility. This alloy exhibits acceptable melting be-havior (309 to 345 °C, or 588 to 653 °F), butthere are difficulties associated with its use be-

Fig. 5.4 Wetting speed of lead-tin solder on copper using arosin-based flux in air and nitrogen atmospheres.

Nitrogen reduces the propensity for the solder and substrate tooxidize and thereby decreases the cleaning action demanded ofthe flux to effect wetting.


cause of the unfavorable combination of highhardness, low ductility, and ease of oxidation.Nevertheless, this solder has a higher thermalconductivity (77 W/m · K) than high-lead sol-ders, lower thermal expansivity (20 × 10�6/K),and, moreover, foil can be produced by hot roll-ing at 250 °C (482 °F). Fluxless attachment ofgold-metallized silicon die to silver-plated cop-per leadframes has been demonstrated with thissolder using a die-bonding machine that pro-vides for a mechanical scrubbing action duringthe heating cycle. The resulting joints, whichmeasured 5 by 5mm (197 by 197mil), were wellfilled and passed standard thermal cycling andwet-high temperature storage tests without dif-ficulty.

Another candidate for a lead-free high-melting-point solder is the ternary alloyBi-11Ag-0.05Ge, which has a liquidus of 360 °C (680 °F)and a solidus of 262 °C (504 °F). Although thewide melting range will be unattractive for some

applications, this alloy does have the merit thatingots can be mechanically processed by hotworking into foil and wire. Molten bismuth oxi-dizes readily in air to an extent that renders thebinary Bi-11Ag alloy unsuitable as a solder. Ger-manium has a higher oxygen affinity than bis-muth and is largely insoluble in the binary basealloy. BecauseGeO sublimes only at 480 °C (896°F), it is reported that at process temperatures inthe region of 400 °C (750 °F), wetting by thissolder can be obtained on nickel, silver, and goldbut not copper, in air, using an inorganic aqueousflux.At room temperature, the alloy is softer thanthe Au-20Sn solder and possesses similar, lim-ited ductility but has a thermal conductivity ofonly 9 W/m · K, which may be restrictive forsome applications [Lalena, Dean, and Weiser2002].

Because of their limitation, both of these al-loys are likely to be of interest only for nicheapplications.

Fig. 5.5 Wetting rate, as measured by the time for the wetting force to reach an acceptable value, for a range of solders on differentcomponent and board finishes. There is no conclusive trend of superiority for any solder or solderable coating, emphasizing

the need to tailor the lead-free process to each situation. OSP, organic surface preservation

Fig. 5.6 Wetting rate, as measured by the time for the wetting force to reach an acceptable value, of solders on different substratesin different atmospheres. The relative ranking varies greatly, depending on the metals and process atmosphere involved.


5.2 Flip-Chip Interconnection

Flip-chip interconnection is a method of mak-ing electrical connection and providing physicalattachment between two components, often asemiconductor die and substrate. Although theprocess was established at the birth of the semi-conductor revolution, itwas largely abandoned infavor ofwire bonding. Flip-chip bondingwas rel-egated toniche applicationsbut nevertheless con-tinued to evolve and develop to a high degree oftechnological sophistication. With the ongoingdrive toward smaller, lighter, and cheaper elec-tronics systems of greater functionality, flip-chipbonding experienced a renascence as an essentialenabling technology. Extrapolation of currenttrends suggests thatwithin ten years, themajorityof semiconductor die may be flip-chip bonded.Flip-chip interconnection is being advocated orused in virtually every advanced packaging con-cept—ball grid arrays (BGAs), chip-scale pack-ages (CSPs), direct chip attach (DCA), and mul-tichip modules (MCMs), to name but a few.

The invention, or, more strictly, the commer-cialization, of flip-chip bonding is credited toIBM Corporation in the early 1960s [Miller1969]. The process was developed at the sametime as the first integrated circuits began to ap-pear, in response to the need to simultaneouslymake a large number of extremely miniaturizedconnections. The technique involved soldering acopper ball at every joint to provide the physicalseparation and electrical interconnection be-tween the transistors on the chip and the sub-strate. Solder-based flip-chip technology cameabout from the realization that the copper ballwas not required and that precise z-axis separa-tion is maintained by the relatively high surfacetension of solder alloys. This modification,known as the C4 process, standing for controlledcollapse chip connection, forms the basis of thetwo flip-chip technologies practiced today. Therenow exist many variants of this process, and onlythe key features are elucidated in the followingsections. For amore complete description of flip-chip technology and methods of its implemen-tation, the reader is referred to Lau [1996, 2000].An overview of flip-chip technology, with par-ticular emphasis on optoelectronic applications,is provided by Lee and Basavanhally [1994].

5.2.1 The Flip-Chip ProcessAschematic illustration of the solder flip-chip

process is given in Fig. 4.19. The two parts to be

joined are first patterned with mirrored, match-ing arrays of a solderable metal. One of theseparts is then further processed to deposit inter-connect metal (solder) on every pad. Either partis then flipped over (hence, the process name)and the metal patterns aligned. Finally, they arebrought together to form a joined assembly.

The flip-chip process flow is set out in greaterdetail in Table 5.7. It is described by reference totwo basic inputs, an electronics wafer and a sub-strate, although either or both of these parts couldbe substituted by other parts of different func-tionality, such as a semiconductor componentpackage and a PCB, as in a μBGA mountingscheme. Processed semiconductor wafers aresupplied coated with a passivation layer, typi-cally either a polymeric or oxide/nitride film, toprovide some protection to the devices duringfinal assembly and packaging. Apertures are cutthrough this passivation layer at discrete pointsto allow contact to the top metal of the semi-conductor fabrication process. This metal tendsto be aluminum or copper on silicon devices andgold on III-V devices, such as gallium arsenide.The substrate will usually have a track pattern ina conductive metal such as copper or gold, withpads at the connection points.

The first step of the flip-chip process involvesapplication of a wettable metal, or underbumpmetal, on the areas of the chip and substrate tobe mated (Fig. 5.7). The deposit is often a mul-tilayer metallization, the functions of which areto provide ohmic contact to and metallurgicalcompatibility between the contacts on the chipand substrate and the interconnect metal. Typi-cally, it comprises a sequence of up to threemetallayers. The first metal is called the foundationmetal, and its role is to provide the initial ohmiccontact and a strong physical bond to the fayingsurfaces of the components. The choice of foun-dation metal varies with the nature of the com-

Table 5.7 Flip-chip process flow

Individual processes differ greatly in their details, but thegeneral scheme is common to both solder bump bonding andcompression bump bonding. The steps shown in parenthesesare optional.

Define underbump metal pattern areasDeposit underbump metals

Define interconnect pattern areasDeposit interconnect metal

Clean and inspect(Fuse interconnect deposit to form alloy)

Align mating partsBond

(Underfill)Inspect


ponents being joined and the method of depo-sition. Common examples are titanium andchromium, if applied by a vapor-phase tech-nique, and zinc, if applied by wet plating. Thesecond metal is the barrier metal and is includedto prevent metallurgical reaction between the ad-hesion metal and the interconnect metal. Thechoice of barrier metal is extremely wide andincludes both pure elements, such as platinum,nickel, and copper, as well as mixtures of metals,such as titanium-tungsten (oxy-nitride) and tita-nium nitride. The barrier metal needs to be ofsufficient thickness to ensure that it remains in-tact during themaking of the joint and the servicelife of the assembly, while remaining electricallyconductive. The former requirement means thatthis barrier layer must be able to satisfy anydissolution into the molten solder and interme-tallic growth in the solid state without being de-pleted. Finally, it is common practice to com-plete the underbump metal with a sacrificialmetal, which is often gold, so as to maintain thesolderability shelf life and/or aid visual inspec-tion. The underbump metal typically has a totalthickness of less than 1 μm (40 μin.).

One of the two parts is removed from theprocessing chain at this juncture, because it isusually neither cost-effective nor often necessaryto have interconnect metal on both parts. Theinterconnect metal is mostly applied by screenprinting or electroplating, although for very smallbumps (<10 μm, or 400 μin., in diameter), vapor-phase techniques are generally necessary. His-torically, lead-tin alloys and pure indium wereused for the interconnect metal, but the processworks equally well with lead-free compositionsand other solders [Lau 1996]. Finally, the com-ponents are cleaned, ready for bonding. A num-

ber of direct deposition processes have been de-veloped for applying the interconnect metal. Oneof the more interesting of these uses a jettingsystem to generate a stream of precisely sizedmolten droplets. These are steered by an elec-trostatic deflection system, and an exact numberof droplets can be deposited at any location overthe entire area of a 6 in. diameter substrate (Fig.5.8) [Priest, et al. 1994; Hayes, Wallace, andBoldman 1996].

The interconnect metal is frequently an alloy.Sometimes, for ease of process control, the sol-der is deposited as sequential layers of the con-stituent metals. It is then common practice toperform an additional fusion step immediatelyafter deposition, whereby the part is heated untilthe layers constituting the solder melt and alloy,forming a hemispherical bump in the process.This opens the possibility of an interesting pro-cess variant, called the wet back process,whereby the diameter of the interconnect metalareas, as deposited, can be up to twice that of theunderlying wettable metal pad. During fusion,surface tension forces pull the molten dropletback to the wettable pad, thereby increasing theheight of the bump while limiting its width. Thistrick simplifies the creation of particularly tallbumps. By making any tracks connected to thebump pad sufficiently narrow, no masking orother measures are required to prevent lateralsolder flow along the tracks.

The interconnects are not by any means re-stricted to a spherical geometry. Indeed, ad-vances in photoresist technology make it pos-sible to fabricate interconnects with aspect ratiosof 5 to 1, even at tiny feature sizes. One pub-lished example shows solder columns 5 μm (200μin.) in diameter and 20 μm (800 μin.) high at 10μm (400 μin.) pitch [Yamada et al. 1998]. Thecolumn grid arrays so formed have substantiallyimproved fatigue life compared with ball gridarrays, because the thermal expansion mismatchstrain between the abutting components is dis-tributed over a greater height ofmaterial.Athree-fold improvement in fatigue life is thereby re-alized [Sturcken 2000].

During the bonding step, the mating intercon-nect areas on the two components are aligned inplane, translation, and rotation and brought intocontact. The connection is made by either coldcompression welding, commonly referred to asindium bump bonding, or by heating to melt theinterconnect metal, known as solder bump bond-ing (Fig. 5.9). Although flip-chip compressionbonding and flip-chip solder bonding superfi-

Fig. 5.7 Schematic illustration of a silicon semiconductor de-vice prepared for flip-chip bonding. A hole has been

cut through the passivation layer, and the three-layer underbumpmetal applied. The foundation metal, M1, is followed by a barriermetal. The interconnect metal has been deposited on top andfused to form a hemispherical bump.


cially appear to be similar processes, they areimplemented by exploiting different joining pro-cesses. Consequently, the resulting interconnectshave significantly different characteristics. Thekey features of each process and the character-istics of the resulting interconnects are summa-rized in Tables 5.8 and 5.9.

It should be pointed out that indium and in-dium alloys are sometimes used as solder inter-connects, and lead-tin alloys can be used as com-pression interconnects, so some care needs to betaken with terminology. Some companies are of-fering flip-chip processes based on conductivepolymeric media. Generally, these have charac-teristics similar to compression interconnects, butthey are not able to achieve interconnect diam-eters much below 100 μm (4 mil), and of coursethey lack the self-alignment feature, which is acharacteristic of molten metal.

The self-aligning feature of solder bump bond-ing requires some explanation. The surface en-ergy of a molten solder pillar between two con-tact pads will always attempt to minimize thesurface area of the liquid. This results in a re-storing force that regulates the relative locationsof each mating pair of pads. The force balancederives from the contact angle of the solder ontothe wettable pads, while the weight of the uppercomponent controls the planarity and magnitude

of the interconnect height. A schematic illustra-tion of the self-aligning process is given in Fig.5.10. Typical alignment accuracy is better than�2 μm (80 μin.) in-plane and �0.5 μm (20 μin.)vertical, and these figures can be improved bynearly an order of magnitude by careful geo-metrical design of the interconnects and atten-tion to process detail. Methods of achieving pre-cise alignment between flip-chip bonded partsare discussed in section 5.2.6 of this chapter.

KHz signal

Liquidmetalsupply

Nitrogengas valve

Temperaturecontroller

Chargedriver

CAD/CAMinput

Piezo crystalCharge electrode

High voltage

Orifice

Catcher

Circuit board

Fig. 5.8 Schematic layout of an apparatus designed to deposit individual solder balls at any desired location on a 6 in. diametersubstrate. CAD/CAM, computer-aided design/computer-aided manufacturing

Fig. 5.9 Flip-chip interconnect schemes. The prepared com-ponents are aligned and joined by either applying

pressure (solid-state compression bonding) or heat (soldering).


5.2.2 Characteristics ofFlip-Chip Technology

Flip-chip bonding can be defined as “amethodof providing simultaneous electrical connectionand physical attachment between two compo-nents.” It is attractive for low-cost, volumemanu-facturing because it is an inherently high-yielding, wafer-scale process. The key featuresand associated benefits of this technology are:

• Attachment and interconnection are per-formed simultaneously. This provides a ma-jor cost and yield advantage over other formsof interconnectionwhere the componentmustfirst be rigidly fixed to the substrate beforeproceeding with wire bonding or lead sol-dering. Fully automated flip-chip bondingmachines are commercially available that canflip, align, place, and reflow (if required) partsat throughput rates well in excess of 500/h.

• Self-alignment occurs between the compo-nents and substrates, through surface tensionforces in themolten solder (this characteristicapplies only to solder bonding, not compres-sion bonding). Thus, optical components canbe assembled in which a semiconductor laseror detector is reliably and accurately posi-tioned relative to other parts, such as opticalfibers.

• Flip-chip interconnects provide a low and re-peatable inductance, essentially a conse-quence of the short bond length. This char-

acteristic is particularly important for high-frequency circuits. A typical 25 μm (1 mil)diameter bond wire interconnection 0.5 mm(0.02 in.) long will have an inductance of0.25 nH, compared with approximately 0.05pH for a typical flip-chip solder bond.

• Parasitic capacitance is small, predictable,and reproducible.Again, this becomes amostimportant issue as the operating frequencyincreases. While the interconnection capaci-tance of a typical bond wire at 150 μm (6mil)pitch is 25 fF, it is only 2 fF for flip-chipbumps at 30 μm (1.2 mil) pitch.

• Interconnections can be accommodated overthe entire area of the component. Wire bond-ing and tape interconnection can only bemade around the periphery of chips. This re-stricts the total number of interconnectionsthat can be made and requires all signal pathson the chip to be suboptimally routed out tothe chip edge. Flip-chip bonding is, for simi-lar reasons, the ideal interconnection tech-nology for pixelated component architec-tures, one example of which is given in Fig.5.11.

• The interconnection area is smaller than thecomponent footprint. Close packing of digi-tal information processing chips is necessaryin order to minimize bus delay times. Allflip-chip connections are located underneaththe chip, whereas all other interconnectionschemes tie up substantial space all aroundthe chip periphery.

• This interconnection method results in lowprofile assembly, with flip-chip bumps just afewmicrons (<1mil) high (Fig. 5.12). There-fore, the chip packages and the whole elec-tronics system can be made correspondingly

Fig. 5.10 Schematic illustration of the self-aligning mecha-nism of flip-chip solder interconnects

Table 5.8 Process characteristics

Key parameterCompression bump

bondingSolder bump

bonding

Process Solid-state diffusion Solid-liquidalloying

Materials In, Au, Pb, Pb/Snalloys

Any solder

Temperature 20–200 °C (68–392°F)

Solder meltingpoint

Pressure 10–100 MPa (1.5–15� 103 psi)

0 MPa

Max height:pitch

1:5 3:1

Table 5.9 Technology characteristicsCompression bump bonding Solder bump bonding

Short (coin) interconnects Tall (pillar) interconnectsFilled joint gap Open joint gapFluxless Flux usually requiredLow residual stress Residual stressService temperature >

bonding temperatureService temperature <

bonding temperatureAlignment as placed Self-aligning


thinner. This characteristic is particularlyvaluable in portable electronics products.

• Flip-chip processing is compatible with deli-cate materials. The upper frequency limit forsilicon-base semiconductors is approxi-mately 5 GHz. Above this frequency, com-pound semiconductors such as gallium ars-enide (GaAs) or gallium nitride (GaN) mustbe used. Compared with silicon, these ma-terials are extremely fragile, so that alterna-tive interconnection techniques are preferredto mechanical wire and tape bonding. Thelatter can generate surface damage on thesemiconductors because of the forces in-volved in these processes.

• The interconnect is packageless, and the join-ingmedium is applied during component pro-cessing. Therefore, application of the mate-rials required for flip-chip bonding onlyrequires a few additional processing steps,and all components may be processed simul-taneously. Removal or minimization of pack-aging costs is always attractive.

5.2.3 Underfill

To improve the operational life of flip-chipassemblies, it is common to apply an underfilladhesive into the gap between the die and thesubstrate on which it is mounted. The objectivesof using an underfill are fourfold:

• Improve environmental protection of thesemiconductor

• Increase the mechanical strength of the as-sembly

• Enhance heat sinking of the die• Extend the fatigue life of the interconnects

The process generally involves placing an ac-curate volume of material along one or moreedges of the die and allowing it to be pulled intothe gap by capillary action. The adhesive is curedby the action of time, temperature, exposure toambient atmosphere, or a combination of these.Provided that the modulus and thermal expan-sion coefficient of the underfill material are ju-diciously chosen, then the fatigue life of solderbumps between a silicon chip and an FR4 PCBcan be increased by as much as two orders of

Fig. 5.11 (a) Semiconductor component of a pixelated imaging device. (b) Close-up view of a single solder ball. The imaging deviceutilizes a 16 by 128 array of 33 μm (1.3 mil) diameter solder bumps for the interconnects. Each pixel is connected to its

own amplifier and processing circuitry electronics. Courtesy of CERN

Fig. 5.12 An array of exceptionally small flip-chip solderbumps ready for bonding. Each is mounted on a

pedestal of polyamide for functional reasons. Solder bumps of thisdimension can only be realized using semiconductor processingtechniques. 5000×. Courtesy of BAE Systems


magnitude on thermal cycling between �65 °C(�85 °F) and 100 °C (212 °F) [Doi et al. 1996].

Flip-chip bonding is a rapidly evolving field,and a number of underfill materials are beingoffered that use different chemistries. Some ofthese are applied prior to assembly of the partsand actually flux the solder during the heatingcycle before providing conventional postbond-ing functions.

5.2.4 Inspection

One of the challenging features of flip-chipbonding is that the interconnects are not easy toinspect. Some process engineers consider this tobe a major benefit, because it restricts the scopefor quality inspectors to reject assemblies!Acom-parative review of flip-chip inspection methodsis given by Burdett, Lodge, and Pedder [1988].The problem is that only the peripheral bondscan be observed visually, and then often onlywith some considerable difficulty. The otherbonds are hidden from view beneath the com-ponents. Special microscopes have been devel-oped that are able to look a considerable distanceinto the gap between bonded parts (see section5.10.3 in this chapter). X-ray inspection is anobvious solution to this problem, stemming fromthe ability of x-rays to penetrate most materials.Fortuitously, the materials making up most com-ponents and substrates are lighter and less ab-sorbent to x-rays than the solder, which thereforeshows up clearly. The main defects that can beisolated using x-ray equipment include missingbumps, bridged bumps, excess bump material,and porosity in the interconnects. However, itcannot easily detect the most usual cause of yieldloss in flip-chip technology, which is intercon-nects that have failed to produce joints, that is,leave gaps. The reason for this is that small airgaps negligibly affect x-ray absorption, particu-larly in comparison with solder balls. Furtherinformation on x-ray inspection is given in sec-tion 5.10.2 in this chapter.

Infrared microscopy can be used to inspectcertain types of flip-chip assemblies. Infraredmi-croscopy is similar to traditional optical micros-copy but uses much longer wavelengths. Siliconand many III-V semiconductor materials aretransparent in this portion of the electromagneticspectrum. Infraredmicroscopy permits the tracksand the backside of the underbump pads to beexamined but not the detail of the bumps them-

selves, because solder bumps are opaque to in-frared radiation.

Arguably, the most appropriate technique forinspection of flip-chip bonded components isacoustic microscopy. An acoustic microscopegenerates and focuses an acoustic pressure waveinto the device under examination and uses thereflected echoes to construct an image of theobject at a specific depth. Themethod is sensitiveto boundaries between regions of different den-sity and is therefore ideally suited to discrimi-nating planar joints and transverse defects suchas unwetted joints, missing bumps, cracks, andvoids.Acoustic microscopy is a rapidly evolvingfield, and instrumentswith truly remarkable reso-lution are available commercially. Defects in 20μm (800 μin.) diameter solder bumps can be re-solved in real-time. Further information onacoustic microscopy can be found in section5.10.1 in this chapter.

5.2.5 ReworkSuccessful rework of a flip-chip mounted die

is difficult but not impossible. Local hot gasnozzles can be used to heat an individual chipuntil the interconnect metal melts, whereuponthe component can be lifted clear. The difficultylies in dressing the used bond pads on the sub-strate so they are left in an adequate state ofcleanliness and parallelism to accept anotherchip. For these reasons, rework tends not to bepracticed for fine-pitch interconnects or with diethat have been underfilled, although it is per-fectly practicable for larger bump bonds such asthose used with BGAs, which have coarser ge-ometries.

5.2.6 Self-Alignment ofFlip-Chip Structures

One of the remarkable features of solder flip-chip interconnects is their ability to self-align.Accuracies at the 100 to 200 nm (4 to 8 μin.)level, with respect to the underbump pattern ac-curacy, are possible. The factors that need to beconsidered to achieve these precisions are nowreviewed.

Self-alignment of flip-chip interconnects wasnoted almost from the inception of this solderingmethod, and subsequently, there was a lot ofinterest in its exploitation for the assembly ofoptical components. This interest has been re-viewed by Tan and Lee [1996]. The best align-ment from self-alignment is quoted as better than0.8 μm (30 μin.) for four 130 μm (5mil) diameter


bumps, while Hayashi [1992] and Tsunetsagu etal. [1997] suggest that better than 0.25 μm (10μin.) can be obtained with a larger number (20�)of smaller bonds (25 μm, or 1mil, diameter). Thebest claimed alignment accuracy for systemswithdeliberately misaligned solder bumps pulling thedevice against a mechanical stop was <0.3 μm(12 μin.) in the stop axis but greater in otherdirections [Lin et al. 1993].Attempts to use mul-tiple stops in different orthogonal directionswereunsuccessful. This is because the friction ofmov-ing the device past the first stop is greater thanthe available restoring force.

As stated previously, the driving force for self-alignment of solder bumps is surface tension,which seeks to minimize itself by adopting theshape of a perfect sphere, having the minimumsurface area for a given volume.At that point, thesurface tension is balanced by the internal pres-sure of the solder. Distorting the sphere, bysquashing, stretching, or shearing sideways, willbe resisted by a restoring force that is approxi-mately proportional to the additional surface areacreated.

When a solder bump is used to bond two com-ponents together, the situation is complicatedslightly by the presence of flat, wettable pads,plus the weight of the top chip. The solder willthen try to adopt a shape as close to a truncatedsphere as the volume of solder and the size andshape of the pads allow. The effect of the weightof the top chip is to slightly increase the internalpressure in the solder, and this will tend to reducethe restoring force. Software is available thatallows the restoring force to be calculated[Brakke 1990], from which the following geo-metric design guidelines emerge for achievinggood in-plane alignment:

• Reduce the compressive loading per bond,because some of the surface tension has to beused to support this weight. This can beachieved by having the smallest chip on topduring reflow, thinning the chip or reducingits size, and increasing the number of solderbonds.

• Use thinner bonds because for a fixed padsize and number, they have a higher restoringforce than ones with thicker solder, becausethe shear distortion per unit height is muchgreater. However, this implies that the soldervolume should be reduced, but this could havea detrimental effect on the joint reliability infatigue situations and on the ease of makinga joint, because the solder volume-to-surface-

area ratio becomes progressively more un-favorable with smaller bonds.

• Use smaller joints because they have slightlyhigher restoring force than a larger joint of thesameaspect ratio.This is because agivenmis-alignment distorts a small sphere much morethan a larger one. On the other hand, becausemany more ball bonds can be accommodatedin the same area than can larger bonds, there isless loading per joint and additional restoringforce from each additional bond.

• Design the pads so that at the neutral posi-tion, all bonds are equally offset. Because therestoring force diminishes rapidly as the axialalignment improves, there can be insufficientnet force available to move the upper com-ponent, particularly if the joint gap is filledwith liquid flux. By offsetting the bonds, theavailable force for a given displacement islarger and is never zero.

The other route of increasing the restoringforce and thus the accuracy of self-alignment isto increase the surface tension of the moltenmetal. In fact, this can be changed reasonablyeasily. Lead-tin eutectic alloy has a surface ten-sion of 0.483 N/m (2.8 � 10�3 lbf/in.) at 350 °C(660 °F) in inert atmosphere, compared to 0.55N/m (3.1 � 10�3 lbf/in.) for tin at the sametemperature and 0.44 N/m (2.51 � 10�3 lbf/in.)for lead. Replacing the lead with cadmium (0.59N/m, or 3.37 � 10�3 lbf/in.), zinc (0.78 N/m, or4.45 � 10�3 lbf/in.), or silver (0.92 N/m, or 5.25� 10�3 lbf/in.) should increase the surface ten-sion of the solder alloy and hence its restoringforce against misalignment. The other major fac-tor altering surface tension is the presence offlux. The surface tension of eutectic solder at 350°C (660 °F) drops from 0.48 N/m (2.74 � 10�3

lbf/in.) in inert atmospheres to 0.41 N/m (2.34 �10�3 lbf/in.) in the presence of nonactivated R-type flux. Additions of chlorine to this flux dropthe surface tension even further, so that using a1% Cl activated flux lowers the surface tensionto 0.31 N/m (1.77 � 10�3 lbf/in.). Liquid fluxesshould therefore be avoided and gaseous fluxesused instead (see Chapter 3, section 3.3 on flux-less soldering).

Z-axis, or height, control is a function of thevolume of the solder sphere, modified as neces-sary by the weight of the upper component.Achieving good z-axis control principally re-quires that accurate solder volumes are placedon pads whose surface areas are also well de-fined. These factors are primarily process con-


trol issues, although some design inputs willhelp. As before, many bonds mean that indi-vidual bond variations are less significant [Mc-Groarty et al. 1993]. For the vacuum depositionand wet back process, where solder is depositedover a greater area than the wettable metal pads,any variation in the thickness of solder depos-ited is exaggerated in the reflowed solder bumps.Similarly, plating a thick solder deposit over asmaller part of the wettable metal will dilutethese thickness errors. Solder deposition sys-tems that work on mass change rather thanthickness control tend to be superior in thiscontext. Bond design also helps, with morespherically shaped bonds being less sensitive todeposited thickness variations than more co-lumnar-shaped interconnects.

Predicting the extent of self-alignment thatcan be achieved from a flip-chip structure is pos-sible using the SURFACE EVOLVER computercode. Good agreement has certainly been dem-onstrated between the model and lead-tin solderbumps wetted onto copper pads [Josell et al.2002]. This suggests that such models can beused as a tool for predicting the alignment andstand-off height of specified interconnect geom-etries.

5.2.7 Surface Topography

One of the difficulties with flip-chip processesis that surface topography is difficult to accom-modate in a reliable manner. Quite simply, theout-of-flatness of the mating parts often repre-sents a significant proportion of the total heightof the interconnects.

One method of addressing this problem is toendow each of the mating bond pads with asolder land that is connected via a narrow path,as shown in Fig. 5.13. This provides a meanswhereby the excess solder on the first pads tomate is progressively drawn off, reducing thetotal height of the interconnect. Thereby, shorterinterconnects are able to form as the two partsare drawn together. The narrow land betweenthe pad and overspill area greatly slows the rateat which the solder height can decrease andtherefore ensures sufficient material is availableto make all of the interconnects. Obviously, thissystem demands additional surface area to ac-commodate these solder “drains,” and the addi-tional electrical capacitance that arises must bewithin design limits.

5.2.8 Step-SolderedFlip-Chip Interconnects

Solder flip-chip interconnects can be madewith essentially every known alloy. Step flip-chip soldering is therefore both possible and prac-ticed. Figure 5.14 is an example of a substratethat contains both wire-bonded and flip-chip-mounted die. By using the high-melting-point

Fig. 5.13 Flip-chip land designed to cope with surface to-pology. The solder is initially confined by surface

tension to the central circle but will slowly flow through theconstricting necks into the overspill areas. The ensuing reductionin solder volume, and hence bond height, enables adjacent in-terconnects of reduced height a chance to reach and wet theirmating pads.

Fig. 5.14 Radio frequency diemounted using high-tempera-ture flip-chip interconnects onto a substrate,which

is itself populated with lower-melting solder balls ready for directattach to a printed circuit board. Courtesy of Intarsia Corporation


Au-20Sn solder (melting point 280 °C, or 536°F) for this application and making the inter-connects as short as possible (30 μm, or 1.2 mil),it is possible to use a second flip-chip process toattach the substrate to a PCB, using, in this case,solder bumps of lead-tin eutectic solder (meltingpoint 183 ºC, or 361 °F).

5.3 Solderability Test Methods andCalibration Standards

The formation of strong joints is contingent onthe ability of the molten filler to wet the joint sur-faces over their entire surface area. Wetting ofmetal surfaces tends tobe inhibitedby surfaceox-idesandcontaminants.Visual inspection isnot re-liable for ascertaining the condition of surfaceswith regard to solderwetting.The surface appear-anceofmetallizations, inparticular, canbedecep-tive, because color and reflectivity are dependenton a variety of factors, such as grain size and type,whether equiaxed or columnar, surface texture,and any films present on the surface. Therefore,wettability has to be determined by a direct mea-surement involving exposure of a part represen-tative of a component to be joined to molten sol-der. These wettability measurement methods canalso be used to assess the effectiveness of a flux,the sensitivity of surfaces to exposure to variousatmospheres, and also shelf life, in general. Theevaluation of solderability shelf life tends to in-volveartificiallyaccelerated testing toobtaindatawithin a reasonable timescale.

Because soldering processes are dynamic innature—that is, they are sensitive to the timeover which the filler is molten—two aspects ofsolderability need to be considered, namely:

• The readinesswithwhich substrates and com-ponents are wetted by a filler

• The extent of spreading that is obtained bythe end of the process cycle, when the fillersolidifies

A number of evaluation procedures and testshave been developed to measure these charac-teristics and are described in the subsequent sec-tions. Reference is also made to the productionof calibration standards of solderability.

5.3.1 Assessment of WettingThe simplest and most popular test used for

assessing wettability is the dip-and-look (DNL)test. It is a standard test method for solderabilityused by the electronics industry for testing the

leads of integrated circuit devices with appro-priate finishes (e.g., ANSI/J-STD-002 and IPC/EIA J-STD-003A). In this test, the componentleads are immersed into a suitable rosin-type fluxfor 5 to 10 s and then immersed into moltensolder (generally, lead-tin eutectic) for 5 to 10 s.The component leads are removed from the sol-der, cleaned, and then visually inspected. Thepercentage of the lead surface that has remainedunwetted by the solder is then visually assessed.The standards specify the acceptance conditionthat “all terminations shall exhibit a continuoussolder coating free from defects for a minimumof 95% of the critical surface area of any indi-vidual termination.”

The chief disadvantage of the DNL test is thattheconditionsuseddifferconsiderablyfromthoseused in amanufacturing environment. In a typicalproduction line, an infraredor convection furnaceor vapor-phase reflow chamber is used to heat theassembly, with solder paste preapplied to thejoints.Thedynamicthermalenvironment is there-foreverydifferent and the relativevolumesof sol-der and flux verymuch smaller than that affordedby the test.

A variant of the DNL test, called the surface-mount solderability test (SMT), was devised tomoreclosely simulate theactual environment thatsurface-mount devices encounter during the sol-der reflow process. In this test, solder paste isscreened onto a thin, unmetallized ceramic plate(0.9 mm, or 35 mil, thick) using a solder stencil.Thepaste print used is the pattern of the footprintsof the leads to be assessed.The device to be testedis thenplacedonto the solderpasteprint.Next, theceramic substrate is passed througha reflowcycleand allowed to cool. The devices are removedfrom the ceramic substrate, and the leads are ex-amined for solder wetting. The advantage of thistest is that leadeddevicesaresubjected toasimilarthermalenvironment to thatexperiencedin theac-tual assembly. Furthermore, the leads receive thesame volume of solder paste as in the actual as-semblyoperation.TheSMTalsoneatlyavoids theoccurrence of solder bridging of fine pitch leads,towhich the conventionalDNLtest is prone.ThistestmethodhasbeenadoptedasanElectronics In-dustriesAssociation(EIA)interimstandard(EIA/IS-86).

The new SMT test method has been shown tobe especially effective for providing a reliableassessment of the solderability of palladium-plated integrated circuit (IC) leads. Palladium-plateddeviceshavebeenmarketedsince1989andnow account for a significant portion of the IC


market.The palladium is applied as a 0.076 μm (3μin.) plated finish over a nickel-plated lead. Itsfunction is simply to act as a protective barrier tooxidation for the nickel. During the solderingcycle, thepalladiumis required tocompletelydis-solve into the solder with the joint being formedwith the underlying nickel layer [Abbott et al.1991]. Trials using the SMT method have dem-onstrated that thepalladiumcoating is completelydissolved, as designed. By comparison, the DNLtest, with its short solder immersion time and ab-sence of preheat, hinders the dissolution of thepalladium from the lead surface.

Neither of the two DNL-type tests describedpreviously provide information about the dynam-ics of wetting, and neither is quantitative. A testthat has been devised to quantitatively measurewettingasa functionof timeis thewettingbalancesolderability test.Themethodhasbeenadoptedasa standard test for measuring the solderability ofelectronic component leads and substrates undera closely specified set of conditions that include aprescribed atmosphere and flux. Solderabilitytesters, as they are known colloquially, are avail-able from several manufacturers, one example ofwhich is shown inFig. 5.15.Atypicalwettingbal-ance comprises:

• A load cell and signal processing system thatfurnishes a measurement of load versus time

and provides automatic tareing of specimenweight

• A temperature-controlled solder bath• A bath lift or specimen fall mechanism with

speed and positional control• A computer to display the force/time curve

and derive key metrics from the data

The specimen under test is held in a holderthat is itself suspended from the load cell. Thebath containing the molten filler metal is raisedat a preselected speed to immerse the testpieceto a given depth (on some equipment, thetestpiece is lowered into the solder reservoir).The bath is held in this position for a set dwelltime and is then returned to its rest position.The values of these parameters are preselectedbefore commencing the test cycle and are de-fined by standards.

The resolved vertical forces acting on thespecimen are recorded as a function of timeover the whole test cycle. Figure 5.16 shows thetypical form of the trace that is recorded, to-gether with the corresponding position of thespecimen relative to the solder bath at eachstage.

The wetting balance provides a measurementof the vertical component of the force exerted onthe testpiece as it is lowered into a reservoir ofthe molten solder or braze, as a function of time.

Fig. 5.15 Commercially available wetting balance. Courtesy of Concoat


This force, FR, is theoretically equal to the sumof the vertical component of the surface tensionforce, F�, between the filler and the testpiece andthe buoyancy of the testpiece, FB. Figure 5.17shows an equilibrium situation appropriate topartial wetting. The resolved force in the verticaldirection, FR, is the parameter measured in thetest. The variation of this force as a function oftime provides information on the dynamics ofthe wetting process.

Typical wetting balance force/time traces aregiven in Fig. 5.18. The graphs show the effect ofdifferent cleaning methods on the wetting be-havior of lead–tin eutectic solder, heated to 235

°C (455 °F), on mild steel fluxed with a mildlyactivated rosin flux. In Fig. 5.18(a), the steelcoupon is in the as-received condition, and wet-ting by the solder is consequently very poor.Abrasive cleaning of the steel improves wetting,as shown in Fig. 5.18(b), but chemical cleaningis necessary in order to meet the quality-acceptance criteria, which are indicated by thebox in the lower left-hand corner of the graph(Fig. 5.18c). The acceptance criterion for elec-tronic components specified in national stan-dards is a wetting force that exceeds two-thirdsof the theoretical maximum force, achieved in atime representative of the soldering process to beused. For wave soldering, this is set at 2.1 s.

A wetting balance test can be performed rap-idly, and the results are quantitative, inasmuch asreproducible numerical data can be obtained fora well-defined set of sample and instrumentalparameters and operating procedures, as ex-plained in Barranger [1989] and Lea [1991]. Fur-thermore, the change in wetting as a function oftime can bemonitored. The surface tension of themolten filler can be calculated from data ob-tained on the wetting balance using nonwetted(� � 180°) substrates, such as polytetrafluoro-ethylene (PTFE) or ceramic coupons, using theequation:

FR � P�cos � � gV (see Fig. 5.17)

From this value and the measured wettingforce, the angle of contact between the moltenfiller and the testpiece can be calculated.

Attempts have been made to correlate wettingbalance datawith the results given by othermeth-ods used for assessing wetting [Thwaites 1981;Wooldridge 1988]. Moreover, adaptations havebeen made to the wetting balance for solderabil-ity testing on specific types of components and,in particular, for surface-mounted electronic de-vices and also in controlled atmospheres, includ-ing vacuum [Gunter and Jacobson 1990].

The only other test in common use that evalu-ates the dynamics of wetting is the measurementof contact angle in an enhanced version of thesessiledroptestdescribedbyMatienzoandSchaf-fer [1991]. In that test, a fixed volume of solder ismeltedonaflat couponof the substrateof interest,which is held at a fixed temperature.Aflux is gen-erallyadded if the test isperformedinair.Thecon-tact angle of the molten pool of solder is dynami-cally observed andmeasured.Theprinciple of themethod is illustrated in Fig. 5.19.

Fig. 5.16 Typical trace of the wetting force during a solder-ability test cycle, with the corresponding position

of the specimen relative to the solder bath

Fig. 5.17 Forces diagram for a solid plate partially immersedin a liquid. P, specimen periphery length; �, liquid

surface tension; �, contact angle, , liquid density; gravitationalconstant, g, 9.81 m2/s; V, immersed solder volume


There is not necessarily a close correlationbetween solderability test measurements and thewetting characteristics of actual joints. One ofthe reasons is that the configuration of the testdoes not strictly mirror that of an actual joint. Inparticular, the volume of themolten filler relativeto the volume of the components, including sur-face metallizations that react with it, may besignificantly different in the two cases. For ex-ample, the extremely fast dissolution rates ofgold coatings in solder, reported for samplesdipped into the reservoirs of solder in a wettingbalance, typically 1 μm/s (40 μin./s) or more, donot apply to joints soldered using foil preformsof much smaller volume. Thus, gold coatingsthat completely dissolve in the solder bath canpartly survive in actual joints after the same timeand temperature of the bonding operation.

Another difference is the nonequivalence ofthe thermal environment, as in the DNL test. Inmany materials systems, the filler metal will re-act differently with the testpiece as the tempera-ture changes. This is a critical aspect because, asexplained in Chapter 1, section 1.2.2, metallur-

gical reaction and not surface tension is usuallythe dominant driving force for wetting in liquid-solid metal systems. Because such a metallur-gical reaction occurs, to an extent that dependson the materials involved and their temperature,the wetted interface changes its composition andgeometry with time. This, in turn, means that themeasured wetting force tends to vary in thecourse of a test, and only in exceptional cases isthe wetting force equation strictly valid, depend-ing as it does on the classical model of wetting.Considerations such as this limit the value ofwetting balance tests for absolute measurements.However, as a means of obtaining comparativedata, and, in particular, for quality-control as-surance purposes (pass/fail determination), thismethod is most useful [Thwaites 1981].

5.3.2 Assessment of SpreadingOne of the simplest and most direct methods

that have been devised for assessing wetting byliquid metals on solid substrates involves mea-suring the area of spreading by the molten metal.

Fig. 5.18 Wetting behavior ofmild steel by lead-tin eutectic solder,measured on awetting balance at 235 °C (455 °F). (a) As-receivedcondition. Wetting occurs slowly and at an inconsistent rate. (b) Following mechanical abrasion of the coupon surfaces

immediately prior to testing. Wetting occurs more rapidly because of denudation of the oxide scale, but the pass condition, which isa wetting force of �4.5 mN achieved within 2.1 s of immersion, is not achieved in this case. (c) Following chemical cleaning. Thecomponent solderability now satisfies the acceptance criterion.


This type of basic spreading test measures wet-ting by a molten solder of a solid over the in-terface of contact and not the enhancement thatoccurs in narrow joints through the action ofcapillary forces (see Chapter 1, section 1.2.4).

The procedure that is widely adopted in area-of-spread measurements is to melt a filler-metalpelletofknownvolumeinaspecifiedatmosphere,with or without fluxes, and to allow it to spreadover the surface of a testpiece for a fixed period oftime under controlled conditions. These condi-tions, where possible, should be representative ofthe intendedapplication,because thespreadingofthe filler metal is usually sensitive to component-specific variables (e.g., surface finish) and to pro-cess variables (e.g., time at temperature).

The area of spreading of the filler metal ismeasured; this provides a relative index of wet-tability for comparative purposes. This index isthe spread ratio, which is defined as:

Spread ratio (Sr) �

Total plan area wetted by the molten metal

Original plan area of a metal pellet(of a specific geometry)

Because the geometry of the solidified fillermetal is seldom perfectly circular, image analy-sis should be used to provide an accurate mea-sure of the total wetted area. The technique ofimage analysis is capable of providing an as-sessment in real-time, so that with suitable in-strumentation, the spread area can be monitoredas a function of time on a single testpiece. Aperpendicular view also enables the contact angleto be similarly monitored. A sequence of spreadtests made to evaluate small changes to the com-position of a filler alloy is shown in Fig. 2.25.

An alternative means of determining relativespreading involves quantitatively defining aspread factor in terms of the volume, V, of fillermetal used and the maximum height, h, of thesolidified pool:

Spread factor (Sf) �(6V / )1/3� h

(6V / )1/3× 100

where (6V/)1/3 � D, the diameter of a spherecorresponding to volume V of the metal pelletbefore the spreading test. This term can be cal-culated from the density of the metal pellet andits mass.

The normal method used for measuring theheight (h) uses a micrometer. However, thisapproach introduces errors because of the ne-cessity to subtract the thickness of the substrate,which is likely to be much greater than h, fromthe measured height. A more accurate methodof measuring h is to determine the peak heightfrom metallographic sections or profilometertraces. Both of these methods also enable si-multaneous measurements to be made of thecontact angle between the resolidified pellet andthe substrate.

Assuming that the initial pellet of filler can beapproximated to a sphere and the resolidifiedfiller to a spherical cap of radius R on the surfaceof the substrate, the spread ratio (Sr) and spreadfactor (Sf) can be expressed in terms of the angleof contact (�), as follows:

Sr �4 cot2 �/2

(1 � 3 cot2 �/2)2/3

(Eq 5.1)

Sf � 1 �1

(1 � 3 cot2 �/2)1/3

(Eq 5.2)

Fig. 5.19 Principle of the sessile drop test used to assesswettability. (a) A controlled volume of filler metal

(solder) is melted onto the substrate under controlled conditions.(b) The contact angle is measured with a calibrated viewfinder.In an enhanced form of this test, the contact angle is recordeddynamically as a function of time.


The contact angle, �, can be written in termsof the radius, A, and height, h, of the resolidifiedpool of filler:

sin � �2

(A/h) � (h/A)(Eq 5.3)

The mathematical derivations of Eq 5.1 to 5.3are given in Appendix A1.2. Some numericalvalues relating spread ratio and spread factor tothe wetting angle (contact angle) are given inTable 5.10. Obviously, the situation representedby these expressions is an idealized one, and theassumptions made to derive them become lessrealistic with increasing solder spread. Never-theless, these relationships do provide a reason-ably close concordance with measured values.Some results of spreading tests are given in Fig.1.14, which provides spread ratio data for a rangeof solder alloys melted on thin chromium met-allizations covered with a flash of gold, as afunction of the process temperature.

The spread ratio becomes progressively moresensitive as the contact angle declines and wet-ting improves, whereas the spread factor variesalmost linearly with contact angle, from a valueof 1 at � � 0° to 0 at � � 180°, as shown in Fig.5.20.Spread ratio, therefore, providesabetterdif-ferentiation between small differences in mea-sured contact angle when the values of the latterare small. However, the converse is truewhen thecontact angle is greater than approximately 60°.

The results of spreading tests over a singlesurface must be treated with some caution whenattempting to relate them to the wetting and fill-ing of joints. Because a joint comprises a pair offacing solid surfaces, with which the filler metalcan react, the capillary forces can govern thespreading characteristics; there are hydrostaticforces to consider as well. The relevance of thespreading test is also questionablewhen the joints

are made using foil preforms. For this type ofconfiguration, it is not necessary for the fillermetal to spread significantly in order to fill thejoint, so that a low spread in the test does notnecessarily mean that a joint formed under simi-lar conditions will be poor. Indeed, a high degreeof spreading can be detrimental to joint filling,because the filler metal can flow out of the joint,resulting in voids and unwanted coverage of otherparts of the component. Nevertheless, it is often,but erroneously, assumed that low spreading inthe test necessarily implies weak interaction be-tween the filler and the substrate and thereforepoor bonding and weak joints.

An additional problem in attempting to ex-trapolate spreading test results to predict jointquality is that the filler metal reacts and spreadsover a single surface in a conventional spreadingtest to produce a somewhat different microstruc-ture from that of an actual joint. This is dem-onstrated in Fig. 5.21. Therefore, the wetting andspreading characteristics might be different inthe two cases. Large-area testpieces are neededfor measuring spread parameters in situationswhere there is good wetting.

5.3.3 SolderabilityCalibration Standards

There are occasions where it is necessary toconsider solderability tests from another stand-point, namely, where the solderability of thetestpiece is defined and some other property isunder investigation, such as occurs during the

Table 5.10 Calculated values of spread ratioand spread factor corresponding to selectedcontact angles. The values are derived from theexpressions given in the text (Eq 5.1 and 5.2).

Contact angle (�), degreesSpread ratio

(Sr)Spread factor

(Sf)

180 0 090 1.6 0.3740 3.7 0.6510 9.7 0.860 Infinite 1.0 Fig. 5.20 Relationships among spread ratio, spread factor,

and contact angle


development of a new flux or when comparingsolders from different manufacturers. In otherwords, there is a need for solderability calibra-tion standards. Perfectly wettable and nonwet-table testpieces are readily obtainable; gold andanodized aluminum are examples of the formerand latter, respectively. The problem ariseswherethere is a need for testpieces having intermediateand known degrees of solderability.

The published literature cites at least twometh-ods of preparing reference materials [NormeFrancaise 1987; Lea 1990]. Neither method isentirely satisfactory. Both depend on chemicaltreatments to modify a clean copper surface andthereby degrade the solderability. The modifiedsurfaces are not chemically stable, so that thetestpiece must be freshly prepared before eachuse. Hence, there is considerable batch-to-batchvariability that becomes amplified when differ-ent chemists at different sites are involved.

In recognition of this deficiency, several teamsof researchers, funded by the United KingdomDepartment of Trade and Industry, were giventhe goal of developing a solderability standardreference material. The solution finally devisedwas a two-layer electrodeposit consisting of 10μm (390 μin.) pure nickel, applied from a low-stress sulfamate solution, overlaid with 5 μm(197 μin.) of pure, soft gold [Hunt and Wallis1995]. The substrate can be any inert material(from a soldering point of view), and nickel wasselected because it does not contaminate thenickel sulfamate plating bath. Gold applied di-rectly to nickel strip was not successful, becausethe composition of readily available nickel stripis not sufficiently tightly controlled. Problemswere encountered with impurities in the basemetal, especially zinc, diffusing through to theouter gold plating. Plating baths are formulatedto very precise specifications from high-puritychemicals, so the quality of the metals obtainedfrom this source tends to be highly repeatable.

The as-prepared testpieces have “perfect” sol-derability, because the wettable surface is puregold, with the plated nickel interlayer providingan effective barrier to diffusion of impurities fromthe nickel substrate to the gold. Because the goldcoating is thick and the diffusion of nickelthrough gold at room temperature is negligible,the reference standards have a long shelf life.Trials indicated no change in solderability whenstored for a year in an open laboratory environ-ment. In order to degrade solderability of thetestpieces in a controlled manner, the plated ma-terial was heat treated in air for 2 h. The long heattreatment time was chosen to ensure that errorsarising from the heating and cooling stages,which tend to be difficult to control, are minimal.It is desirable to load plated sheets directly intoa preheated furnace and, at the end of the allottedtime, remove them to ambient to achieve thefastest possible temperature change without in-troducing surface contamination.

The heat treatment temperature needs to bequite tightly controlled, because the tightlygrouped process temperatures of 295, 310, and328 °C (563, 590, and 622 °F) yield three in-termediate degrees of solderability. Because ofthe small temperature interval between thesethree temperatures, some care needs to be takento ensure that the furnace temperature is cali-brated and the referencematerial experiences thespecified set-point temperature.

During the heat treatment, nickel will diffusethrough the gold and, on reaching the free sur-

Fig. 5.21 (a) Solder spread sample on a gold-plated sub-strate. The solder is Pb-60Sn, and the substrate is

copper plated with 5 μm (200 μin.) of nickel and then 5 μm (200μin.) of gold. At least three distinct microstructural bands arevisible. (b) Micrograph of a joint made using the same substrateand solder described in (a). The joint has a regular microstructure,and all of the gold coating has dissolved in the solder.


face, will oxidize. Because nickel oxide is onlyremoved by the most aggressive of fluxes, thereference material has variable solderability, de-pending on the proportion of nickel oxide in thegold. Figure 5.22 shows the concentration of oxy-gen (atomic percent) in the gold at a depth of 3nm (0.1 μin.). Figure 5.23 represents the wettingforce measured for each solderability referencestandard using three commercially availablefluxes having different activities. The resultswerefound to be consistent, with scatter bars of �2mN (�0.007 ozf) able to fully accommodate allof the experimental results obtained from an in-terlaboratory comparison involving five sourcesof solderability reference material, six test labo-ratories, and three different makes of solderabil-ity test equipment. The fluxes used in all the trialswere of the same type, obtained from one manu-facturer, but were from different batches and

dates of production. This reference material withthe plated layers therefore fulfills the require-ment of being a readily reproducible calibrationstandard for assessment of solderability.

5.4 Amalgams as Solders

An amalgam is a mechanically alloyed mix-ture of liquid metal and solid powder. They areusually designed such that metallurgical reactionat room temperature or when slightly warmedresults in isothermal solidification, and the mix-ture sets solid. Amalgams therefore merit con-sideration as solders. Themolten constituents area low-melting-point metal—either mercury, gal-lium, or indium—or amixture of these, while thesolid powder contains metals having a consid-erably higher melting point. A list of liquid met-als and solid powders that have been explored asamalgams is given in Table 5.11.

Because amalgams initially exist as pasty flu-ids, they offer a number of unique advantagesover conventional solders. The bonding processoccurs at low temperature, without flux, so thatit greatly reduces the equipment complexity andconfers much process flexibility. In particular,amalgams can accommodate large out-of-planeengineering tolerances, because they are semi-solids, but yet can be dispensed precisely at thebond location. Furthermore, because amalgamstake a finite time to set, often many tens of min-utes, and are electrically conductive while mol-ten, a product can be tested for electrical func-tionality while the amalgam is still liquid, withtime to replace any faulty components before thebond is cured. This property is also likely to beof benefit in the assembly of optoelectronic de-vices, where it is frequently necessary to under-

Fig. 5.22 Oxygen concentration at 3 nm (0.1 μin.) depth inthe gold surface of solderability reference standard

material as a function of the heat treatment condition. The sub-strate material is nickel sheet, electroplated with 10 μm (390 μin.)nickel, overlaid with 5μm (200 μin.) gold.

Fig. 5.23 Wetting force at 2 s, determined using a wettingbalance, for three commercial fluxes, as a function

of the heat treatment condition used to produce the solderabilityreference standard

Table 5.11 Solid and liquid metals that havebeen evaluated as amalgams

Melting pointSolid powders Liquid metals °C °F

Antimony Mercury �39 �38Chromium Gallium 29 84Cobalt Indium 159 318Copper Gallium-indium-tin 5 41Germanium Gallium-indium 15 59Gold Gallium-tin 16 61IronNickelManganesePalladiumPlatinumSilverVanadium


take active alignment before the parts are fixedin place, yet one does not want the further move-ment that can occur when a higher-melting-pointsolder solidifies and cools. Amalgams are alsosuited for integration in step-assembly pro-cesses, because the melting point of a solidifiedamalgam is typically 200 to 600 °C (390 to 1110°F). Therefore, multiple joints can be made se-quentially using the same process, which can becarried out below the upper service temperatureof the product. The liquid component of commonamalgam systems also possesses the character-istic of being able to wet directly many non-metals, including ceramics, glass, and, of course,tooth enamel.

There have been some attempts over the yearsto devise amalgam systems for electronic as-sembly. These have met with only partial suc-cess. It is not clear from this work whether thereis a fundamental limitation that will preclude thewidespread availability of amalgams with favor-able functional properties and process charac-teristics or whether it is simply a lack of researcheffort. Certainly, dental amalgams have benefitedfrom very detailed study and, as a result, arehighly effective for the function for which theyare designed.

5.4.1 Amalgams Based on Mercury

Dental amalgams have been in existence sincebefore 659 A.D.; the alloy first appears in a Chi-nese book with a title translated into Latin asMateria Medica, written by SuKung in the fourthyear of theTang emperor Hsien Ch’ing (659 A.D.)[Chu 1958], but the formulation currently usedin restorative orthodontology was not deviseduntil the early 20th century.

The basic method for making amalgam in-volves a diffusion reaction between silver pow-der and mercury in a ratio such that all of themercury is eventually consumed in the formationof an intermetallic compound. The reaction maybe simply written as:

2Ag � 3Hg � Ag2Hg3

(melting point � 127 °C, or 261 °F)

The intermetallic compound is often referredto as the gamma phase in the solidified alloy.

Modern amalgams contain approximately 25%Sn. The presence of tin accelerates the consump-

tion of the liquid mercury and also results in analloy that has a small, controlled expansion onsetting. This is necessary to effect a good seal tothe walls of the tooth cavity but obviously mustnot be so great as to cause pain or crack the tooth.Copper is added to tie up the tin as Cu6Sn5 in-termetallic compound, because this phase isproven to have superior mechanical propertiesand resistance to corrosion by oral fluids than theHgSn7 phase that would otherwise prevail.Amal-gam metallurgy is a relatively complex subject,and a whole family of alloys is available withvarying proportions of minor elements to tailorthe properties of the amalgam for its position inthe mouth and the type of cavity it is being usedto fill.

The mechanical properties of mercury-baseamalgams cannot be expressed in the same lan-guage as used in mainstream metallurgical tech-nology without distinct risk of misinterpretation.For example, the tensile and compressive be-havior are totally different, and the creep behav-ior does not correspond with conventional be-havior. The most accurate description ofmechanical properties is obtained when usingrheological concepts, which are applicable to vis-coelastic materials. The deformation process in-volves dislocation climb and grain-boundarysliding, and, when overstressed, a dental amal-gam fails by intergranular brittle fracture [Wa-terstrat 1990]. A stress-strain curve for dentalamalgams at low strains is given in Fig. 5.24.

The mercury in dental amalgams is bound upin theAg2Hg3 intermetallic compound, but thereis concern that some mercury might be slowlyreleased through extended contact with saliva.Mercury is classed as a hazardous substance.Efforts have been underway to develop resin ma-terials for dentistry that are color-matched toteeth, but despite recent advances in the tech-nology, these remain inferior to conventional

Fig. 5.24 Stress-strain curve for a dental amalgam (mercury-silver base) at low strains. Adapted from Dickson,

Oglesby, and Davenport [1968]


amalgams, particularly in their durability. Theymostly contain beryllium, which is also hazard-ous in some forms, including the oxide.

The requirements of an amalgam intended foruse as a solder are very different to those used fordentistry. A dental amalgam works by filling aspecially shaped cavity in the tooth that is cre-ated by the dentist. On curing, the amalgamswells slightly to lock the cohesivemass of metalin place. By contrast, a solder must wet and jointwo parallel surfaces. Also, a dental amalgammust set in minutes and attain useable strengthwithin 2 h. The amalgam used as solder mustafford sufficient persistence of fluidity to givesensible bench life, yet only slightly elevatedtemperature is desired to accelerate the harden-ing process. Nevertheless, the detailed under-standing of mercury amalgams that now existsmay facilitate the development of amalgams formanufacturing based on gallium and indium.

5.4.2 Amalgams Based on GalliumGalliummelts at 29 °C (84 °F) and is therefore

a potential base for formulating very low-process-temperature amalgams without the toxichazard associated with mercury. Gallium is nota particularly expensive metal, having a com-parable price by weight to solder pastes. Mostgallium amalgams exhibit significant volumechange during curing, but certain compositionshave volume change that is similar to mercuryamalgams. Indeed, because of toxicity concernswith mercury, gallium-base amalgams have beendeveloped for dentistry and are commerciallyavailable on the Japanese andAustralianmarkets[Reusch, Geis-Gerstorfer, and Zeigler 1988].Over the typical service temperature range ofelectronics equipment, the electrical and thermalconductivities of gallium amalgams are similar

to other solder alloys, as can be seen from the listof properties cited for a range of gallium-nickel-copper amalgams in Table 5.12 [Mackay 1993].

Gallium forms amalgams with several metalpowders, of which copper and nickel are themost intensively studied. Unfortunately, the pub-lished literature does not elucidate on the im-portant parameters of powder condition, particlesize and shape, nor often the amalgamationmethod. However, from the available micro-graphs, it is possible to deduce that the nickel andcopper powders usedwith gallium amalgamwereprobably spherical and roughly 25 μm (1 mil) indiameter. Gallium is also known to form amal-gams with silver, and it is regrettable that thispotentially interesting combination has not beenevaluated further.

To form a gallium amalgam, a mechanicalmixture of the constituent metals is warmed untilthe temperature reaches 35 °C (95 °F) and thenthoroughly blended.A convenient method of do-ing this is in a commercial dental amalgamator,which contains a pestle to assist in obtaining anintimate mixture of powder and liquid in a shortperiod of time. The amalgam will then directlywetmost commonmetallizations used in the elec-tronics industry and also many nonmetals, in-cluding alumina. Due to the high proportion ofsolids in the amalgam, some mechanical assis-tance in the form of scrubbing or wiping is nec-essary to spread the mixture. Although galliumalloys will cure at room temperature, to achieveacceptable process times, the use of elevated tem-perature is beneficial. The graphs in Fig. 5.25show the cure behavior of Ga-5Ni-30Cu amal-gams at different temperatures (for comparison,solidified lead-tin eutectic has an equivalent hard-ness, on the same Shore Durometer scale, ofapproximately 80).

When fully cured, high-strength and well-filled joints with good fatigue resistance can be

Table 5.12 Properties of some gallium-copper-nickel amalgamsPowder content(a), wt%

Property 5Ni-20Cu 5Ni-30Cu 5Ni-40Cu 10Ni-20Cu 10Ni-40Cu

Set-up time at 35 °C (95 °F), min 530 40 15 250 8Curing time at 100 °C (212 °F), min 90 10 2 70 1Expansion on curing, % 31 12 21 15 8Shear strength, MPa (psi) 23.5 (3410) 48.5 (7040) . . . 29.9 (4340) 39.2 (5690)Elongation, % 8.5 7.4 . . . 9.3 12.4Electrical resistivity, μohm · cm 18.2 11.8 10.8 . . . 12.1Thermal expansivity, 10�6/K 10.4 4.6 5.3 7.3 �0.2

(a) Balance gallium


obtained, which gives an indication of the po-tential of this soldering method. Gallium amal-gams have been demonstrated as filler metals ina number of electronics-oriented applications,including semiconductor die attach, flip-chipinterconnection, and via-filling in PCBs. In thelatter example, the amalgam was applied byscreen printing [Bhattachatya and Baldwin2000]. However, it is generally the case that theresults to date show considerable scatter in termsof joint quality, and further research is requiredto develop processes satisfactory for industrialuse [Baldwin, Deshmukh, and Hau 1996].

5.4.3 Amalgams Based on Indium

Indium is another liquid metal that can beconsidered as a base for amalgam systems. Par-ticularly after alloying with tin and bismuth, themelting point of the liquid can be depressed tobelow 100 °C (212 °F). Because indium alloysmelt at temperatures that are substantially abovethe melting points of mercury and gallium, it ispotentially possible to premix indium amalgamsusing liquid indium. Then, the curing reactioncan be suppressed temporarily by quench cool-ing the liquid-powder mixture until it is needed.

The authors are only aware of attempts to formindium amalgams using silver powder. As withthe gallium amalgams, the very restricted trialdid not yield a useable mixture, which was at-tributed to the availability of only relativelycoarse crystalline silver powder. Amalgams ap-pear to be more successful if the powder used isfine and spherical. Nevertheless, the mechanicalproperties of the silver-indium intermetallic com-pounds are known to be favorable in indium

joints to silver-metallized surfaces, and there-fore, further endeavor in this area may prove tobe fruitful.

5.5 Strengthening of Solders

Compared with most metals and alloys ineveryday use, soft solders have intrinsically poormechanical properties, that is, low strength andinferior resistance to creep and fatigue. For thecommonlyused lead-tin solders, application tem-peratures can easily be up to 90% of the meltingpoint, expressed in Kelvin. This means that themicrostructure of solidified lead-tin joints tendsto be unstable under typical service conditions,and frequently, grain growth will occur, to thedetriment of the creep and fatigue resistance ofjoints.Many solders are also susceptible to stress-induced microstructural changes. The superpo-sition of these microstructural changes on acyclic strain regime results in concentrations ofshear bands developing in the alloy microstruc-ture. Such features are prone to initiating fa-tigue cracks. Hence, solders differ from com-mon metals in that fatigue failure is initiatedinternally, whereas in engineering structures,fatigue cracks almost always start at an exteriorsurface. The poor mechanical properties be-come most apparent when the joint gap is wide(>25 μm), so that a proportion of the filler metalthrough the joint width is free from the con-straint of the parent materials.

It is precisely because the failure mode of sol-dered joints in electronics systems is so complexand depends on many interrelated variables thatthe test methods used to assess their integrity andreliability are specific to that industry, with thetesting usually being carried out on fabricatedassemblies. Further details are given by Coombs[1988].

So long as the primary function of solders hasbeen to provide electrical contact, their mechani-cal weakness could largely be tolerated, althoughcreep and fatigue failure arising from thermalexpansion mismatch of abutting componentshave caused occasional reliability problems.Thiswas the situation in electronics prior to the adop-tion of surface-mount technology. This devel-opment and, in particular, the reliance on supportstructures of solder, coupled with the continuingtrend toward miniaturization, have focused at-tention on the mechanical weakness of solderedjoints.

Fig. 5.25 Curing curves for Ga-5Ni-30Cu amalgams at arange of temperatures


The key to boosting strength and, in particular,conferring fatigue resistance to joints made withsolder is the development of a thermally stableand fine-grained microstructure. Three strategieshave been pursued in an attempt to address thisneed:

• Grain refinement• Dispersion strengthening• Composite solders

All three approaches actually lead to a similarresult, namely, refinement of the solder micro-structure. They are discussed here mostly in re-lation to the lead-tin solder system, for whichmuch of the development work has been carriedout because of their ubiquitous use and poormechanical properties, although the same ap-proaches could be adapted to other solder sys-tems.

5.5.1 Grain RefinementGrain refinement is conventionally achieved

by small additions (0.001 to 0.5%) of selectedelements, such as lithium, beryllium, indium, andgallium, to lead-tin solder [Klein Wassink 1989;Wade 1999; Tribula and Morris 1990]. This ap-proach operates by creating a fine dispersion ofeither oxide or nitride particles or stable inter-metallic phases when the solder is molten. Thesethen act as sites from which the solid phases cangrow on solidification (heterogeneous nucle-ation). The presence of a large number of suchfine particles when the alloy ismolten is reflectedin a fine alloy microstructure. The grain size thatcan be achieved is typically 150 μm (6 mils),compared with 300 μm (12 mils) for pure lead-tin solder solidified at the same rate. Althoughgrain refinement is beneficial to the mechanicalproperties ofmetals, improving both strength andductility, the fine microstructure is not thermallystable and will gradually coarsen by solid-statediffusion when thermal energy is provided byexposure to elevated temperature (typically,above 75 °C, or 165 °F).As can be seen from theexample given in Fig. 5.26, the creep-rupturetime of a solder doped in this manner is an orderof magnitude longer than that of regular lead-tinsolder. Similar improvements to the creep prop-erties have been achieved for other binary tin-base solders by grain refinement [McCabe andFine 2000]. Similarly, examples of ternary alloysolders, which have been shown to benefit fromgrain refinement accomplished by minor addi-tions, include 5In-87Sn-8Zn (melting point, 188

°C, or 370 °F) and 3.5Ag-95.5Sn-1Zn (meltingpoint, 217 °C, or 423 °F) containing additions ofsilver or copper, respectively, as the source of theheterogeneous phase at concentrations in the re-gion of 0.1 to 0.5% [McCormack and Jin 1994].

This method of improving resistance to failureby creep or fatigue of soldered joints has thebenefit that it is simple to achieve, because itmerely involves controlled doping of the solderduring manufacture.

5.5.2 Oxide-Dispersion-StrengthenedSolders

Dispersion strengthening involves the incor-poration of fine and intrinsically insoluble par-ticles into the solder, usually by mechanicalmeans. Examples are TiO2, SiO2, and Al2O3, allof which are cheap and inert compounds (not-withstanding the normal hazards associated withhandling dust) and readily available as fine pow-der with precisely controlled size distributions.The particle size is typically 0.1 μm (4 μin.)diameter but frequently much less. The loadingof refractory compound in the solder is typically3% by volume. These dispersoids provide somegrain refinement on solidification of the solderby heterogeneous nucleation, but their principalrole in improving themechanical properties is byother means. First, they inhibit coarsening of themicrostructure; because they are insoluble in thesolder, it is thermodynamically favorable for theparticles to reside at regions in the alloy matrixwhere there are natural departures from a regularatomic lattice, such as the boundaries betweenthe tin- and lead-rich phases in lead-tin solder.Secondly, provided the particles are sufficientlyfine, they will, for similar reasons, impede themovement of dislocations through the solderma-

Fig. 5.26 Creep curve for lead-tin eutectic solder and a dis-persion-hardened equivalent alloy containing 0.5

wt%Ag, 0.5wt%Sb, 0.1wt%Cu, and 0.003wt%Gaat a constantstress of 10 MPa (1450 psi) and a test temperature of 60 °C (140°F)


trix. Dislocations are defects in the atomic latticethat can be induced to move through it by theapplication of mechanical stress; the movementof many such dislocations results in plastic flowof the material. Hence, by restricting the move-ment of dislocations, the mechanical propertiesof the solder are enhanced. This type of disper-sion is less sensitive to thermal degradation, andtherefore, the strengthening effect is more resil-ient to elevated temperatures than is simple grainrefinement frommetallic precipitates. Table 5.13demonstrates the boost to the creep resistance oflead-tin eutectic solder that can be obtained evenat elevated temperature through dispersionstrengthening.

To be effective obstacles to dislocation move-ment, the particles must be within a certain sizerange, be stable in size and interparticle spacing,and have a higher flow resistance than thematrix.Dispersion strengthening is quite well under-stood, from a theoretical perspective, and theproperties of suitable dispersoids can be calcu-lated from first principles. It transpires that fineoxide particles are a good choice for inclusion insolders.

Because the dispersoids are insoluble in thesolder, dispersion strengthening remains effec-tive even when the alloy is heated almost to itssolidus temperature. No change in microstruc-ture is reported even after heating for 48 h at 120°C (248 °F). Furthermore, because the particlesare unreactive toward the constituents of the sol-der, the dispersion strengthening occurs equallyin both the tin-rich and lead-rich phases. Thisresults in the modified solder exhibiting im-proved ductility, tensile strength, and resistanceto creep [Mavoori and Jin 1998]. With a suffi-cient dispersion of nanosized TiO2 particles, ithas been possible to boost the creep resistance ofsoft solder at room temperature to a level com-parable to that of theAu-20Sn alloy and increaseits tensile strength fourfold, albeit with a corre-sponding reduction in ductility [Mavoori and Jin2000]. One useful side benefit from stabilization

of the alloy microstructure is that it helps sig-nificantly with simplifying reliability predictionmodeling.

The principal impediment to the widespreadadoption of dispersion-strengthened solders isthe difficulty of manufacture. In order to be ef-fective, the particles need to be present as a fineand uniform dispersion in the solidified joint.The problem is that particles tend to agglomer-ate, trap porosity, and adversely affect the vis-cosity—and hence, spreading ability—of the sol-der. While solder alloys can be producedsatisfactorily on a small scale in the laboratory,the specialized equipment and processes re-quired for their preparation and use currentlypreclude them from being drop-in replacementsfor conventional solder in most applications.

5.5.3 Composite Solders

Composite solders are not fundamentally dif-ferent from the other two types of strengthenedsolders described previously, in that they are con-ventional solder alloys with improved mechani-cal properties that arise from the presence ofsmall, hard particles. Thus, the strengtheningmechanisms are as mentioned previously,namely, a combination of grain refinement, grain-boundary pinning, and impediment of the move-ment of dislocations. What differentiates thesematerials is the choice of particle to provide thereinforcement, being predominantly copper-tin(Cu3Sn, Cu6Sn), silver-tin (Ag3Sn), nickel-tin(Ni3Sn4), or copper-nickel-tin (Cu9NiSn3) inter-metallic compounds [Guo and Subramanian2002; NEPCON West 1992; Betrabet, McGee,andMcKinlay 1991;Marshall et al. 1991]. Theseadditions satisfy the following conditions [Guoand Subramanian 2002]:

• The molten solder matrix wets the additivesand bonds strongly to it.

• The reinforcements are partially soluble inthemolten solder at normal temperatures usedfor reflow, so that the reinforcements are es-sentially stable during reflow and aging.

• The density of the reinforcements and soldermatrix are sufficiently similar so that gravi-tational separation does not occur, and themixture remains homogeneous when the sol-der is molten.

• These reinforcing phases do not appreciablycoarsen during normal service.

Table 5.13 Compressive creep rates of normaland dispersoid-strengthened solders at 100 °C(212 °F) and an applied stress of 1.7 MPa (247psi)Solder Steady-state creep rate at 100 °C, s�1

Pb-63Sn 1.26 � 10�6

Pb-63Sn � 3 vol% Al2O3 4.25 � 10�8

Pb-63Sn � 3 vol% TiO2 2.61 � 10�8

Au-20Sn 8.40 � 10�8


• They tend to retard the growth of the inter-metallic layers at the solder/substrate inter-face. Although such layers are essential tosuccessful bonding, their growth in thicknessresults in deterioration in the mechanicalproperties of joints.

Composite solders come in two flavors. Themore common contains fine particles (<5 μm, or0.2 mil, in diameter) at low volume fractions (0.5to 20 vol%) to produce grain refinement. How-ever, it has been found that the reinforcing par-ticles are most effective in stabilizing the mi-crostructure of the solder when they are of theorder of 1 μm (40 μin.) in size. Figure 5.27 andTables 5.14 to 5.17 illustrate some of the re-ported benefits to mechanical properties. In gen-eral, metallic additions are less effective in boost-ing creep resistance of soft solders than lowerpercentages of dispersed oxides (compare Tables5.14 and 5.13). Metal additions have a beneficialeffect on creep resistance to high cycle fatiguebut are detrimental for low cycle fatigue (Tables5.15, 5.16). Copper particulate reinforcement issuperior to silver from the point of view of creepresistance (Table 5.14). However, in situ Cu6Sn5-reinforced solders are superior to those strength-ened with copper or silver. Solders containingdistributed Ni3Sn4 particles are better still, in re-spect of high cycle fatigue and tensile properties(Tables 5.15, 5.17).

A second and less common type of compositesolder is filled with larger particles (typically 5to 25 μm, or 0.2 to 1 mil, in diameter) at volumefractions of 10 to 40%. These are akin to metal-matrix composite materials and benefit from en-hanced tensile strength, resistance to creep, andimprovements to other mechanical properties atthe expense of ductility and ease of application.The combination of higher volume fraction andlarger particle size is effective in strengthening

the solder matrix by restraining the yield behav-ior of the soft matrix. The material combinationsthat have been evaluated are largely the same asthose described previously for the first type ofcomposite.

To date, both types of composite solders havebeen difficult to use in joints without a substan-tial concentration of micropores developing,which offset much of the gains in properties out-lined previously. Indeed, while there is a rea-sonable body of literature largely extolling themechanical properties of composite solders, thedata almost exclusively pertain to specially pre-pared bulk samples of the filler metal and not tothe properties of joints of typical geometry andaspect ratio.

There are three common methods by whichcomposite solders are prepared: powder blend-ing, mechanical mixing, and rapid solidification.In powder mixing, powder of the reinforcementof interest is simply stirred into a vat of moltensolder prior to use.The difficultywith thismethodis that it is difficult to obtain a uniform disper-sion, and small particles tend to dissolve andreprecipitate on the surface of larger particles,resulting in a relatively coarse dispersion. Me-chanical mixing entails placing solder powderand powder of the reinforcing constituent in aball mill and working the mixture until the de-sired distribution is obtained. Obviously, someof the benefits of preparation in this manner willbe lost when the solder ismelted to effect joining.The third method involves gas atomization ofessentially off-eutectic composition alloys. Thereinforcement arises from the fine dispersion ofprimary and secondary phases of intermetallicthat precipitate out on cooling. Again, somecoarsening is inevitablewhen the solder ismeltedin the joining operation. Mechanical mixing andgas atomization yield powder that then needs tobe added to a binder and flux to form paste.

Composite solders generally have inferiorwet-ting and spreading characteristics, compared tonormal solders, and the visual appearance of thejoints is impaired. A frequent and very apt de-scription is that the solder is “gritty.” Wettingbalance tests confirm that the wetting time ofcomposite solders is affected detrimentally andthe spreading is decreased by as much as 25%[Steen and Becker 1986]. However, the wettingangles of molten composite solders, in sessiledrop tests using flux, are not appreciably higherthan the unmodified alloys [Subramanian, Bieler,and Lucas 1999], as shown in Table 5.14. Pro-ponents of composite solders argue that the poor

Fig. 5.27 Yield strength of composite solders at room tem-perature plotted as a function of volume fraction of

the added intermetallic powder [Yost, Hosking, and Frear 1993]


spreading is not a great impediment for elec-tronics assembly applications where solder(paste) is usually preplaced and minimal furtherspreading is required to effect joining. The stifferrheology of the semimolten composites results inwider joints than for conventional solders, whichmay be advantageous where there is a desire tobridge gaps, although wider solder joints are in-trinsically weaker, counterbalancing potentialbenefits to mechanical properties. Wide jointsare a decided disadvantage for most electronicand optical assembly because of the relativelypoor electrical and thermal conductivity of sol-

ders, compared with metal conductors, notably,copper, silver, and gold.

The improvements in materials properties re-sulting from the inclusion of dispersoids in thefiller metal do not greatly expand the possibili-ties for using lead-tin solders in load-bearingapplications. This is because the reinforcing pro-cesses are only effective at low strain rates, ascan be seen in Fig. 5.28. At higher strain rateslikely to be experienced in load-bearing struc-tures (10�2 s�1 and above), properties such astensile strength are no different between com-posite and conventional solder [Mavoori and Jin1998, 2000]. Furthermore, although the disper-soids improve the resistance to creep, high cyclefatigue life, and stress-rupture life of solder, it ismostly at the expense of ductility, low cycle fa-tigue life, and fracture toughness, which are prop-erties likely to be of greater importance in struc-tural applications.

An interesting variation of composite soldershas been proposed that utilizes insoluble par-ticles in the form of iron powder. Iron is wetted,but not consumed, by alloying at a particularlyfast rate by tin-base solders and therefore pro-vides a similar degree of reinforcement as otherdispersants. However, because iron is a soft mag-netic material, the natural shape of the solder,when molten, can be altered by application of anexternal magnetic field, and the modified profilewill be frozen when the solder solidifies. Fieldsin the region of 0.05 to 0.5 T (500 to 5000 G) willproduce significant height change of molten sol-der spheres. It is suggested that this effect mightbe exploited to help remove the effects of joint

Table 5.14 Wetting and creep properties of composite solders reinforced with metallic additions

Average wetting angle on copper Creep rate at a steady stress of 17 MPa, s�1

Solder (fluxed) substrates, degrees 25 °C 65 °C 105 °C

Sn-3.5Ag 18.8 2.62 � 10�5 1.50 � 10�4 1.9 � 10�3

�15 vol% Cu 45.7 (18.0 at 6 vol%) 4.70 � 10�6 2.03 � 10�5 3.95 � 10�4

�15 vol% Ag 19.8 1.91 � 10�5 8.37 � 10�5 1.80 � 10�3

�20 vol% Cu6Sn5,(produced in situ fromadded Cu and Sn in thesolder)

17.6 7.6 � 10�6 (at25 MPa)

9.8 � 10�5 (at13 MPa and85 °C)

5.8 � 10�4 (at12 MPa and125 °C)

Adapted from Guo and Subramanian [2002]

Table 5.15 High cycle fatigue life (expressedas cycles to failure) of composite solders,determined at different stress levels, reversed30 times/min (0.5 Hz)Solder 28 MPa 34 MPa 42 MPa

Pb-63Sn 42,000 18,000 4,000�13 vol% Cu6Sn5(a) 116,000 32,000 10,000�18 vol% Ni3Sn4(b) 285,000 167,000 54,000

(a) Cu6Sn5 dispersoids, 0.42 μm (16.5 μin.) in diameter. (b) Ni3Sn4 dispersoids,0.25 μm (10 μin.) in diameter

Table 5.16 Low cycle fatigue life (expressed ascycles to failure) of composite solders,determined at strain rates of 0.05 and 0.005 Hzand a total strain of 1%Solder 0.05 Hz, 1% strain 0.005 Hz, 1% strain

Pb-63Sn 980 2100�13 vol% Cu6Sn5 750 460�18 vol% Ni3Sn4 160 75

Table 5.17 Tensile properties of composite solders

Solder0.2% offset yield,

MPaUltimate tensilestrength, MPa

Elastic modulus,GPa

Elongation tofailure, %

Reduction inarea, %

Pb-63Sn 35 37 12 48 331�13 vol% Cu6Sn5 52 60 21 18 37�18 vol% Ni3Sn4 63 73 24 15 37


gap variation when attempting to make multiplejoints in parallel, for example, flip-chip inter-connection [McCormack, Jin, and Kammlott1994]. Iron-containing composite solders couldbe of interest for microelectromechanical sys-tems (MEMS) fabrication.

At the time of writing, neither grain-refined,dispersion-strengthened, nor composite soldershave yet been adopted in commercial use byindustry. Although they are undoubtedly attrac-tive for some applications, because of the prob-lems outlined previously, there is clearly a re-quirement for further research before thetransition can proceed with confidence.

5.6 Reinforced Solders(Solder Composites)

Reinforced solders are filler metals that in-corporate a reinforcing medium that is physi-cally large in relation to the joint width. On asimplistic level, they can be considered as visu-ally equivalent to the use of steel bars withinreinforced concrete constructions. Because thestrengthening mechanism is bulk physical con-straint of the solder by the reinforcement, it isfundamentally different from themechanisms forstrengthening solders described in section 5.5 ofthis chapter.

Reports in the published literature reveal someof the results of attempts made to load solderalloyswith a uniformdistribution of strong (high-modulus) particles or fibers (typically, 100 μm to1 mm, or 4000 μin. to 0.04 in., in size) of non-

metallic materials. Refractory metals have beentried, but these tend to agglomerate when thefiller is molten, because of their higher density[Ho and Chung 1990]. The most successful av-enue to date has been to incorporate into the fillerchopped carbon fibers, electroplated with nickelor copper so they are wettable by solder. Withlead-tin eutectic solder, the results show a three-fold enhancement of the shear and tensilestrength of the joints with respect to the unmodi-fied fillers and, more particularly, a significantreduction in the thermal expansivity of the filler.

The published data show that the tensilestrength of joints formed with lead-tin with re-inforcement can be as high as 65% of the rule-of-mixtures for the relative proportions by vol-ume of solder and reinforcement present.Approximately 15% by volume of fibers is themaximum that can be incorporated into the sol-der while retaining acceptable workability in themolten state. With this level of fiber loading,joint strengths in excess of 250 MPa (5.2 lb/ft2)at room temperature can be achieved [Ho 1996].

An alternative approach to attempting to fill ajoint with a molten alloy loaded with insolublefibers is to pack the joint with dry fibers and usea conventional solder, with or without flux, asrequired, to infiltrate and fill the interstices. Ob-viously, the fibers need to be metallized so theycan be wetted by the solder. Much higher load-ings of reinforcement can then be achieved—upto 55% of the volume of the solder. As alludedto previously, because carbon fibers have a smallbut negative coefficient of thermal expansion, at42% by volume, the thermal expansivity of thesolder composite declines to zero over the tem-

Fig. 5.28 Tensile strength of lead-tin eutectic solder with and without 3 vol% of Al2O3 dispersoids as a function of strain rate,measured at 80 °C (176 °F)


perature range of 25 to 100 °C (77 to 212 °F).Solders reinforced in this manner might there-fore offer benefits where there is a requirementto join low-expansivity materials, such as ce-ramics to metals, where a wide joint gap is man-datory. The high thermal expansivity of a con-ventional solder alloy (typically, >20 � 10�6/K,or 20 ppm/K) introduces a shear stress at thecomponent/solder interface, which is overcomeby using a carbon-fiber-loaded solder. The re-duction in the thermal expansivity brought aboutby the fiber addition reportedly accounts for athreefold enhancement in fatigue life on thermalcycling that is observed in bonded assemblies ofthis type.

An additional benefit of reinforced solders isthat carbon fibers can have very high longitudi-nal thermal conductivity, so the presence of amatof fibers in a joint can help to redistribute localhot spots within the plane of the joint. Clearly,there is scope for further research in this area.

When contemplating using fiber-reinforcedsolders, one of the key targets is to obtain avoid-free joint; otherwise, poor joint filling miti-gates the strengthening effect. This end is greatlyassisted when the infiltration of solder into thefiber bundle is promoted not only by metallur-gical wetting of the solder, but when surfacetension forces are exploited to achieve sponta-neous infiltration into the interstices. This hasbeen studied from a theoretical standpoint, albeitsimplified, and some of the key results are pre-sented in Table 5.18. In summary, provided thatthe wetting angle of the solder to the metalliza-tion on the reinforcement material is below 45°,then spontaneous infiltration should take placeirrespective of the aspect ratio of the reinforce-ment.

If the reinforcement medium is not closelypacked, then the critical wetting angle decreasesaccordingly. The corollary is that unless themini-mumconditions given inTable 5.18 are achieved,

the resulting joint will contain voids, unless ex-ternal pressure is applied to force the moltenmetal into the interstices of the reinforcementmaterial [Yang and Xi 1995].

5.7 Mechanical Properties andNumerical Modeling of Joints

This section considers methods for quantify-ing the mechanical integrity of joints and pre-dicting their dimensional stability under speci-fied environmental conditions. These are reallytwo independent issues that are addressed sepa-rately. Although there have not been recent sig-nificant advances in methods of measuring themechanical properties of solders and joints, val-ues for various material parameters, such asYoung’s modulus, are required for numericalmodeling techniques. It is therefore important tohave an appreciation of the limitations inherentin the derivation of parameters and the scope oftheir applicability. Modeling of the lifetime ofjoints, when subject to cyclic conditions, hasmade considerable progress in recent years andis also considered in this section.

5.7.1 Measurement ofMechanical Properties

Measurement of the mechanical properties ofbulk solder specimens and soldered joints shouldbe an uncomplicated process. Suitable test meth-ods, specimen designs, and methods of prepa-ration for metallurgical specimens are all definedby standards. However, even obtaining whatmight appear to be relatively straightforwardproperty data on bulk solders is fraught withproblems. A few of the factors that influence themechanical strength of solders include:

• Test method used• Procedure used to prepare the samples

Table 5.18 Calculated critical angle for a liquid to spontaneously infiltrate the interstices inselected close-packed structures, and the minimum packing density necessary to achieve filling evenwith perfect wetting. Above the minimum packing density, spontaneous infiltration is relatively easy toachieve, even when the wetting is relatively poor.

Reinforcement typeCritical wetting angle

for spontaneous infiltration, degrees

Minimum volume fraction ofreinforcement for spontaneousinfiltration at 0° contact angle, %

Unidirectional fibers 45 40Body-centered cubic packed mono-sized

spheres65 20

Face-centered cubic/hexagonal close-packedmono-sized spheres

50 40


• Precise sample geometry• Test temperature• Strain rate employed in the test• Microstructure of the filler metal

These sensitivities help to explain why pub-lished values of the properties of solders can varyby at least an order of magnitude for what, to theinexperienced observer, might otherwise appearto be identical samples [Plumbridge 1996]. Thecomplexity of solder behavior, especially that ofthe softer indium-containing alloys, arises partlyfrom the fact that at room temperature these al-loys are working close to their melting point,expressed in Kelvin. This means that atomic dif-fusion can occur rapidly, and hence, the pro-cesses of annealing, alloying, and precipitationare observed during testing, in addition to con-ventional bulk metallurgical phenomena such asnecking. In actuality, the response of a solder tomechanical stress is a complex combination ofelastic, anelastic, and viscoplastic behavior.

Obtaining consistent values for the mechani-cal properties of soldered joints is even moredifficult. In addition to the variables cited pre-viously in respect to bulk solders, the strength ofa joint is additionally influenced by a number offactors, including:

• Dimensions (thickness and area)• Method used for making the joint• Heating excursion (integrated temperature

and time)• Cooling rate• Substrate materials• Impurity content of the filler• Age of the joined assembly

Variables such as strain rate and temperaturealone can be manipulated to give joint strengthsthat vary by more than two orders of magnitudefrom identical samples prepared under rigorouslaboratory conditions [Jones et al. 1997; ITRI1987].

Notwithstanding the previously mentionedconsiderations, a further issue then arises as towhether the mechanical properties measured aresatisfactory in the context of the service require-ments of the assembly. For example, a solderjoint between a silicon die and a ceramic packagemay achieve a shear strength of, say, 40 MPa(5800 psi). Because the die weighs only a fewgrams and there are nomechanical contactsmadeto it, then clearly, the joint is more than strongenough to hold the die in place. The majority ofsoldered joints in electronic products are not

made to meet load-bearing requirements but toeffect electrical connectivity, thermal conductiv-ity, or a hermetic seal between components. Cor-relation between these variables and basic me-chanical properties (e.g., tensile and shearstrength) is often obtuse. Failure of the joinedassembly to pass a die-shear test is only an in-dicator that it falls short of a minimum require-ment.Ajoint can possess high levels of voids andtherefore possess impaired local thermal con-ductivity and leakage paths, giving rise to her-meticity failure long before this is reflected in itsshear strength.

In the electronics and photonics industries, thedifficulty of obtaining consistent mechanicalproperty data and correlating this informationwith the parameters of real interest have beenrecognized, and, as a result, simple mechanicaltesting has largely fallen out of favor. Instead, itis more common to build either complete prod-ucts or representative subparts and subject themto some form of accelerated or extended test thatis considered to encompass the rigors of life inservice. The components are then assessed forsigns of functional degradation.Any deficienciesthat are detected and can be attributed to unsat-isfactory joints are addressed accordingly.

More information about mechanical propertytests can be found in the planned companionvolume Principles of Brazing. Brazed joints areoften expected to carry mechanical loads, andhence, issues such as the choice of test methodused to measure the mechanical integrity of ajoint become more directly relevant.

5.7.2 Numerical Modeling of JointsComputational methods are finding growing

use in modeling how components and productswill respond to various stimuli, such as changesin temperature, mechanical loads, and exposureto chemicals. These naturally include consider-ation of soldered joints.

There are basically two types of model: thosethat calculate the dimensions of the product, andothers that predict the lifetime of joints. Modelshave also been devised that describe the profileof molten solders wetted onto solid substrates.An example of this type of model is referred toin section 5.2.6 in this chapter.

5.7.2.1 Dimensional Stability ofSoldered Joints

One of the better known methods of calcu-lating the dimensional stability of joints is finite-


element analysis (FEA). In FEA, a component orassembly is modeled in a geometrical manner interms of a mesh of smaller units or elements,with other dimensions of each element scaling toa set of properties of interest. The visible meshusually represents two or three orthogonal di-mensions of the part, while the other dimensionsare used to set how each element can respond tostimuli. Common parameters areYoung’s modu-lus for mechanical behavior, and thermal expan-sivity and thermal conductivity for thermal be-havior. In a realistic model, these parameters arenot fixed but vary with temperature. Constraintsare then applied to the mesh, with matrix algebraused to obtain a comprehensive and numericallyconvergent solution. The large number of cal-culations that are needed requires a computer forthis task.

There are now a number of commercial FEAsoftware packages available to suit a range ofapplications. A FEA prediction of the deforma-tion that occurs in a ceramic-metal bond on cool-ing from the solidus temperature of the filler isgiven in Fig. 5.29.

Despite the complexity of the modeling tech-niques, simplifyingassumptionshave tobemade.Often, an axis of spatial symmetry is defined,and only half or one-quarter of the assembly ismodeled. The oversimplification represented bythis assumption can easily be demonstrated formany situations, especially when transient con-ditions and thermal gradients are taken intoconsideration. Also, when these models are ap-

plied to soldered assemblies, it is normally as-sumed that the joints are uniform in their com-position and physical properties. This is patentlynot correct. A further restriction of applicabilityarises because most modeling programs assumethat the interface between the components andthe solder is relatively abrupt, whereas, in prac-tice, it is a highly complex region that hassignificant compositional and microstructuraldifferences, generated by reaction across inter-faces, and often these are not stable with time,temperature, or stress. Fortunately, the greatersophistication of software and increased com-puting power are enabling these features to bebuilt into FEA models.

To successfully apply such models requiresknowledge of the properties of materials in anassembly and how those properties change indifferent environments. For mechanical models,the environmental variables are usually stress,strain, and temperature. Thus, there is often a callon the joining technologist to provide data on themechanical and other physical properties of fillermetals. For the reasons outlined in the precedingsection, this is almost an impossible requirementto fulfill reliably. However, the potential savingsafforded by modern computer modeling tech-niques in their ability to assist in achieving right-first-time designs and thereby speed product timeto market means that there is often considerablepressure to provide appropriate material prop-erty data! Table 5.19 represents an attempt toprovide indicative data of the bulk properties of

Fig. 5.29 Finite-element analysis prediction of the geometry of a ceramic-metal brazed joint, at its periphery, at the solidustemperature of the filler alloy and on cooling to room temperature


a few common solders and some of their com-mon constituents at room temperature.

When running a model, the values given inTable 5.19 can be taken as a starting point andmodified to allow for the joint thickness, inter-metallic formation, compositional change, andthe properties of the components on either sideof the joint. Temperature-dependent terms canthen be added as appropriate. A literature searchwill usually throw out values for temperature-and strain-rate-dependent terms that can beadopted if they appear valid for the situationunder consideration. Often, it is necessary to de-sign and test joints of simplified geometry, inorder to provide confidence in the values andtheir sensitivity to second-order effects.

While acknowledging the many limitations ofthe data, as mentioned previously, Table 5.19does reveal some interesting trends in comparingthe alloys with the constituent pure elements.Almost invariably, solders possess greatly infe-rior thermal and electrical conductivity, com-paredwith puremetals. These characteristics fol-low from their heterogeneousmicrostructure andphase boundaries, which impede the flowof elec-trons and phonons. The effect is more pro-nounced on thermal conductivities because ofthe dominant contribution of phonons, or latticewaves, to thermal conductivity at room-ambientand elevated temperatures. Phonons are morestrongly scattered by the inhomogeneous micro-structure than are electrons,which dominate elec-trical conductivity and low-temperature thermalconductivity. At first sight, it might be expectedthat the gold-rich solders would have propertiesnot dissimilar to pure gold. However, the largedifference in density between gold, on the onehand, and the alloying elements, on the other,means that in terms of volume fraction, the gold-

silicon eutectic alloy, for example, possessesclose to 20 vol% Si. The gold-tin eutectic solder,Au-20Sn, is not a mixture of the constituents butan alloy of two intermetallic compounds of goldand tin (AuSn and Au5Sn), both of which havevery different characteristics to the pure ele-ments. In general, if compound formation occursin the creation of solders, then the propertiescannot be predicted reliably without directknowledge of the mechanical and physical prop-erties of the constituent phases.

5.7.2.2 Prediction of Joint Lifetime

The traditional methods for predicting the life-time of a fabricated part and the joints containedtherein are accelerated and extended testing. Be-cause these are time-consuming and expensiveto undertake, much effort has been devoted toattempting to predict joint life by numericalmod-eling. The ability to do this is particularly rel-evant to the needs of the PCB industry, where itis crucial to know the lifetime of the product andalso for manufacturers to convincingly demon-strate to customers that the soldered joints willmeet long-term expectations of reliability.

The approach to this problem startedwith clas-sical fatigue theory, which was rapidly found tobe wholly inadequate for describing the behaviorof solders. Gradually, more and more rate- andtemperature-dependent parameters were addedto the models. Although the quality of the analy-sis steadily improved, unexpected outcomeswereas common as accurate predictions. Some test-ing, albeit less extensive, remained necessary tovalidate the model.

Over the last few years, an alternative ap-proach that uses a concept termed strain energydensity has been gaining favor [Morrow 1964;

Table 5.19 Indicative property values of selected solders and pure metals in bulk form at roomtemperature

Elementor solder

Hardness,HV

Tensilestrength,

MPa

Young’smodulus,

GPaPoisson’s

ratio

Elongationto failure,

%

Electricalresistivity,

� · cm

Thermalconductivity,

W/m · K

Thermalexpansivity,

10�6/K

Ag/Au 20 150 80 0.40 40 2 350 20In 5 5 10 0.45 50 10 80 25Pb 10 15 15 0.42 50 20 35 30Sn 15 30 50 0.35 30 15 70 25Ag-97In 5 5 10 0.45 50 10 50 25Ag-96Sn 15 60 40 0.35 30 15 50 25Bi-50Sn 25 60 . . . . . . 1 35 20 20Cu-99Sn 10 30 50 0.35 30 15 50 25In-50Sn 5 20 . . . . . . 50 30 20 20Pb-63Sn 15 40 30 0.25 30 15 40 25Au-3Si 100 300 80 . . . 1 20 30 15Au-20Sn 100 275 60 0.30 5 10 60 10


Vaynman and McKeown 1993]. In very simplis-tic terms, thismodel attempts to examinewhetherthe deformation of a joint on single or multiplestress cycles exceeds the ability of the solder toabsorb and/or dissipate the energy imparted to it,through creep ormicrostructural changes.Wheremicrostructural changes do occur, the model isable to account for the material change on sub-sequent cycles.Although its complexities are be-yond the scope of this book, the important pointto note is that the quality of the joint lifetimepredictions now being made generally accordwith experimental results. For examples of ap-plication of this method, reference may be madeto Lau [2000], Lau and R. Lee [2001], Lau et al.[2002], and similar publications.

5.8 Solders Doped withRare Earth Elements

A relatively new area of solder research con-cerns doping of traditional filler metals with rareearth elements. Some results are reported in theliterature, and patents have been filed in respectof certain composition ranges. Rare earth dopingof solders is of interest because it affords a pos-sible means of favorably modifying the charac-teristics of solders in the molten and solid statesand enabling them to approximate more closelyto the materials properties desired by the manu-facturing industry.Alarge proportion of theworkdone in this area has been in conjunction withlead-free solders. These solders are discussed insection 5.1 of this chapter.

The rare earth elements are so described be-cause they were originally thought to have a lowabundance in the Earth’s crust and to be difficultto win from it and to separate from each other.It is now known that lanthanum, cerium, andneodymium are actually more abundant thanlead, and vast ore reserves have been found inChina and the United States. There are thirty rareearth elements, which is really another name forthe elements contained in the lanthanide and ac-tinide series of group three of the periodic table.However, one element of the lanthanide series(promethium) andmost of the actinides are trans-uranium elements, that is, manmade and atomi-cally unstable. Sometimes, referencewill be seento a rare earth called mischmetal. This is actuallyan alloy containing a high concentration of therare earth elements in proportion to their naturalabundance, and therefore, its composition varies

with the ore from which it was obtained: mona-zite, xenotime, or bastnasite. Mischmetal issometimes given the chemical symbol M, and itcontains, typically, 50% Ce and 30% La. Thegeneral symbol used for rare earth is RE. Rareearth elements have the common attribute thatthey are extremely reactive toward other metalsand most atmospheres.

The addition of rare earth elements to soldershas a number of effects on properties that varywith concentration.

5.8.1 Effect of Rare EarthAdditions on Solder Properties

Recent research has shown that doping of sol-ders with rare earths appears to enhance theirproperties while introducing few drawbacks, andthe melting range of solder alloys is hardly af-fected (Fig. 5.30) [Ramirez, Mavoori, and Jin2002; Wang, Yu, and Huang 2002; Chen et al.2003]. Small additions of rare earths:

• Grain refine solders and produce within thema dispersion of hard, insoluble intermetallicparticles that benefits their mechanical prop-erties. Thus, for example, the addition of0.4% rare earths to the Ag-96.5Sn solderhalves the average grain size, and this pro-duces a 12% boost in strength and ductility,as can be seen in Fig. 5.31 [Wang, Yu, andHuang 2002]. Likewise, a 0.083% additionof rare earths to 3.33Ag-4.83Bi-91.84Sn sol-der has similar effects on grain size, strength,and ductility [Xia et al. 2002]. However, theeffect of rare earth doping and the associatedgrain refinement of themicrostructure on duc-tility appears to be solder specific—in thecase of tin-zinc eutectic, this property is ac-tually impaired (Fig. 5.32) [Wu et al. 2002].

Thermodynamic analysis shows that thedispersed intermetallic phase, in tin-base sol-ders, is based on the composition Sn3RE [Maand Yoshida 2002]. These intermetallicphases predominantly congregate at grainboundaries [Ying, Hongyuan, and Yiyu1994].

The agglomeration of rare earth interme-tallic compounds at grain boundaries is re-sponsible for considerably increasing the re-sistance of solders to creep at roomtemperature, as shown in Fig. 5.33 and 5.34,because creep in solders occurs primarily bygrain-boundary sliding [Chen et al. 2002].Progressively increasing the concentration of


the rare earth elements increases the prefer-ence for discrete compound formation, whichwould explain why the creep properties peakat 0.1% RE addition in 3.8Ag-0.7Cu-95.5Snsolder. Exposure to elevated temperature re-duces the benefit to be derived from grain-

boundary pinning by the rare earth interme-tallic compounds, because other materialtransport mechanisms take effect. Thus, at 65°C (149 °F), the creep-rupture life is onlydouble that of the simple ternary solder, andany advantage disappears completely above

Fig. 5.30 Solidus and liquidus temperatures of 3.8Ag-0.7Cu-95.5Sn solder as a function of the added rare earth (RE) concentration

Fig. 5.31 Tensile strength and elongation to failure of Ag-96.5Sn solders dopedwith cerium and lanthanum.

The samples were chill-cast ingots of solder, 20 mm (0.8 in.) longby 10 mm (0.4 in.) in diameter, tested at room temperature anda strain rate of 4 � 10�3/s. RE, rare earth

Fig. 5.32 Tensile strength and elongation to failure of Sn-9Znsolders doped with lanthanum. The samples were

chill-cast ingots of solder, 25 mm (1 in.) long by 5 mm (0.2 in.)in diameter, tested at room temperature and a strain rate of 5 �10�3/s. RE, rare earth

Fig. 5.33 Stress-rupture life of Sn-3.5Ag and Sn-3.5Ag-0.25RE solders for an applied stress of 20 MPa

(2900 psi) at 50° C (122 °F)

Fig. 5.34 Stress-rupture life of 3.8Ag-0.7Cu-95.5Sn solder asa function of the rare earth (RE) content for an

applied stress of 16.5 MPa (2400 psi) at room temperature


100 °C (212 °F). This result puts a questionmark on the practical benefit of the observedimprovement in creep resistance. Also, theavailable data pertain to tens of hours of testduration, not thousands of hours of servicelife, and this issue needs to be investigated.

• Decrease the wetting angle of the solder onmetal substrates and generally enhancespreading. The effect on contact angle, wet-ting force, and wetting time is shown in Fig.5.35 for silver-tin eutectic solder on copper.Rare earth doping of the Sn-9Zn solder im-proves its wetting on metal substrates in asimilarway (Fig. 5.36) [Wu et al. 2002]. Like-wise, small additions of rare earths have beenfound to reduce the contact angle of indiumsolders on silver under inert atmosphere, asdescribed in Chapter 3, section 3.3.8.4 (Fig.3.28). The tests described were conducted inair using flux. It is presumed that the role ofthe rare earths is to reduce the interfacialtension between the solder and flux [Wang,Yu, and Huang, 2002]. It would seem that thereactive rare earth elements also help desta-bilize oxide films on the surface of the solderand substrate. Certainly, once the rare earthconcentration exceeds approximately 0.2%,the modified solders are able to directly wetand bond to nonmetals, such as silica. Thishas been demonstrated for both silver-tin eu-tectic solder and the high-melting-point Au-20Sn alloy. The bonding mode is as de-scribed in Chapter 4, section 4.1.2.2, namely,migration of the rare earth element to thesolder/substrate interface and chemical bond-ing by reduction of the nonmetal surface.Unfortunately, the concentration of rare earthelements necessary to achieve this behaviorresults in a loss of many of the boosted

mechanical properties referred to earlier andimpairs the ductility of all solders [Ramirez,Mavoori, and Jin 2002].

5.8.2 Implications forSoldering Technology

Most results reported to date relate to bulksolder specimens. The strength of lap joints incopper testpieces was found to be indifferent torare earth doping. This somewhat disappointingresult was attributed to porosity in the joints,originating from the use of a simple organicflux for the joining operation. Whether this po-rosity is an intrinsic characteristic of rare-earth-doped solders or indeed merely due to inappro-priate choice of flux is uncertain, becausecurrently, there are few reported studies of me-chanical properties of joints made and assessedunder controlled conditions.

As far as the authors are aware, solders dopedwith rare earth elements cannot yet be found inmanufacturers’ catalogs.

Fig. 5.35 Contact angle and spread area of Ag-96.5Sn solders doped with cerium and lanthanum melted on copper at 300 °C (572°F) for 30 s under cover of RMA flux. RE, rare earth

Fig. 5.36 Wetting force of Sn-9Zn solders doped with lute-tiumon copper using rosin-activated flux at 245 °C

(473 °F) in air. RE, rare earth


5.9 Diffusion Soldering

In soldering, wetting of the component sur-faces is not always easy to achieve, and when itdoes occur, the resulting alloying between thefillerandcomponentscancauseexcessiveerosionof the parent materials, embrittlement of jointsdue to the formation of phases with inferior me-chanical properties, andotherundesirable effects.These problems notwithstanding, solders havethe singular merit of being able to fill joints of ir-regulardimensionsandproducewell-roundedfil-lets at the edges of the joint.

Diffusion bonding sidesteps the need for wet-ting and spreading by a filler metal (see Chapter1, section 1.1.7.2). Once formed, diffusion-bonded joints are stable to high temperatures,so that the service temperature of the assemblycan actually exceed the peak temperature of thejoining process without risk of the joint remelt-ing. While the formation of undesirable inter-metallic phases can also occur in diffusion bond-ing, because there are usually fewer constituentsinvolved, it is easier to select a safe combina-tion of materials. However, diffusion bondingtends to be limited as a production process,because it is not tolerant to joints of variablewidth, and moreover, its reliability is highlysensitive to surface cleanliness. High loads (typi-cally 10 to 100 MPa, or 1.4 � 103 to 1.4 � 104

psi) have to be applied during the bonding cycleto ensure good metal-to-metal contact acrossthe joint interface. Also, the duration of theheating cycle is typically hours, compared withseconds for soldering, because solid-state diffu-sion is much slower than wetting of a solid by aliquid. These factors and the absence of anysignificant fillets to minimize stress concentra-tions at the edges of joints (see Chapter 4,section 4.2.4) considerably limit the applica-tions of diffusion bonding.

There exists a hybrid joining process that com-bines the good joint filling, fillet formation, andtolerance to surface preparation of conventionalsoldering, together with the greater flexibilitywith regard to service temperature and metallur-gical simplicity that is obtainable from diffusionbonding. This process is called diffusion solder-ing, sometimesalso referred toas transient liquid-phase (TLP) joining. Its higher-temperature ana-logue, diffusion brazing, is an established joiningand repair process that has been used for decadesin the aerospace industry and is described in theplanned companion volume Principles of Braz-ing.

5.9.1 Process PrinciplesDiffusion soldering uses a thin layer, typically

5 μm (200 μin.) or less, of molten filler metal toinitially fill the joint clearance, but during theheating stage, the filler diffuses into the materialof the components to form solid phases, raisingthe remelt temperature of the joint. At this stage,isothermal solidification occurs, and further re-action proceeds by solid-state diffusion until theprocess cycle is complete. Due to the generationof liquid in the joint, the necessary applied pres-sures are much less than those required for nor-mal diffusion bonding and are typically in therange of 0.5 to 1 MPa (70 to 140 psi). Diffusionsoldering therefore provides the ready means tofill joints that are not perfectly smooth or flat (afeature of liquid-phase joining) while offeringgreat flexibility with regard to process tempera-ture in relation to the service temperature of theproduct. Because of these features of the processand also the precise conditions used in imple-menting it, especially the controlled thinness ofthe layer of low-melting-point filler and speci-fied loading applied to the joint, the followingadditional advantages are obtained:

• Good reproducibility of joints• Excellent joint filling that applies to both

small- and large-area joints. Joints free ofvoids ensure leaktightness, which is impor-tant in situations where the joint is made toprovide a seal.

• Tight control of edge spillage, which can bekept to a minimum

• Attainment of very narrow joints, typicallyless than 10 μm (400 μin.), which benefits themechanical properties, as comparedwith con-ventional solder

Exploratory work on diffusion soldering wasdone in the mid-1960s [Bernstein and Bartho-lomew 1966; Bernstein 1966]. This activity,which was described as solid-liquid interdiffu-sion bonding, was primarily concernedwith low-temperature bonding of power semiconductor dieinvolving indium and gold. In the processes thatBernstein investigated, the emphasis was on lowtemperatures, typically below 160 °C (320 °F),and the endpoint of the process left thick inter-metallic phases in the joint, which tended to com-promise the mechanical integrity, although it isadequate for the intended function.

Alloy systems suitable for diffusion solderingwill possess a phase constitution that includes arelatively low-melting-point constituent—ideally, a eutectic reaction—to initiate the melt-


ing process and a higher-melting-point phase orsolid solution on which to terminate solidifica-tion. Wilde and Pchalek [1993] identified thebinary combinations of the precious metals (sil-ver, gold, palladium) and also copper and nickel,together with either tin or indium as the low-melting-point constituent, as among the mostsuitable systems for diffusion soldering. Some ofthe process details for these combinations, asreported in the published literature, are listed inTable 5.20.

In most of the published studies of diffusionsoldering, the emphasis has been on limiting theprocess temperature to make the joining opera-tion suitable for the attachment of dies and otherelectronic components, which are temperature-sensitive. In consequence, a compromise has hadto be struck whereby intermetallic phases aretolerated at the expense of mechanical strength.This trade-off is usually acceptable, because thestrength requirement in die attachment is gen-erally low. For example, the gold-indium jointsobtained by Wang et al. [2000], which achieveda shear strength of 12 MPa (1740 psi), weresufficiently strong to satisfy MIL STD 883E,method 2019.7 acceptance criteria.

At first sight, the cost of the precious metals,especially gold, might be considered an impedi-ment. However, the depth of interaction of thesemetals with the filler metal is shallow in diffu-sion-soldering processes, so that they can be ap-

plied as thin metallizations (<10 μm, or 400 μin.)to the component surfaces without the danger ofthe component materials entering the reaction.The contribution to component cost from suchthin layers is relatively small. The use of copper,silver, or gold as surface coatings confers theparticular advantage that, being relatively noble,they are readily wetted, even in slightly oxidiz-ing atmospheres. This makes diffusion-solderingprocesses using these metals relatively tolerantto the condition of the atmosphere in which thejoining operation is carried out. In the publishedreports referred to in Table 5.20, the authors em-phasize their ability to achieve satisfactory jointswithout the use of fluxes in moderately protect-ing atmospheres, such as a shroud of nitrogen.

The sequence of steps involved in making adiffusion-soldered joint is shown schematicallyin Fig. 5.37.

5.9.2 Diffusion Soldering of SilverThe authors have found the silver-tin system

to be one of the most versatile for diffusion sol-dering of components for use in the electronicsindustry, because well-filled and ductile jointscan be produced using compressive loads of aslittle as 0.5 MPa (70 psi) [Jacobson and Hump-ston 1992]. Silver-tin solder of eutectic compo-sition reacts with silver to form the Ag3Sn phaseas a continuous interfacial layer, which is both

Table 5.20 Selected diffusion-soldering systems and process details, as reported in the publishedliterature

SubstrateFillermetal

Maximumfiller

thickness,μm

Processtemperature,

°C Time, minApplied

load, MPa

Remelttemperature,

°C Ref

Copper Indium 1.5 180 4 0.5 >307 Sommadossi et al. [2000]Copper Tin 1.0 280 4 0.5–1.0 >415 Bartels et al. [1994]

5.0 300 20 Notspecified

>676 Kato, Horikawa, andKageyama [1999]

30–50 300 300 Notspecified

>676 Kang et al. 2002

Silver Tin 5.0 250 60 1.0 >600 Jacobson and Humpston[1992]

2.0 250–350 10 0.17 >600 Wilde and Pchalek [1993]Silver Indium 6.0 175 120 0.5 >880 Jacobson and Humpston

[1992]Gold Tin 4.0 450 60 1.0 >900 Humpston, Jacobson, and

Sangha [1993]2.25 310 13 0.28 >278 Matijasevic, Lee, and

Wang [1993]2.0 260 15 0.3 >278 Lee and Wang [1992]

Gold Indium 5.0 200 0.5 1.25 >495 Wang et al. [2000]2.0 160–240 10 0.05 >495 Wilde and Pchalek [1993]

Nickel Tin 1.8 300 6 0.8 >400 Khanna, Dalke, and Gust[1999]

1.8 400 2160 0.8 >977 Khanna, Dalke, and Gust[1999]


tough and strongly adherent to the other phasesin this alloy system. The rate of reaction betweensilver and molten tin has been characterized andis represented graphically in Fig. 2.35. The con-trollable nature of the alloying reaction in theconventional soldering system is indicated bythe general profile of the erosion curves, whichshow that the reaction is self-limiting in char-acter, within the context of realistic processingcycle times and temperatures. If a thin layer oftin, typically 5 μm (200 μin.) thick, is sand-wiched between two components, each coveredwith a 10 μm (400 μin.) thick layer of silver, andheated to 250 °C (480 °F), the tin will melt andreact with the silver to formAg3Sn. On continuedheating, the tin is progressively converted to thiscompound until no liquid tin remains. By keep-ing the tin layer thin, it completely reacts to formsolid phases at the joining temperature in lessthan 1 min.

The remelt temperature of a silver-tin diffu-sion-soldered joint is determined by the phasesthat are present. Immediately after the liquid tinhas been consumed by reaction, the remelt tem-perature is that of theAg3Sn compound, which is480 °C (895 °F). Longer heating times promote

continueddiffusionof tin fromtheAg3Sn reactionzone into the silver. Consequently, the width ofthis zone decreases as it is replaced first byAg5Sn(�) and, ultimately, by a solid solution of tin in sil-ver, as anticipated from the silver-tin phase dia-gram given in Fig. 2.9. This progression is illus-trated by the series of micrographs shown in Fig.5.38. As the reactionwith the silver proceeds, theremelt temperature risesprogressively toward themelting point of silver (962 °C, or 1764 °F). Me-chanical propertymeasurements have shown thatthe shear strength of diffusion-soldered jointscontaining theAg3Snphase is close to the 25MPa(3600 psi) value for conventional soldered jointsmadewith theAg-96.5Sneutectic solder to silver-coated components. As the joint microstructureconverts to a silver solid solution, themechanicalproperties shift in tandemtoward thoseofpuresil-ver, with the shear strength increasing toward 75MPa (11,000 psi). This can be significant from anapplicationspoint ofview,because the strengthofsilver is approximately three times that of theAg-96.5Sn eutectic alloy, which itself is superior inthis respect to the common Pb-62Sn solder atroom temperature by a factor of two to three[Harada and Satoh 1990].Another attractive fea-tureof the silver-tinalloysystemfordiffusionsol-dering is that there is negligible volume contrac-tion as the reactionproceeds,which is a fortuitousconsequence of the closely similar specific vol-umes of the various phases. Therefore, the ten-dency to formvoidsorcracksasa resultofvolumechange is minimal.

Diffusion-solderingprocessesarenot routinelyencountered but are used commercially. One ex-ample is as a method of attachment of siliconpower devices tomolybdenumheat sinks [Jacob-son and Humpston 1992; Humpston et al. 1991].Replacing brazed joints made using an industrystandard process, involving the Al-12Si alloy,with silver-tin diffusion soldering provided ameans for reducing the process temperature fromover 600 to 275 °C (1112 to 527 °F), which de-creased the bimetallic, center-to-edge bow by240% for a typical 50mm (2 in.) component. Be-sides silver-tin, a silver-indium diffusion-solder-ing process is an alternative, offering a lower pro-cess temperature [HumpstonandJacobson1990].However, the associated processing involved ismorecomplex—theplatingof indiumislessstan-dardthanthatof tin,andthemorerefractorynatureof indiumoxidesmakes it necessary to apply spe-cial surface treatments to exposed indium sur-facesprior to thebondingoperation[Sommadossiet al. 2002].

Fig. 5.37 Schematic illustration of the steps involved inmak-ing a diffusion-soldered joint


5.9.3 Diffusion Soldering of Gold

The gold-tin alloy system has provided thebasis for the diffusion-soldering process for join-ing items of high-carat gold jewelry below 450°C (842 °F). The traditional gold jewelry manu-facturing route involves the use of the so-calledcarat gold solders, which are actually brazingalloys with working temperatures above 800 °C(1470 °F) [Rapson and Groenewald 1978; Nor-mandeau 1996]. The high temperatures involvedare detrimental to the mechanical strength ofhigh-carat gold jewelry, because they anneal andsoften rapidly when heated above approximately450 °C (842 °F). Further details on carat goldsolders and themetallurgy of gold jewelry can be

found in the planned companion volume Prin-ciples of Brazing. Diffusion soldering provides asatisfactory alternative joining process. In trials,it was found that a tin coating 3 to 4 μm (120 to160 μin.) thick was generally sufficient to ensurecomplete joint filling and the formation of smalledge fillets. Provided that the peak process tem-perature exceeds 419 °C (786 °F), the meltingpoint of the AuSn intermetallic compound, thetin will transform initially to the high-gold in-termetallic compound Au5Sn and, on continuedheating, to gold solid solution. The Au5Sn com-pound contains approximately 90wt%Au and someets the 18 carat requirement of the jewelryitem. Prolonged heating is undesirable, becauseit results in softening of the gold assembly, as

Fig. 5.38 Series of micrographs showing the progressive change in joint microstructure that occurs on making a diffusion-solderedjoint using tin in combination with silver metallizations applied to copper substrates. 400�


reflected by the grain growth, and also in Kirk-endall voiding in the centerline of the joint. Onehour at 450 °C (842 °F) under a compressiveloading of 1 MPa (145 psi) was found to be anacceptable compromise for the processing con-ditions [Humpston, Jacobson, and Sangha, 1993].Figure 5.39 shows a bracelet and matching ear-ring set that was assembled by this method andexhibited at theWorld Jewelry Trade Fair held inBasel in 1992.

5.9.4 Diffusion Soldering of CopperCopper-tin and copper-indium are also suit-

able systems for diffusion soldering. However,when the copper-tin joining process is carried outbelow 676 °C (1249 °F) and the copper-indiumprocess is operated below 631 °C (1168 °F), theyresult in the formation of planar intermetallicphases that have limited fracture toughness andare responsible for relatively weak joints. In thecase of copper-tin, these intermetallic phases areCu6Sn5 (�) and Cu3Sn ( ), while Cu7In4 (�) andCu7In3 (�) are found in the copper-indium sys-tem. The Cu3Sn ( ) and Cu7In3 (�) phases formadjacent to the surface of the copper layer [Kato,Horikawa, and Kageyama 1999; Kang et al.2002].

In the copper-tin process, it has been shownthat the formation of the brittle intermetallics canbe suppressed by raising the joining temperatureabove the melting point of the Cu3Sn interme-tallic (676 °C, or 1249 °F). The successful dif-

fusion-brazing process that has been developedusing this approach is described in the plannedcompanion volume Principles of Brazing (seealso Sangha, Jacobson, and Peacock 1998).

An interesting variant of this diffusion-soldering process has been investigated wherebythe tin solder in the copper-to-copper joint isreplaced by the In-49Sn eutectic alloy (meltingat 120 °C, or 248 °F). The joining operationswere carried out at temperatures up to 400 °C(752 °F) [Sommadossi, Gust, and Mittemeijer2002]. In this case, two intermetallic compoundsform by reaction, but these are different from the� and � phases produced in the absence of in-dium. They are both ternary alloys. One of theseis based on the Cu10Sn3 (�) phase, which, in thebinary alloy system, is only stable at elevatedtemperature but is stabilized at room temperatureby the addition of indium. The other phase isdesignated (confusingly, here) as � but is basedon the Cu2In rather than the Cu6Sn5 intermetallic.This phase can also dissolve the third element, inthis case, tin. Below 200 °C (392 °C), only the� phase grows, and above this temperature, bothgrow together, with the � phase steadily out-growing its sister phase. The joints made wererelatively thick (typically 50 μm, or 2 mils),which did enable significant dilution of the in-dium and tin in copper to occur, so as to dissolvethe intermetallic phases. However, jointsmade at350 °C (662 °F) using 10 μm (400 μin.) thicksputtered layers of In-49Sn thatwere etch cleanedprior to bonding achieved a shear strength inexcess of 155 MPa (22,500 psi) and a tensilestrength of 160 MPa (23,200 psi) [Sommadossiet al. 2002]. These relatively strong joints con-tained only the homogeneous � [Cu10(Sn, In)]phase, which grows entirely by solid-state dif-fusion and has a relatively small grain size.

5.9.5 Practical AspectsThere are often practical difficulties with ap-

plying the layer of filler as an electroplated orvapor-deposited coating to the intended joints,including the need to mask off other areas of thesurfaces of the components. It has often provedmore convenient to use a foil preform of therelevant precious metal, typically 25 to 100 μm(1 to 4 mils) thick, that is coated with the nec-essary thickness of the low-melting-point filler(tin or indium). An appropriate area of the foil iscut out and sandwiched in the joint gap. The useof a foil of soft precious metal offers the furthermerit of acting as a stress absorber, which is most

Fig. 5.39 Parts of an 18 carat gold bracelet and matchingearring set assembled using the gold-tin diffusion-

soldering process at 450 °C (842 °F). The unusually low processtemperature enables the face plates to retain much of their work-hardened strength and thereby accept a particularly high surfacepolish. Each box of the bracelet measures approximately 8 mm(0.3 in.) wide. Courtesy of the World Gold Council


useful in situations where the parts to be joinedhave local topography or significantly differentcoefficients of thermal expansion. This latter as-pect is treated in greater detail in the plannedaccompanying volume devoted to brazing. Be-cause of the higher temperatures involved inbrazing operations, thermal expansion mismatchstresses can be a more serious problem in thatcontext.

5.9.6 Modeling ofDiffusion-Soldering Processes

Some attention has been devoted to the theo-retical modeling of transient liquid-phase joiningprocesses, but the published studies to date havebeen confined to the higher-temperature diffu-sion-brazing process [Isaac, Dollar, and Massal-ski 1988]. The analytical models generally as-sume that the process kinetics are governed bydiffusion, so that the phases that solidify from themelt at the joining temperature grow at a rate thatis proportional to the square root of the bondingtime. Clear evidence for the classical diffusion-controlled relationship has been provided for thecopper-tin system at 300 °C (572 °F) [Kato,Horikawa, andKageyama 1999]. However, therehas been little systematic work in modeling thekinetics of the various solid phases that grow andsubsequently are replaced by other solid phases,or of the primary solid solution, as usually occursin diffusion soldering. Such an endeavor wouldgreatly facilitate the design and development ofthis interesting joining method.

5.10 Advances in JointCharacterization Techniques

Of the many techniques available to assess theintegrity of soldered joints, particularly thoseused to attach electronic components to PCBs,three have benefited greatly from the advent ofcomputer technology. These are ultrasonic in-spection, x-ray inspection, and optical inspec-tion.

5.10.1 Ultrasonic Inspection (ScanningAcoustic Microscopy)

Ultrasound is defined as pressure waves withfrequencies higher than sound waves and thatcannot be heard—in practice, in the range of 0.5MHz to 5 GHz. The particular characteristic of

ultrasound that is exploited in nondestructivetechnology is its ability to travel through solidmaterials while obeying the same laws of re-flection and refraction as light. Because ultra-sound travels with fixed velocity through a givenmaterial, echoes produced by reflection at dif-ferent surfaces and interfaces will be temporallyresolved, and the corresponding distances canthen be calculated. This is the basic principle ofpulse-echo ultrasonic inspection.

In commercial instruments, ultrasound is gen-erated by a piezoelectric transducer mounted asa probe that is coupled to the surface of thetestpiece via a liquid or pasty coupling agent.Two or more probes tend to be used: one totransmit the pulse and the others to detect ech-oes. The higher the ultrasonic frequency, thehigher is the resolution, but the stronger is thesignal attenuation due to absorption by the ma-terials through which it travels. Ultrasonic sig-nals, like other forms of wave energy, can re-solve features down to approximately the sizeof a single wavelength. Accordingly, ultrasoundof 10 MHz frequency should be capable ofdetecting cracks down to 0.5 mm (0.02 in.) incross section in metal. This is clearly inad-equate for the inspection of defects in joints thatmay themselves be of comparable size or evensmaller, for example, those made to surface-mount electronic components. To improve onthis level of discrimination, higher frequenciesmust be used.

The scanning acoustic microscopy (SAM)technique has been developed to operate in thefrequency range of 20 MHz to 2 GHz and offersthe finest level of resolution of the ultrasonic testmethods, although the depth of penetration islimited to below 10 mm (0.4 in.).

The SAM technique is capable of nondestruc-tively assessing the distribution of voids, cracks,inclusions, and other hard phases over the areaof essentially parallel-sided joints [Matuasevic,Wang, and Lee 1990; Kauppinen and Kivilahti1991]. It involves focusing an acoustic wave,generated from a piezoelectric transducer, via asapphire lens onto the specimen and scanning itin a raster fashion. Changes in the reflected acous-tic signal from boundaries between features hav-ing different acoustic properties are recorded andmapped to produce the image. The correspon-dence that can be obtained between the imagesof a voided joint produced by x-radiography andby SAM is illustrated in Fig. 4.25 and 4.26.

The interpretation of a scanning acoustic mi-crograph of a joint can present difficulties, for


example, determining whether a certain featurecorresponds to a crack, void, or intermetallic par-ticle. Similarly, the need to mechanically scanthe probe in very close proximity to the surfaceof the components restricts the geometries thatcan be examined. Moreover, the limited depth ofsample from which clear images can be obtainedmeans that one of the components must be thin.

Modern SAM equipment overcomes all thesedifficulties by using a variable or multiple fre-quency system, often with multiple transducers,and by automating the probe positioning and im-age processing.Amodern SAM system is able todiscriminate between internal and externalboundaries of components and present the userwith a three-dimensional image that can be elec-tronically rotated and sliced similar to a com-puter-aided design drawing. By this advance,SAM has changed from being a laboratory di-agnostic tool to part of the suite of in-line quality-assurance methods essential to ensure low-defect-ratemanufacturing. Some examples of theuse of SAM to inspect the interior of electroniccomponents in a nondestructivemanner are givenby Adams [2001].

5.10.2 X-RadiographyThe x-ray spectrum comprises electromag-

netic radiation of short wavelengths in the range1016 to 1021 Hz. The high frequency and energyof x-rays enables them to penetrate materials andreveal internal features, including defects, pro-vided they absorb the radiation to a differentextent than does the surrounding material. It isthis characteristic that provides the contrast inthe x-ray image. Accordingly, voids, inclusionsof heterogeneous material, and cracks parallel tothe x-rays will be more conspicuous to this tech-nique than cracks and interfaces that are per-pendicularly oriented. Radiographs revealingvoids in joints are shown in Fig. 1.17 and 4.25.

In a modern industrial x-ray machine, a tele-vision camera system is employed in place oftraditional film. This enables the x-ray image ofthe object to be viewed in real-time. The com-ponent undergoing inspection is held by a free-space manipulator. The combination of move-ment of the testpiece coupled with real-timeviewing permits a comprehensive examinationto be made rapidly. The nature of visual percep-tion is such that an area that differs marginally incontrast from its surroundings is more easily de-tected when in motion, so that defects are morereadily noticed. Additionally, the controlled

movement enables defects to be viewed at anoptimal angle.

An important innovation in x-radiography isthe development of systems incorporating mi-crofocus sources. Small focal-spot sizes can beachieved at the x-ray target by electromagneti-cally focusing the incident electron beam, ex-ploiting a technique that is widely used in elec-tron microscope technology. The focal-spot sizecan be as small as 1 μm (40 μin.) in diameter. Thebenefits offered by a microfocus x-ray systemare:

• Fine rod anode sources can be inserted intohollow assemblies, such as the cavity of anoptoelectronic package, permitting single-wall exposures to be obtained in situationswhere conventional x-ray systems could onlyprovide double or multiple-wall radiographsbecause of their bulky tube heads. This sim-plifies the projection geometry of the radi-ography and increases the relative sensitivityof the radiograph to defects with respect tothe absorbing material of the assembly.

• There is the possibility of obtaining geo-metrically magnified images of high defini-tion by distancing the camera from thetestpiece. The magnification obtainable froman idealized point source, M, is given by theexpression:

M �t � Do

Do

This is illustrated in Fig. 5.40.

Fig. 5.40 Representation of geometric magnification in mi-crofocus x-radiography


Steady improvements in microfocus x-raytubedesignhave resulted in geometricalmag-nification of over 2000� being realized incommercial systems,with the tubes operatingin the low-kilovolt range.With further digitalprocessing of the optical image, total magni-fications of 7000� are now possible. This en-ables features of the order of 250 nm (10 μin.)to be resolved, and the system is then said tobe operating in the nanofocus mode.

• A diminution of the focal-spot size meansthat the geometric unsharpness is correspond-ingly reduced; that is, the precision withwhich edges can be resolved is improved.

• A consequential benefit of spatially separat-ing the image sensor from the testpiece is areduction in the scattered ray fraction gen-erated by the testpiece itself, which contrib-utes to the image.

• The x-ray beam emerging from a microfocussource can be profiled to give a controlledcone of radiation, again decreasing scatterand improving the sensitivity to flaws.

• Finally, by simplymoving the testpiece alongthe axis of the x-ray beam, it is possible tocontinuously zoom in on detail.

Multifocus x-ray tubes have been introducedthat enable the user to choose between micro-focus, nanofocus, and high-power modes, usingsoftware control and a single mouse click. Themicrofocus mode is preferred for applicationswhere the area being examined is millimeters ona side, such as inspection of circuit boards anddiscrete modules. By contrast, the nanofocusmode is suitable for examining flip-chip assem-blies and wire bonds. High-power settings areappropriate for denser samples, such as diemounted on heat studs and generally, bulkieritems where the boosted x-ray energy providesdeeper penetration of the radiation.

5.10.3 Optical InspectionOptical inspection of joints normally de-

scribes examination by eye, often with the aid ofa magnifying device. Usually, this will be anoptical microscope but may be stretched to in-clude scanning electron microscopy operated inits usual backscattered mode. Despite being avery old technology, optical inspection methodscontinue to improve at a remarkable rate. Thequality of the image that can be obtained from amodern microscope in terms of its resolution,depth of focus, field area, and working distance

is vastly improved compared to even only a fewyears ago. Unfortunately, these improvementscome with a price, and a standard objective lenscan easily carry a four-figure (dollars) price tag.Notwithstanding the fundamental improvementin lens design and manufacture, two relativelyrecent innovations in optical inspection are hav-ing an impact on soldering technology. These areautomated optical inspection (AOI) and endo-scopes.

Automated optical inspection is the product ofa highly successful marriage between optical anddigital electronic technology. In essence, the partto be inspected is placed under an optical mi-croscope, and the image is electronically pro-cessed to identify certain features, usually againstspecified pass/fail criteria. The viability of thisinspection method hinges greatly on the camerasystem. The camera used in an AOI system is adigital camera with a sensor having a minimumresolution of six million pixels. To maximize thequality of the information acquired by each pixel,it is exported directly to the computer in digitalformat, without conversion to a video signal.Very sophisticated analysis is now possible onthe acquired digital image. Some of the outputsthat are available from a commercialAOI systeminclude:

• Reading of component labels and markers tocheck that correct device types and valueshave been placed on a PCB

• Reading of serial numbers on larger compo-nents, for quality and archiving purposes

• Verification of component orientation andmeasurement of alignment

• Checking for solder bridges and lack of sol-der

• Measurement of contact angles of fillets andranking of surface reflectivity

• Validation ofwire bond patterns and isolationbetween adjacent loops

• Confirmation of electrical and optical cablerouting and termination

• Identification of regions requiring furthermanual inspection or rework

• Creation of an archive of board integrity

The continued development of improved op-tics, cameras, and more powerful software,coupled with improved processing speeds, canonly enhance the adoption of AOI as an integralpart of electronics and photonics assembly lines.

Endoscopes are essentially miniaturized mi-croscopes that use optical fibers or other meansto transmit the image from the location of the


examination to the viewing point. They arewidely used in medicine for conducting internalexaminations in a noninvasive manner. Endo-scopes have been developed for examining hid-den soldered joints. Examples include intercon-nects of ball grid arrays (BGAs), chip-scalepackages (CSPs), and flip-chip components. Cus-tom optical heads enable side-on viewing of fea-tures, while their small size and long workingdistance permit inspection of otherwise inacces-sible areas. Inspection of solder fillets, solderpaste print profiles, via-hole plating integrity, andconformal coating uniformity are but a few ex-amples of their application.

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Index

A

Abbreviations and symbols, 243Abietic acid, 116Acceptance criteria, 231Acetic acid vapors, 132(F)

for Pb-62Sn solders, 131as reactive atmospheres for fluxless soldering, 130toxicity of, 131

Acetylene, 34Acoustic microscopy, 204Activated filler metals, 151Activated solders, 8Activation energy, 8, 24, 124Activation temperature of fluxes, 117Active brazes, 103Active filler metals

for ceramics, 149contact angle in, 17high temperatures of, 149

Active hydride process, 151Active solders, 147, 152–153Additives, 135Adhesive bonding, 2–3Adhesively bonded joints, 3, 173Adhesives, 184Aesthetic requirements, 50Airborne particle size, 179(F)Alcohol solvent carrier, 118Allotropic transformation, 77(F), 195–196Alloy constitution, defined, 78Alloy Constitution (reference text), 78Alloy J, 60, 61Alternative atmospheres for oxide reduction, 111Alumina, 121

alloy matching thermal expansion, 164and gallium, 62nonmetallic bonding to, 152

Aluminum and aluminum alloysAl-94Zn solders with, 54brazes for, 6corrosion of, 54diffusion bonding of, 10diffusion soldering for, 137fluxes for, 121–122heat capacity of, 62as impurity, 76, 77oxides on, 9and temperature uniformity, 50thermal conductivity of, 62thermal expansivity of, 62thermal heat capacity of, 50zinc alloys with, 54zinc-bearing solders for, 61

Aluminum fluxes, 121Aluminum-gallium-magnesium-zinc solders, specific

typesAl-3Ga-3Mg-90Zn, lead-free solders, 66Al-3Ga-3Mg-90Zn, zinc-base solders as, 66

Aluminum-germanium eutectic alloy, 8Aluminum nitride, 152Aluminum quaternary alloy, 83Aluminum-silicon alloys, 6Aluminum-silicon alloys, specific type

Al-12Si, additions for wetting and spreading, 134Al-12Si, diffusion soldering processes with, 232Al-70Si, alumina matched thermal expansion, 164

Aluminum-silicon carbide composites, 152, 175Aluminum-zinc phase diagrams, 66(F)Aluminum-zinc solders, specific types

Al-94Zn, with aluminum, 54Al-94Zn, zinc alloys, 54Al-94Zn, zinc-bearing solders, 62

Amalgamsadvantages of, 214based on gallium, 216–217based on indium, 217based on mercury, 215–216defined, 214dental, stress-strain curve for, 215(F)as solders, 214–217solid and liquid metals evaluated for, 214(T)

Ammonia fluxes and intergranular corrosion withbrasses, 114

Anisotropically conductive adhesives, 3ANSI/J-STD-002 (test method), 207Antimony

added to lead-tin solders, 57as additive to indium, 135effects of, on surface tension of tin, 193as impurity, 76Pb-Sb-Sn ternary system, 57as solder constituent, 53–54and solid-solution strengthening, 53–54

Antimony-tin phase diagrams, 60(F)Application methods, 148Argon, 36, 109Arrhenius-type rate relationships, 24Ashby materials selection chart, 159(F)Asthma, 43Atmospheres. See also inert atmospheres

categories of, for joining, 103chemically active, described, 104chemically inert, described, 104controlled gas, 104effect of, on spreading, 20and fluxes, 8and heating method, 33



Atmospheres (continued)joining, interrelationship of, 104(F)nonoxidizing, and heat treatment for wetting, 148and oxide film reduction, 105–106reactive gas, and oxide reduction, 130–131reduced pressure, and zinc bearing-solders, 62, 65soldering, costs and benefits of, 116(T)soldering, thermal conductivities of, 114(T)types of, 104

Atmospheric corrosion, 37Atmospheric quality, 136(F)Atomic diffusion, 224Atomic fraction of constituents, 96Atomic hydrogen

creation of, 127reduction of silver oxide by, 130(F)solder oxides reduction by, 127–128

Atomically clean, 50ATV 2003, 127Automated optical inspection (AOI), 237

B

Bakeout temperature, 36Ball grid array integrated circuits (BGAICs), 179Ball grid arrays (BGAs), 199, 204Balling up, 62–63Balls (of lead-tin solders), 129(F)Barrier coatings, 133, 149Barrier metal, 200Barrier metallizations, 155Basic spreading test, 211Berthoud equation, 25Beryllia, 152Beryllia dust, 163Beryllium, 104, 216Bimetallic expansion, 50Bimetallic strip, 158(F)Binary alloys and phase diagrams, 79Binary compounds, 84Bismuth

added to lead-tin solders, 57as additive to indium, 135Bi-Pb-Sn ternary system, 57effects of addition to, on liquidus and solidus

temperatures of silver-tin off-eutectic alloys, 193(F)effects of addition to, on tensile strength of silver-tin

solders, 193(F)effects of, on surface tension of tin, 193expansion on freezing, 53in hermetic soldered points, 53as impurity, 76, 77as solder constituent, 53

Bismuth-antimony-tin system, 96Bismuth-bearing solders

dissolution of chromium metallizations to, 149inferior fluidity of, 53inorganic fluxes for, 53

Bismuth-lead-tin ternary eutectic solders, 93

Bismuth-lead-tin ternary system, 57Bismuth-silver-germanium alloys, specific type

Bi-11Ag-0.05Ge ternary, 197–198Bismuth-tin lead-free solders, 122Bismuth-tin phase diagrams, 56(F)Bismuth-tin solders, specific types

Bi-43Sn, 31Bi-43Sn, elongation, 194Bi-43Sn, for hermetic joints, 173Bi-43Sn, melting point of, 173Bi-43Sn, physical properties of, 194

Boiling/sublimation temperatures, 107Boiling, temperatures of, 109(T)Bolometers (thermal imaging), 35Boltzmann constant, 24, 78, 98Bond formation in pressure welding, 9Bond quality, 212Bond wire, 175Bond wire interconnections, 202Bonding process

at low temperatures, 214stages of, 9

Bonding temperature, peak, 38Boron nitride, 152Bow distortion

of bimetallic assembly, 165of bimetallic strip, 158(F)equation for, 157

Brassesand ammonia, 105intergranular corrosion and ammonia flux, 114vacuum atmospheres, and zinc metallization, 107

Braze alloy familiesand melting ranges, 7(F)temperature ranges of, 6

Braze alloy, specific typeAl-4Cu-10Si, 6

Brazes and brazingfiller metal temperatures in, 5service temperature of, 8

British Standards (BS) 9430 (void free joints), 183Brittle failure, 28Brittle intermetallic layers, 28Brittle joints, 52. See also embrittlementBrittle materials, 28Brittle phases, 49Bulk filter mechanical properties, 29Bulk properties of solders, 225–226Bulletin of Alloy Phase Diagrams (phase diagrams), 79Butt joints, 177, 177(F)Butt welding, 9

C

C-charts, 42C4 process, 199Cadmium and cadmium alloys

health hazard of, 63restrictions in use of, 50, 51




toxicity of, 51, 191Cadmium-base solders

melting point depression, 93melting points for, 93

Cadmium-indium-tin-zinc quaternary eutectic system,93

Cadmium-indium-tin-zinc quaternary system, 94(T)Cadmium solders as substitutes for mercury, 93Calibration standards, 212–214Capillary action, 23Capillary dams, 23Capillary flow

in narrow gaps, 15time for molten tin and copper in, 27(F)

Capillary forces, 15, 15(F), 212Carbon-carbon fiber composites, 152, 163Carbon dioxide, 109Carbon-fiber-loaded solders, 223Carbon fibers, 222Carbon monoxide

as reactive atmospheres for fluxless soldering, 130uses of as reducing gases, 109

Carboxylic acids, 116Cast iron, 147Cathodic sputtering, 180CE7 alloy, 163Ceramics

active filler metals for, 149soldering to, 149wide-gap joints in, 158zinc oxide, metallization of, 150

Cerium, 227as additive to indium, 135

CFC-free cleaning operations, 40CFCs. See chlorofluorocarbonsChadwick peel tests, 73Channeling, 22Chemical cleaning, 37, 209Chemical displacement, 180Chemical fluxes. See fluxesChemical properties of lead-free solders, 193–196Chemical reduction of metal oxides, 105Chemical vapor deposition (CVD)

metallization techniques, 180on nonmetallic components, 149–150

Chemically active atmospheres, 104Chemically inert atmospheres, 104Chip-scale packages (CSPs), 199Chloride-based fluxes, 121Chlorides, 111Chlorofluorocarbons (CFCs), 39, 111, 115, 117, 118,

196Chopped carbon fibers, 222Chrome oxide, 110Chromium, 150Chromium metallizations, 149Clausius-Clapeyron equation, 97, 98Clausius’ theorem, 138Clean room class designation, 179(F)

Cleaningalternatives to, 41benefits and costs of, 41CFC-free, 40chemical, 37costs of, 118electronic assemblies, 40measure of effectiveness of, 119–120mechanical, 37methods of, 40postjoining, 39–41by reverse-gas bias mode, 150

Cleaning agentschlorides as, 111fluorides as, 111relative effectiveness of, 41(F)

Cleaning treatments, 37Cleanliness of IEC board test coupons, 119(F)Coatings

gold, and embrittlement, 133gold, shelf life of, 133, 148(T)non-metallic removal of, 114onto component surfaces, 37solderability shelf life of, 133solderable, 149–152soluble, 149storage shelf life of, 50thickness of, 148thin, by autocatalytic method, 180types of, 149vapor deposition, 234

Cobalt metallization, 71Coefficient of thermal expansion (CTE)

of carbon fibers, 222of components, 26of copper-molybdenum alloys, 161of copper-tungsten alloys, 161of iron-nickel alloys, 160and melting point for metals, 160(F)of metals, and their melting point, 159of molybdenum, 160of non-nickel alloys, 162(F)of Osprey controlled expansion alloys, 163(F)of titanium, 160of tungsten, 160

Coelectroplating, 32Cold compression welding, 200Compatibility, characteristics of, for solder, 49Compliant structures

accommodating thermal expansivity difference,165–167, 166(F)

economics of, 167for mitigation of mismatch expansivity, 166(F)

Component cleaning methods, 172Component surfaces

dissolution of, by brazing, 8dissolution of, by soldering, 8

Component testing, 224Composite materials

controlled expansion materials, 163–164

Index / 247



Composite materials (continued)families of, 163

Composite soldershigh and low-cycle fatigue life of, 221(T)intermetallic compounds in, 219iron containing, 222with large particles, 220preparation methods of, 220of sessile drop tests, 220tensile strength of, 221(T)types of, 220wetting and spreading characteristics of, 220wetting properties of, 221(T)wide joints and, 221yield strength of, 220(F)

Compositional equilibrium, 78Compound formation predictions, 79Compound semiconductors, 105Compression bump bonding, 202(T)Compressive loading

fluxless soldering by, 135(F)and joint shear strength, 136(F)

Concentration of solid metal in liquid metal, 25(F)Conductive adhesives, 3(T)Conductive polymeric materials, 201Conductivity, 30Constraints

imposed on by components and solutions, 168–179joint area, 169–173

Contact angleof Ag-96.5Sn solders, 228(F)of copper-silicon brazes, 16(F)effect of on fillet formation joint filling, 17(F)effective, 22and fillet formation, 17in lap joint, 176(F)on lead-tin solders, 18(F)of lead-tin solders, 22(F)measurement of, 211–212metallurgical modification of, 133of Pb-60Sn solder, 20and quality of wetting, 17rare earth doping of indium solders, 228in reactive wetting, 16and spread factor, 44–45and spread factor and spread ratio, relationship

between, 212(F)and spread ratio, 44–45temperature effects of, 25time dependence of, 15and wetting area, 13–14

Contact angle (of droplets), 45Control charts, 42Controlled expansion materials

alloys, 164(F)components of, 163composite materials, 163–164copper-molybdenum alloys, 160–161copper-surface laminates, 162–163copper-tungsten alloys, 161–162

interlayers, 164–165iron-nickel alloys, 160–161mechanical constraints and solutions, 159–164

Cooling rate, 38Cooling stages, 39, 39(F)Copper

diffusion soldering of, 234effects of, on surface tension of tin, 193lead-free tin-base solders for, 116lead-tin solders for, 116nonmetallic bonding to, 152Pb-63Sn solder wetting using rosin flux, 115(F)rate of dissolution of, in molten Pb-60Sn solder, 84as wettable metallizations, 147wetting of, by Pb-63Sn solder, 115(F)

Copper abiet, 116Copper-base alloys, 10Copper-base brazes, 6Copper coupons, 131(F)Copper-indium alloy systems, 234Copper-indium intermetallic compounds, specific types

Cu2In, for diffusion soldering, 234Copper-Invar-copper laminates, 163Copper-lead-tin phase diagrams, 84–85Copper-lead-tin system

diagram sector of, 86(F)isothermal section of, 86(F)liquidus surface of, 85(F)

Copper metallizations, 132(F)Copper-molybdenum alloys

coefficient of thermal expansion (CTE) of, 161as composite materials, 163controlled expansion materials, 160–161

Copper-molybdenum-copper laminates, 163Copper-nickel-tin phase diagrams, 84Copper particle reinforcement, 220Copper powder, 216Copper-surface laminates, 162–163Copper-tin alloy systems, 234Copper-tin intermetallic compounds

effect of thickness on, 88(F)growth of, 87(F), 88(F)interfacial, rate of formation of, 87from lead-tin solders, 84presence of thick layers of, 88properties of, 87

Copper-tin intermetallic compounds, specific typesCu10Sn, for diffusion soldering, 234Cu3Sn, binary compounds, 87Cu3Sn, precipitates at copper/solder interface, 89Cu6Sn5, as stoichiometric compounds, 91Cu6Sn5, binary compounds, 84, 85

Copper-tin phase diagrams, 85(F)Copper-tin solders, specific types

Cu6Sn reinforced, 220Copper-to-aluminum direct bonding, 162Copper-to-copper joint, 234Copper-tungsten alloys

coefficient of thermal expansion (CTE) of, 161as composite materials, 163




controlled expansion materials, 161–162Copper-zinc compounds, 64Cored silver, 80Corrosion

atmospheric, 37in joints, 49mechanisms of, 29of Sn-9Zn eutectic alloy, 96

Corrosion resistance of lead-free solders, 195Costs. See also economics

of 3.6Ag-1.6Au-92.8Sn eutectic solders, 96of cleaning, 41, 118of forming gas, 111of furnace joining, 104of gold-germanium vs. gold-silicon solders, 70of lead-free solders, 192of low-alpha lead, 190–191of machining molybdenum, 160of machining tungsten, 160of nitrogen, 109of precious metals in metallizations, 231of solder coated substrates, 134of solder elements, 52

Costs and benefits of soldering atmospheres, 118(T)Crack repair, 179Cracked ammonia, 109Cracks

from low-cycle-fatigue, 74orientation of, and x-rays, 236from volume contraction, 67

Creepand heat treating, 158of indium-base solders, 73to relieve mechanical stress, 178resistance of metals to, 195and strain rate measurement, 194thermal cycling without, 165types of, 195

Creep behavior of indium solder alloys, 158Creep curve for lead-tin eutectic solders, 218(F)Creep properties of composite solders, 221(T)Creep rates of various solders, 219(T)Creep resistance

effect of metal additions on, 220of lead-tin eutectic solders, 219of solder, 217

Creep ruptureof indium-base solders, 73and intermetallic thickness, 88(F)

Creep stress, 10Critical angle, 223(T)Critical temperatures, 107Curie point, 160Curing curves, 217(F)

D

De Gennes model, 20, 22Decomposition reaction, 180

Defect-recognition software, 189Defective items, 42Defects. See also tests/testing; voids

distortion, 38, 50, 168–170dross formation, 115, 115(F)inspection of, 235joint embrittlement, 52, 90, 153levels of, and oxide thickness effect of, 124(F)in liquid solder film, 134rates of, in joints, 42and x-ray inspection, 204

Delay effect on joint strength, from cleaning toassembly, 131(F)

Dendritesarm spacing and tensile strength, 32(F)bridging, 120on circuit board, 120(F)growth of, 119–120primary, of silver, 80

Dendritic growth, 120(F)Dental amalgams, 215, 215(F)Depression of melting point, 93–96Dermatitis, 43Design criteria of soldering processes, 28–30Design guidelines for various metals for in-plane

alignment, 205Diamond

as intermediate plate material, 165nonmetallic bonding to, 152

Die attach of gold-metallized chips, 71Differential scanning calorimeter, 155Differential thermal expansion

steps to reduce stress from, 158stress from, 157

Diffuse heating, 33Diffusion

activation energy for, 78rate of, 78surface, 10volume, 10

Diffusion bondingof aluminum-base alloys, 10of copper-base alloys, 10of gold, 4–5, 43of gold, temperature/pressure curve for, 43(F)of indium, 4–5of indium, temperature/pressure curve for, 44(F)interlayers for dissimilar metals, 10limitations of, 230process, 9with standard solders, 43of tantalum, 10of titanium, 10of titanium alloys, 10of tungsten, 10of various metals, 10

Diffusion brazing, 10, 175Diffusion rates

inequality of, 10in solids, 83

Index / 249



Diffusion reaction, 215Diffusion-soldered joints

micrographs of, 233(F)steps involving, 232(F)strength of, 232

Diffusion soldering, 7advantages of, 230for aluminum, 137binary combinations for, 231of copper, 234copper-indium alloy systems for, 234copper-tin alloy systems for, 234Cu10Sn intermetallic phase for, 234described, 230of gold, 233–234interlayers in, 10modeling of, 235practical aspects of, 234–235process principles of, 230–231of silver, 231–232

Diffusion-soldering processeswith Al-12Si alloy, 232basis of, 80for gold jewelry, 233jewelry made by, 234(F)with silver-indium alloys, 232

Diffusion soldering systems, 231(T)Dimensional stability of soldered joints, 224–226DIN WNr 1.3981 (iron-nickel alloys), 160Dip-and-look (DNL) test, 207Dipping methods, 148Direct-bonded copper, 162Direct chip attach (DCA), 199Disc preforms, 170Dislocation climb, 215Dispersion-stabilized alloys, 175Dispersion-strengthened solders, 219(T)Dispersion strengthening, 218, 219Dispersoids, 218, 219, 221Disproportion reaction, 180Dissimilar materials and thermal cycling, joints with,

73Dissimilar-metal joints, 9Dissimilar metals, 10Dissolution

of chromium metallizations to bismuth-containingsolders, 149

of component surfaces, 8of parent materials, 153rate of, 24–25rates of, 153

Dissolution of parent materials and intermetallicgrowth, 24–25

Dissolution rateof copper in molten Pb-60Sn solders, 84, 89of gold in molten lead-tin solder, 89of metals and metallizations in lead-tin solders, 87(F)of platinum, in gold-tin eutectic solders, 71of platinum, in tin-base solders, 148(T)and saturation limit, 25

of silver with lead-tin solders, 53(F)Dissolution rate constant, 24Distortion

of assemblies, causes of, 168–170from bimetallic expansion, 50during heating, 38from thermal expansion mismatch strain, 50

Doping, 227–229Doping additions, 135(F)Drop-in replacement, 191Dross formation

interference of, with wetting and spreading, 115oxygen concentration and rate of, 115(F)

Ductile foils, 155Ductility

improvement of, in filler metals, 155of indium-base solders, 73of indium-bearing solders, 52of silver-tin intermetallic phases, 52of tin-based solders, 52of zinc-based alloys, 64

Duration of diffusion bonding process, 9Dwell stages, 38

E

Economics. See also costsof compliant structures, 167of fluxless soldering, 123of forming gas, 111of joining process, 145of lead-free solders, 192of low oxygen atmospheres, 115of thin foil preforms, 136of vapor-phase techniques, 171

Economics and availability of lead-free solders,191–193

Edge fillets examination, 168Effective contact angle, 22EIA/IS-86 (test method), 207Elastic modulus, 194Electric conductance of mechanical fasteners, 1–2Electrical conductivity

of adhesive joints, 3of composite solders, 221joint requirements, 30

Electrical properties of joints, 50Electrically conductive adhesives, 3Electrode potential of selected elements, 54(F)Electrodeposition, 155Electronic assemblies

cleaning, 40joining methods for, 116

Electroplatingdifficulties with, 234as manner of solder deposition, 31in metallizations, 37tin-lead solder, 32




Ellingham diagramadjustments for plasma state, 127application of, 107–111free energy change for oxidation of several metals,

106(F)for oxides with hydrogen at a partial pressure, 129(F)for selected oxides, 108(F)temperature and O/H ratio for metal oxide reduction,

110(F)Elongation, 194Elongation to failure

of Ag-95Sn solders, 228(F)of Sn-9Zn solders, 228(F)

Embrittlementcaused by impurities, 49of copper-zinc compounds, 64and gold-based coatings, 133liquid metal, 30of solder, 23tin-based solders and AuSn4, 133

Embrittling phases, 4Endoscopes, 237Engineering ceramics thermal expansion, 159Enthalpy of fusion, 97Entropy

changes in, 98–99in chemical bond, 105defined, 137, 138, 139of fusion, 97

Environmental concerns, 191Environmental considerations. See also health and

safety issueschlorofluorocarbons (CFCs), 196volatile organic compounds (VOCs), 115

Environmental durability, 29Environmentally friendly chemicals as cleaning agents,

111Equilibrium constant and oxidizing reaction, 141Equilibrium contact angle, 17–18Equilibrium partial pressure of oxygen. See oxygen

partial pressureEquilibrium phase diagrams, 25Erosion

changes to rate of, in parent materials, 154conditions for, 153of gold, by molten indium, 81of gold by molten tin, 82(F)of parent materials, 49of silver by molten tin, 82(F)of substrate surfaces, 33

Erosion of parent materials, 153–154European Space Agency (ESA) Specifications (void

free joints), 183Eutectic alloys and alloying

benefits of, 81grain refinement of, 81melting point depression by, 93–96melting point of, 81vs. noneutectic alloys, 20spreading characteristics of, 19

spreading of, 81theoretical modeling of, 97–99used for solders and brazes, 6

Eutectic solderscomposition of, 8melting temperatures of, 52

Eutectiferous character of alloys, 51Evolved vapor, 172Exchange reactions, 122Expansion coefficients. See coefficient of thermal

expansion (CTE)Expansion mismatch

compliant structures for mitigation of, 166(F)and cooling stages, 39(F)steps to reduce stress, 158

Expansion on freezing, 53Explosion risk, 110Explosive limit, 111Exposure limits for hazardous dusts and vapors, 43External magnetic fields of iron powder solders, 221Extrapolation of results in spreading test, 212Eye and nose irritation, 43

F

Fast atom bombardment, 128, 180Fast atom cleaning, 131(F)Fatigue cracks

internal initiation of, 217wide-gap joints and, 217

Fatigue failure, 157Fatigue fracture, 165Fatigue life, 200

of carbon-fiber-loaded solders, 223Fatigue resistance

of Alloy J, 61high-cycle, 220indium solder alloys and, 158low-cycle, 220ranks of lead-free solders, 195of solder, 217for soldered joints, 4

Fatigue situations, joint reliability in, 205Fatigue theory, 226Ferromagnetic alloys, 160Filler metal temperatures in brazing and soldering, 5Filler metals

and dissolution with component pieces, 26form of, 31–33with limited component solubility, 25–26molten, flow influences to, 12molten, solidification shrinkage of, 169partitioned, 155(T)partitioning of, 155–157spread characteristics of, 12spreading of, 20surface area to volume, ratio of, 133for welding, 4

Filler spreading characteristics, 19–22

Index / 251



Fillet formation for aesthetic requirements, 50Fillet radius, 176(F)Fillets

beneficial effects of, 167enhancements of, 167with frosty appearance, 81inspection of, 238integrity of, 168quality of, 29to reduce stress concentrations, 176role of, 167–168and stress concentration reduction, 168(F)and stress concentrations, 167as stress reducers, 4

Finishes, 149Finite element analysis (FEA)

of ceramic-metal brazed joint, 225(F), 226(T)modeling assumptions in, 225

Fire risks, 43First Law of Thermodynamics, 137–138Flammability, 43Flip-chip assembly operation, 123Flip-chip bonding, 200(F)Flip-chip bonding process

defined, 201described, 166Pb-5Sn solders for, 166

Flip-chip components pull-in alignment, 132(F)Flip-chip daisy chain, 74(F)Flip-chip inspection methods, 204Flip-chip interconnection

high-temperature, die mounted using, 206(F)self-alignment of, 202(F)self-alignment of solder bumps during, 127thermal expansion mismatch in, 200

Flip-chip interconnects, 179Flip-chip joining process, 167(F)Flip-chip joints with high-lead solders, 129Flip-chip lands, 204(F)Flip-chip process flow for range of solders, in different

atmospheres, 199(T)Flip-chip processes

Au-20Sn eutectic solders for, 207interconnection schemes for, 201(F)lead-tin eutectic solders for, 207step soldered flip-chip interconnects, 206–207surface topology of, 206

Flip-chip solder bumps, 204(F)Flip-chip structures, 204–206Flip-chip technology

characteristics of, 202–203rework, 204underfill, 203

Flow, molten filler metal influences to, 12Fluid flow, 18–19Fluidity

inferior levels of, with bismuth-bearing solders, 53of molten filler metals, 49

Fluorides, 111Fluorine chemistry, 131

Flux action mechanisms, 114Flux activators, 118Flux boil, 123Flux carriers, 117Flux chemistry, 116Flux-cored solder wire, 31, 114Flux-formulations, 114Flux removal, 39Flux residues, 123Flux vapors, 25–26Fluxes

activation temperature of, 117activity of, 116–117for aluminum, 121–122and atmospheres, 8based on organic compounds, 66chemical activity of, 118chemical, described, 111–112chloride-based fluxes for, 121–122classification of, 118(T)commercial designations of, 117containing 2-ethylhexanoate, 195with gold-tin eutectic solders, 122with high-lead solders, 122high-molecular-weight hydrocarbons as, 122high-temperature, 122–123IA type, 117ingredients in, 116for lead-tin eutectic solders, 122for magnesium, 122no-clean, 39, 40, 118–119OA type, 117organic, 121with oxide scale, 34R, RMA, RA type, 117requirements for, 113–114role of, in wetting and spreading, 114SA type, 117for stainless steels, 122that require cleaning, 116–118for tin-base solders, 116–120for “unsolderable materials,” 120–122water soluble, 39, 40WS type, 118for zinc-bearing solders, 65–66

Fluxless joininggold as solderable metallization, 89using Au-20Sn eutectic solders, 71

Fluxless joining process, 30–31Fluxless process, 111Fluxless soldering

of aluminum, 136–137by compressive loading, 135(F)economics of, 123of gallium arsenide (GaAs), 134gold coating for, 124process considerations for, 132–133reactive atmospheres for fluxless soldering, 130using 82Au-18In solders, 72using In-48Sn solder, 135–136




Fluxless wettingon copper metallizations, 132(F)on gold-on nickel metallizations, 132(F)of Pb-62Sn solder, 132(F)

Fluxless wetting angle of indium-silver substrates,135(F)

Foil preformsdeficiencies of, 171production of by solidification casting, 32(F)production of by strip casting, 32(F)and spreading tests, 212

Foils and sheets, 23, 31Force diagram for immersed solid plate, 209(F)Formic acid vapor, 111

concentration, effect of, 132(F)as reactive atmospheres for fluxless soldering, 130–131in self-alignment of flip chip assembly, 130

Forming gas, 111Foundation layers, 150Foundation metal, 199Free energies

comparative values of metal oxide formation, 105(T)of oxide formation, 125

Free-energy changefor oxidation of several metals, 106(T)for oxidation reactions, 106

Free energy of formation, 110Fuel gases, 34(T)Furnace joining, 104Fusible coatings, 149Fusion, 97

G

Galliumand alumina, 62amalgams based on, 215–216liquid alloys based on, 93

Gallium alloy systems, 93Gallium arsenide (GaAs), 134Gallium-base amalgams, 216Gallium-copper amalgams, 216(T)Gallium-indium-tin alloys

Ga-In-Sn, melting points of, 93Gallium-indium-tin solders

Ga-In-Sn, effect of additions to, 94(T)Gallium-nickel-copper amalgams, specific types

Ga-5Ni-30Cu, 216Ga-5Ni-30Cu, curing curves for, 217(F)

Gamma phase intermetallic compound, 215Gap size, 178–179Gaps and x-ray inspection, 204Gas atomization, 220Gas evolution from polymeric materials, 172Gas Law, 139Gas torch, 34Gaseous fluxes

as fluxless process, 111narrow joints with, 178

Gaseous reagents, 140Geometry of adhesive joints, 3Germanium, 77Gibbs free energy

changes in, 105, 141defined, 137depression of, as a function of temperature increases,

98pressure dependence of, 139–141reference point of, 140

Gibbs free energy function (G), 16, 139Gold

as additive to indium, 135characteristics of, 43as constituent of high-melting-point solders, 51critical levels of, 91diffusion bonding of, 4–5, 43diffusion soldering of, 233–234effect of, on tin-base eutectic solders, 153(T)effects of addition to Ag-97.5Pb-1Sn solder, 154(F)effects of addition to In-18Pb-70Sn solder, 154(F)erosion of, by molten indium, 81erosion of, by molten tin, 82(F)high rate of dissolution of, in molten lead-tin solder, 89melting point of, 54in oxidizing atmospheres, 35as solder constituent, 54as solderable metallization, 89temperature/pressure curve for diffusion bonding of,

43(F)as wettable metallizations, 147

Gold-antimony alloys, 66Gold-antimony phase diagrams, 69(F)Gold-base brazes, 66Gold-base solders, 111Gold-bearing solders

for gold-metallized components, 66melting point of, 66used as solders, 67(F)

Gold-coated componentselectrodeposition of, 155solderability shelf life of, 148(T)

Gold coating, 133Gold flash, 10, 147–148Gold-germanium, 66–71Gold-germanium eutectic alloy, 70Gold-germanium phase diagrams, 68(F)Gold-germanium solders, specific types

Au-12Ge, characteristics of, 70Au-12Ge, contact angle of, 72(T)Au-12Ge, iron germides, 70

Gold-indium intermetallic phases, 83(F)Gold-indium joints, 231Gold-indium noneutectic alloys

82Au-18In solders, 72Gold-indium phase diagrams, 52(F)Gold-indium solders, specific types

82Au-18In, fluxless soldering, 7282Au-18In, gold-indium noneutectic alloys, 7282Au-18In, highest melting-point solders, 72

Index / 253



Gold-indium solders, specific types (continued)Au-18In, contact angle of, 72(T)

Gold jewelry, 233Gold layers in metallization, 150Gold-lead-tin ternary system

gold limit in, 89liquidus projection of, 89(F)liquidus surface of, in phase diagrams, 89vertical section through, 90(F)

Gold limit, 89Gold metallizations

characteristics of, 89erosion of, by molten indium, 53(F)and indium-bearing solders, 52–53

Gold-metallized componentsgold-bearing solders for, 66In-48Sn solder to, 136(F)

Gold-on-nickel metallizations, 132(F)Gold-plated tin foil, 156(F)Gold-silicon, 66–71Gold-silicon intermetallics, 66Gold-silicon phase diagrams, 67(F)Gold-silicon solders

Au-2wt%Si, 66cost advantage of gold-germanium eutectic alloy, 69–70as a desiccant for hermetic packaging, 184for silicon semiconductor chips to gold-metallized pads,

66vapor-phase technique, 66

Gold-silicon solders, specific typesAu-2Si, 71Au-2Si, spreading behavior of, 69(F)Au-2Si, titanium as additive to, 69Au-2Si, zinc as impurity of, 77Au-2wt%Si, alloy additions for spreading, 67Au-2wt%Si, gold-silicon solders, 66Au-2wt%Si, high molten viscosity of, 66Au-2wt%Si, silaceous dross, 66Au-2wt%Si, tin as alloy element for, 68

Gold-silicon-tin alloy system phase diagrams, 69Gold-silicon-tin phase diagram, 70(F)Gold-silicon-tin-titanium solders, specific types

Au-2Si-8Sn-1Ti, silicon wetting, 69Gold-tin, 68(F), 71–72Gold-tin alloy system, 233Gold-tin alloys, specific type

Au-20Sn, wetting effect or rare earth doping, 228–229Gold-tin eutectic solders

fluxes with, 122high-melting-point solders, 72intermetallic compounds of, 226

Gold-tin intermetallic compounds, 32AuSn4, and embrittlement, 133AuSn4, and tin-base solders, 133AuSn4, as stoichiometric compounds, 91

Gold-tin intermetallic compounds, specific typesAu-Sn4 phase, and joint embrittlement, 90Au-Sn4 phase, in Au-Pb-Sn ternary system, 89

Gold-tin partitioned filler metals, 155Gold-tin phases and joint embrittlement, 90

Gold-tin solder, 32Gold-tin solders, specific types

Au-20Sn, application of, 71Au-20Sn, bismuth as impurity of, 77Au-20Sn, cobalt metallization for, 71Au-20Sn, fabrication costs of, 155Au-20Sn, fluxless joining of, 71Au-20Sn, foils and preforms of, 71Au-20Sn, for flip-chip process, 207Au-20Sn, for soldering to gallium arsenide, 153Au-20Sn, forms of, 71Au-20Sn, hermetic sealing of ceramic semiconductor

packages, 71Au-20Sn, high-lead solders as alternate to, 72Au-20Sn, intermetallic compounds of, 226Au-20Sn, melted in controlled atmosphere, 113(F)Au-20Sn, palladium metallization for, 71Au-20Sn, rapid solidification casting, 71Au-20Sn, titanium as additive to, 69Au-30Sn, foil melting point, 156(F)

Graduated joint structures, 165Grain-boundary sliding, 215Grain refinement, 81, 218Graphite brazing, 30Green manufacturing, 190

H

Halide atmospheres, 111Halide fluxes, 117Hallmarking regulations, 50Halogen-base fluxes, 122Halogen gases, 130Health and safety issues

asthma, 43from beryllia dust, 163of beryllium, 216of cadmium and cadmium alloys, 50, 51, 63, 191of electronics equipment disposal, 190explosion risk of hydrogen atmospheres, 110exposure limits for hazardous dusts and vapors, 43eye and nose irritation, 43flammability, 43of hydrogen, 37of lead, 50of mercury, 93, 191, 215of nickel, 50of organic metal compounds toxicity, 131soldering fumes, 43of thallium, 191

Heat-affected zone (HAZ), 4Heat capacity, 62Heat treatment

and creep, 158in non-oxidizing atmosphere, 148vs. oxygen concentration, 214(F)prior to joining, 37–38stress relaxation, 158temperature of, 39




Heat treatment temperature, 213Heating cycles

of joining operations, 38–39parameters of, 38profiles of, 38(F)

Heating methods, 33–34Heating rate, 38Heavy metal deposition, 122Helium, 184Helium leak rate, 184Hermetic sealing

Bi-43Sn solders for, 173of ceramic semiconductor packages, 71with compressive force, 174dryness of, 184–185In-48Sn solder, 44Sn-40Pb solder, 44with standard solders, 43

High creep levels of high-lead solders, 72High-cycle fatigue life of composite solders, 221(T)High-cycle fatigue resistance, 220High-lead solder alternative, 197High-lead solders

98Pb-2Sn solders, 73as alternate to Au-20Sn solders, 72fluxes with, 122high creep levels of, 72mechanical properties and corrosion resistance of,

73(T)mechanical properties of, 73step-soldering sequence for, 72

High-melting-point solders, 54constituents of, 51eutectic alloys of, 7(F)fast atom cleaning effect of, 131(F)lead-free, 197–198lead rich, 72(T)for thick film metallizations, 72

High-molecular-weight hydrocarbons, 122High-purity solders, 77High-volume contraction, 64Highest melting-point solders, 72Highly active elements, 103Holding time, 38Homologous temperatures, 6Hooke’s law, 27–28Hot shortness, 26Hydrochloric acid (HCL), 116Hydrogen

explosion risks of, 37in inert atmospheres, 36solder oxides reduction by, 126–127uses of as reducing gases, 109

Hydrogen atmospheresand explosion risk, 110and gold-base solders, 111oxides with hydrogen at a partial pressure, 129(F)tin oxide reduction in, 111

Hydrogen plasma, 128Hydrogen poisoning, 105

Hydrogen safety, 127Hydrostatic forces, 15, 27, 212Hydrostatic pressure, 178–179Hypereutectic alloys, 19Hypoeutectic alloys, 19

I

IA type fluxes, 117Ideal substrate, 19Image analysis, 211Immersion plating, 180Impurity

aluminum as, 76, 77antimony as, 76bismuth as, 76, 77cobalt and iron as, 77germanium as, 77silver as, 77tin as, 77zinc as, 77

Indentation welding, 9Indicative property values of selected solders and pure

metals, 226(T)Indium

amalgams based on, 217diffusion bonding of, 4–5fluidity of, 135melting point of, 75(F)molten, and gold erosion, 81in pressure-welded joints, 43production levels of, 192pure, as a solder, 74temperature/pressure curve for diffusion

bonding of, 44(F)and tin in intermetallic compounds, 51

Indium amalgams, 217Indium and indium alloys, 201Indium and lead, 52–53Indium-base solders

alloy additions to, 74benefits of platinum in, 147–148composites and melting points of, 75(T)creep rupture of, 73ductility of, 73intermediate melting temperature solders, 75low melting point of, 73, 74in optoelectronic applications, 74phase segregation failure of, 74in photonic applications, 74rare earth doping and contact angles of, 228superheats for, 75

Indium-bearing soldersductility of, 52and gold-metallizations, 52–53

Indium bump bonding, 200Indium-gold reaction, 82Indium-lead alloys

non-eutectiferous character of, 51

Index / 255



Indium-lead alloys (continued)peritectic reaction, 75

Indium-lead phase diagrams, 61(F)Indium-lead-silver solders, specific types

In-15Pb-5Ag, 172(F)In-15Pb-5Ag, indium-base solders, 75In-15Pb-5Ag, intermediate melting temperature solders,

75In-15Pb-5Ag, thermal fatigue performance of, 75

Indium-lead-tin solders, specific typesIn-18Pb-70Sn, gold addition, effect of, 154(F)In-18Pb-70Sn, gold addition effects on, 154

Indium oxidesalloying additions to facilitate removal of, 74–75stability of, 125temperature for reduction of, 126

Indium-silver substrates, 135(F)Indium solder alloys

creep behavior of, 158fatigue resistance and, 158

Indium solders, 73–75, 116Indium-tin phase diagrams, 55(F)Indium-tin solders

for magnesium, 122Indium-tin solders, specific types

In-48Sn binary, melting point of, 96In-48Sn, contact angle for, 128(F)In-48Sn, continuum between stress strain and creep

data for, 74(F)In-48Sn, elongation, 194In-48Sn, fluxless soldering, 135–136In-48Sn, hermetically sealed enclosures-hermetically

sealed enclosures, 44In-48Sn, joining operations with, 135In-48Sn, oxide growth on, 125In-48Sn, preform of, 136(F)In-48Sn, stress-strain curve for, 74(F)In-48Sn, to gold-metallized components, 136(F)

Indium-tin-zinc solders, specific types5In-87Sn-8Zn, grain refinement in, 218In-46Sn-2Zn ternary, melting point of, 96In-86Sn-9Zn, melting point of, 96

Inert gas atmospheresdescribed, 35–36industrial quality, 107soldering in, 107–109types of gases in, 36uses of, 109

Infrared microscopy, 204Inorganic acid fluxes, 117Inorganic acids, 116Inorganic fluxes, 53Inspections of joint interior, 183Instantaneous melting properties, 93Interatomic force

per unit area, 27variations of, 27(F)

Interdiffusion, 148–149Interfacial compound formation, 84Interfacial compounds, 26, 103

Interfacial reactions, 19, 24Intergranular brittle fracture, 215Intergranular corrosion, 114Interlayers

controlled expansion materials, 164–165for diffusion bond dissimilar metals, 10gold flash, 10nickel foil, 10silver foil, 10

Intermetallic compoundsbarrier metallizations to avoid, 155in composite solders, 219duration of growth, 25effects of, 153elastic modulus of, 154formation of, 10, 51, 154gamma phase, 215growth of, 16, 25indium and tin in, 51mechanical and physical properties of, 87(T)in solders, 5in ternary systems, 84thickness of, and creep rupture, 88, 88(F)

Intermetallic phase layer, 16Internal energy, 137–138International Electrotechnical Commission (IEC),

119(F)International Organization for Standards

flux classifications of, 117soldering flux classifications, 118(T)

International Programme for Alloy Phase DiagramData (IPADA), 79

Invar, 160Invariant reactions, 193Ion-aided deposition process, 180Ion-assisted vapor spray deposition, 178Ion plating, 180Ionic contamination, 119Ionized-cluster beam deposition, 180IPC/EIA J-STD-003A (test method), 207Iron and tin dendrites, 77Iron germides, 70Iron-nickel alloys

controlled expansion materials, 160–161machinability of, 161trade names of, 160

Iron-nickel alloys, specific type17Co-54Fe-29Ni, coefficient of thermal expansion

(CTE), 16017Co-54Fe-29Ni, iron-nickel alloys, 16017Co-54Fe-29Ni, Kovar, 160Fe-36Ni, coefficient of thermal expansion (CTE), 161Fe-36Ni, thermal expansion characteristics of, 162(F)Fe-36Ni, total expansion of, 162(F)Fe-42Ni, coefficient of thermal expansion (CTE), 161

Iron-tin intermetallic compound, specific typesFeSn2, as stoichiometric compounds, 91

Irreversible processes, 139




J

Jetting system schematic, 201(F)Jewelry, 175, 233Jigging, 30–31Joining atmospheres, 35, 103–111Joining environments, 103–141Joining methods, 1–12, 2(F)Joining operations, heating cycle of, 38–39Joining process

cost tolerance of, 145fluxless, 30–31

Joining process development, 146(T)Joining temperature, minimum practical (liquidus),

145Joint defect rates, 42Joint design

dimensions, 12dimensions and mechanical properties, 25issues regarding, 30to minimizes concentration of stresses, 175–178remedy of problems in, 146strengthened solders, 178trials of, 145

Joint embrittlementeffects of, 153effects of intermetallic phases on, 153and gold-tin phases, 90of indium-bearing solders, 52of tin-based solders, 52

Joint filling requirements, 26Joint gaps

compressive forces applied to, 134control of, 26limits to, 25–26optimal balance of, 26self regulation of, 27significance of, 25–27size of, 26upper practical limit to, 26voids in, 169(F), 170(F)

Joint geometries, 29for brazed joints, 4length vs. void content, 169(F)of mechanical fasteners, 1for soldered joints, 4for welding, 4

Joint integrityadvances in techniques for assessing, 235–238optical inspection, 237–238scanning acoustic microscopy, 235–236ultrasonic inspection, 235–236x-radiography, 236–237

Joint quality degradation and power cycling, 65(F)Joint strength

of Ag-96.5Sn eutectic solders, 232of Ag3Sn intermetallic phase, 232effect on, from delay from cleaning to assembly, 131(F)and wetting of component surface, 136

Joint weakness, 28

Joints, 175–178butt joints, 177, 177(F)cleaning for solid-state joining, 5cleanliness of, 4corrosion in, 49dimensional stability of, 224–226with dissimilar materials and thermal cycling, 73fitness for purpose tests, 224landed butt, 177lap joints, 175, 176(F), 176–177, 177(F), 229large area, 168, 171(F), 174(F)lifetime prediction of, 226–227measurement of mechanical properties of, 223–224mechanical properties of, 26, 224modeling lifetime of, 223narrow, 178numerical modeling of, 224–227peel force profiles of, 168(F)quality of, 12recommended designs for, 177(F)scarf butt joints, 177shear strength and compressive loading, 136(F)shear strength as a function of thickness, 178(F)shear strength of, 136(F), 178(F), 222solidification shrinkage, 173step butt joints, 177strap joints, 177, 177(F)strength factors influencing, 224surface roughness of, 22tensile strength of, 222tongue and groove joints, 177trapped gas, 169–173trapped gas sweeping, 170(F)voids in, 49wide gaps in, 158

Journal of Phase Equilibria (phase diagrams), 79

K

Kelvin-Planck statement, 138Kinetics of reaction, 153Kirkendall porosity, 10Kirkendall voids, 71, 234Kovar, 160

L

Landed butt joints, 177Lanthanum, 227Lap joints, 176–177

failure in, 176(F)geometry of, 176(F)recommended designs for, 177(F)shear stress on, 175strength of, and rare earth doping, 229stress distribution in, 176(F)

Large-area jointsdefinition of, 168

Index / 257



Large-area joints (continued)due to trapped gas, 174(F)pressure variation and void levels, 174(F)vapor pockets in, 171(F)voiding in, 171(F)

Latent heat of fusion, 97Lattice waves, 226Lead

as constituent of high-melting-point solders, 51effects of, on surface tension of tin, 193health restrictions on, 50in landfills, 190pure, as solder, 72–73

Lead-antimony-tin system, 62(F)Lead-antimony-tin ternary system, 57Lead-free solder paste, 155Lead-free solders

Ag-20In-77Sn alloy for, 96Al-3Ga-3Mg-90Zn solders, 66Alloy J, 60availability of potential alloying elements for, 192(T)composition ranges for, 192(T)corrosion resistance of, 195costs of, 192drive for, 190–191economics and availability of, 191–193economics of, 192elongation of, 194fatigue resistance ranking of, 195and fluxes, 196health, safety and environmental aspects of, 191high-melting point, 197–198history of, 190In-48Sn binary eutectic solders, 96In-86Sn-9Zn solders, 96lead-tin solder, alternatives to, 191–193lead-tin solder, compatibility with, 191literature on, 189–190mechanical properties of, 194–195melting point depression, implications for, 95–96melting ranges of, 191, 196metallurgical, physical, and chemical properties of,

193–196other physical properties of, 194physical and chemical characteristics of, 191plastic flow of, 194for plumbing applications, 51for printed circuit board (PCB) components, 95process window for, 196–197reduction in superheat for, 196shear strength of, 194silver-copper-tin ternary phase equilibria, 193strain rate of, 194surface tension of, 193–194tin-based solders as key to replacement for, 192tin pest and tin whiskers, 195–196ultimate tensile strength of, 194wetting and spreading characteristics of, 197yield strength of, 194

Lead-free tin-base solders, 116

Lead-rich alloys, 6Lead-silver solders, specific types

1.5Ag-92.5Pb-5Sn, micrograph of, 73(F)Lead-tin alloys

as compression interconnects, 201tensile strength of, 81(F)

Lead-tin-copper solders, specific types61.75Sn-38.05Pb-0.2Cu(wt%), lead-tin eutectic solders,

84Lead-tin intermetallic compounds, specific types

Pd3Sn2, with palladium metallization, 71Lead-tin phase diagrams, 58(F)Lead-tin solders

additions to, 57advantages of, 56antimony in, 57application of hydrogen plasma to, 128balls of, 129(F)bismuth in, 57change of dissolution rate of silver, 154characteristics of, 191contact angle of, 22(F)for copper, 116and copper substrate, reaction phases formed by, 87(F)copper-tin intermetallics from, 84creep curve for, 218(F)creep-resistance of, 219dissolution rate of metals and metallizations in, 87(F)dissolution rate of silver in, 53(F)drop-in replacement for, 95, 96, 191for flip-chip process, 207fluxes for, 122grain refinement, 218history of, 56lead-free solders, alternatives to, 191–193lead-free solders, compatibility with, 191molten, high rate of dissolution of gold in, 89repairs to, 191shear strength of joints with, 54(F)silver in, 57vs. silver-tin solders, 60solder alloy systems, 56–60tensile strength of, 222(F)wetting angle of, 14(F)wetting behavior of, on mild steel, 210(F)wetting speed of, 197(F)

Lead-tin solders, specific types98Pb-2Sn, Chadwick peel tests of, 73Pb-3Sn, effect on joint strength on delay from cleaning

to assembly delay, 131(F)Pb-3Sn, fast atom cleaning effect of, 131(F)Pb-3Sn, flip-chip joints, 129Pb-3Sn, high melting point, 131(F)Pb-4Sn, 3095Pb-5Sn, oxide growth on, 125Pb-5Sn, for flip-chip bonding process, 166Pb-26Sn, eutectic composition, 89Pb-60Sn, 51Pb-60Sn, contact angle of, 20Pb-60Sn, effect of major ternary additions to, 76




Pb-60Sn, impurity concentrations producing detrimentaleffects for, 76(T)

Pb-60Sn, lead-tin eutectic solder, 51Pb-60Sn, molten, rate of dissolution of copper in, 84Pb-60Sn, spreading of, on gold-plated sample, 213(F)Pb-62Sn, elongation, 194Pb-62Sn, fluxless wetting of, 132(F)Pb-62Sn, for copper components, 131Pb-62Sn, for gold-nickel components, 131Pb-63Sn, copper wetting using rosin flux, 115(F)Pb-80Sn, effect of alloying additions on wetting of,

76(T)Pb-80Sn, effect of major ternary additions to, 76

Leadless ceramic chip carriers (LCCs), 40Lever rule, 80, 80(F), 92–93Lifetime prediction of joints, 226–227Limitations of solderability test measurements, 210Linear expansion coefficient vs. thermal conductivity,

159(F)Liquid flow, 19Liquid infiltration casting, 161Liquid lake condition, 134Liquid metal embrittlement, 30Liquid nitrogen, 109Liquid phase sintering, 157Liquid solder film, 134Liquid-solid metallurgical reactions, 34Liquid spreading, 19Liquidus projection, 89(F)Liquidus surface

of Ag-Cu-Sn system, 63(F)of Ag-Sb-Sn system, 64(F)of Cu-Pb-Sn system, 85(F)of Pb-Sb-Sn system, 62(F)of Si-Pb-Sn system, 65(F)

Liquidus temperature, 49of 3.8Ag-0.7Cu-95.5Cu solders, 228(F)calculating depression of, 98of eutectic silver solders, 80increases of, from silver additions, 53

Load applied to preforms vs. void level, 134(F)Local heating, 33Local interfacial mismatch stresses, 28Low-alpha-emission solders, 190Low-alpha lead, 190–191Low-cycle fatigue life, 221(T)Low-cycle fatigue resistance, 220Low-expansion materials, 161(T)Low-expansion metals, 160Low-expansivity materials, 223Low-melting-point eutectic alloys, 55(F), 149Low-melting-point metals, 192(T)Low-melting-point solders

Ag-96Sn properties of, 53eutectic alloys of, 7(F)of indium-base solders, 73stress-rupture life of, 195(F)

Low-solid fluxes, 118Low spreading and bond quality, 212Lower-melting-point solders, 51–52

Lutetium, 152Lutetium oxide (Lu2O3), 152

M

Machinability and machining costs, 160, 161Magnesium, 122Magnetostriction, 160Maps of brazes and solders, 7Materials

storage of, 43strength of, 27–28

Materials systems approach, 146(T)Maximum exposure limits, 43Mechanical cleaning, 37, 209Mechanical constraints and solutions, 157–168Mechanical fasteners and fastening, 1–2Mechanical integrity, 29Mechanical properties. See also tests/testing

of intermetallic compounds, 87(T)of joints, 26of joints, measurements of, 223–224of lead-free solders, 194–195of selected solders, 194(T)of soldered joints, 224tests/testing of, 224

Mechanical strengths in solders, factors in, 223–224Melting, instantaneous, 93Melting point

of Ag-20In-77Sn solders, 96of Ag-Cu-Sn ternary system, 96of aluminum-silicon alloys, 6effect of, on multiple alloying additions, 97–98of eutectic alloys, 81of eutectic solders, 52gold addition effects on, 153–154of gold-bearing solders, 66of In-46Sn-2Zn ternary eutectic, 96of In-48Sn binary eutectic, 96of In-86Sn-9Zn solders, 96of indium, 75(F)of liquid metal, 15of lower-melting point solders, 51of metals and their coefficient of thermal expansion

(CTE) of, 159, 160(F)and peak operating temperature, 29of silver-tin binary system, 96of solid metal, 15

Melting point depressionbehavior of, 93–95cadmium-base solders, 93by eutectic alloying, 93–96general features of, 93–95lead free solders, implications for, 95–96liquid alloys based on gallium, 93

Melting rangesof Alloy J, 60and braze alloy families, 7(F)of lead-free solders, 191

Index / 259



Melting ranges (continued)and solder alloy families, 6(F)

Mercuryamalgams based on, 215–216health hazards, 215health safety issues of, 93melting point of, 93and silver powder, diffusion reaction of, 215substitutes for, 93toxicity of, 191

Mercury-base amalgams, 215Metal-ceramic composites, 163Metal-loaded glass frits, 181Metal-matrix composite (MMC), 175Metal-metalloid composites, 163Metal oxide reduction, 110(F)Metal oxides

bond strength of, with parent metals, 106–107chemical reduction of, 105and formic acid, reaction between, 131

Metal-to-oxygen chemical bond, 105Metallic impurities, 75–77Metallization layer, 151Metallization techniques

advantages and disadvantages of, 181as a barrier, 153characteristics of, 181(T)chemical vapor deposition, 180coating quality of, 182(T)of porous ceramic materials, 151(F)process parameters of, 182(T)relative merits of, 182(T)with silver electroplating, 151(F)thick-film formulations, 180–181wet plating, 180

Metallizationsof alumina, 151and control of spreading, 148costs of, 152, 231firing on glass and ceramics, 151gold layers in, 150in layers, 150and metals in diffusion bonding, 11(T)moly-manganese process of, 151noble metals, 181of oxide ceramics, 151of parent materials, 49for refractory materials, 130solderable, gold as, 89stresses in, 151with titanium, 150widely used, 37with zirconium, 150

Metalloid-metalloid composites, 163Metallurgical considerations for solder selection, 78Metallurgical constraints and solutions, 147–157Metallurgical incompatibility of materials and

processing conditions, 147Metallurgical modification of contact angle, 133Metallurgical properties of lead-free solders, 193–196

Metallurgical reactions, 12erosion rate changes to parent materials, 154liquid-solid, 34melting point changes, 153–154

Metallurgical stability of soldering processes, 29Metals

cohesive strength of, 27and their properties, 161(T)wetting of by solders, 147–157

Metals and metallizations in diffusion bonding, 11(T)Microcracks, 28Microelectromechanical systems (MEMS), 123, 222Microflame torch, 34, 34(F)Microfocus x-ray systems, 236(F), 236–237Micrographs

of Ag3Sn intermetallic phase, 92(F)of controlled expansion (CE) alloys, 164(F)of diffusion-soldered joint, 233(F)of lead-tin solder on gold-plated copper substrate,

213(F)of peritectic transformation, 83(F)

Microstructural coring, 75Military Standards (MIL-STD)

883D, (void free joints), 183883E, method 2019.7 (acceptance criteria), 231

Mischmetals, 227Mismatch stress concentration, 164Mismatch stresses, 39

elimination of, 165–166Modeling

joint lifetime, 223, 224of lifetime of joints, 224

Modeling assumptions in finite element analysis (FEA),225

Modeling elements, 226Modulus of elasticity, 157Moisture content

as a function of bakeout time and temperature, 185(F)as a function of bakeout times, 184

Moisture in semiconductor failure mechanisms, 123Moisture permeation, time predicted for, 184(F)Moisture removal, 184Molar volumes, 140Molecular hydrogen, 127Molten acetamid fluxes, 122Moly-manganese process, 151Molybdenum

of coefficient of thermal expansion (CTE), 160joining atmospheres for, 104as low-expansion metals, 160

Monatomic hydrogen, 127–128Monolithic plates, 165(F)Montreal Protocol on Substances That Deplete the

Ozone Layer: 1991, 111Multichip modules (MCMs), 199Multilayer metallic coatings, 50Multilayer metallization, 199Multiphase materials, 160Multiple frequency SAM, 236




N

Narrow joints, 178Native oxides, 126(T)Neodymium, 227Neutral flame, 34Nichrome, 150Nickel

health restrictions on, 50toxicity of, 191

Nickel foil interlayers, 10Nickel powder, 216Nickel-tin intermetallic compounds, specific types

Ni3Sn2 phase, binary compounds, 84Ni3Sn4 phase, binary compounds, 84

Nilo-Alloy 42, 160Nilo-K, 160Niobium (columbium), 104Nitrogen

costs of, 109dry, for moisture removal, 184in inert atmospheres, 36

Nitrogen-based soldering, 197No-clean fluxes, 39, 40, 118–119Noble-metal metallizations, 181Noble metals

embrittlement from, 147oxides of, 105stability of, 105

Non-nickel alloys, 162(F)Nonmetallic bonding, 152Nonmetallic components

chemical vapor depositions on, 149–150physical vapor depositions on, 149–150wet plating on, 149–150

Nonmetallic materials, wetting and spreading with,103

Nonmetallic phases and wetting problems, 147Nonmetallic surface coating, 114Nonmetals

foundation layers for, 150solderable coatings on, 149–152wetting of by solders, 149–153

Nonoxidizable metallizations, 133Nonwettable patches, 17Numerical modeling, 223–224

O

OA type fluxes, 117Off-eutectic compositions, 54Optoelectronic devices, 214–215, 221Organic acids, 116Organic acids fluxes, 117Organic coatings, 149Organic films, 37Organic fluxes, 121Organic metal compounds, 131Osprey spray-forming process, 163

Outgassing, 30Oxidation

rate of, 107water vapor as source of, 36

Oxidation reactions, 106Oxide-dispersion-strengthened solders, 218–219Oxide films

destabilization of, and rare earth doping, 228effects of, 111reduction of, 105–106regrowth of, 129removal of, 35

Oxide formation and removal, 124–125Oxide growth

on 95Pb-5Sn solders, 125on base metals, temperature dependence, 125(F)on base metals, time dependence, 124(F)equation for, 124on molten solders, 126(F)

Oxide layers, 122Oxide reduction

alternative atmospheres for, 111dynamics of, 126thermodynamic aspects of, 106–107thermodynamic principles for analysis of, 106

Oxide reduction rate by hydrogen, 126Oxide scale, fluxes with, 34Oxide thickness

and defect levels, 124effects of, on defect levels, 124(F)vs. oxidation time, 125(F)temperature effects on growth of, 124

Oxideson aluminum, 9growth of, 125, 148with hydrogen at a partial pressure, 129(F)mechanical removal of, 128–130native, superheats to dissolve, 126(T)reduction of, by reactive gas atmosphere, 130–131surface, removal of, 37

Oxidizing atmospheresgold in, 35graphite effect in, 30platinum-group metals in, 35

Oxidizing flame, 34Oxygen atmospheres and solder spreading, 127(T)Oxygen concentration

effect of, on rate of dross formation, 115(F)vs. heat treatment, 214(F)

Oxygen partial pressure, 35–36to effect oxidation reaction, 141and metal oxide bond strength, 106method for reduction of, 107and water vapor desorption, 108

Oxygen, residual levels of, 35Ozone depletion, 39

Index / 261



P

P-charts, 42Package leak rates, 184Palladium metallization, 71Palladium oxide, 107Palladium plated devices, 207–208Parent materials

erosion of, 49, 153–154metallization of, 49wettability of, 49

Partial pressure. See also oxygen partial pressure, 36Partitioned filler metals, 155(T), 155–157Pastes, 31Path length, 169–170PCBs. See printed circuit boards (PCBs)Peak bonding temperature, 38Peak operating temperature, 29Peak process temperature, 33Peel force, 167(F)Peel force profiles, 168(F)Peel fracture and height of the solder film, 167(F)Peeling stresses, 175Peritectic reaction

of aluminum quaternary alloy, 83described, 83indium-lead alloys, 75

Peritectic transformation, 83(F)Phase diagrams

availability of, in literature, 79binary alloy systems, 79–83and binary alloys, 79binary eutectic composition solder with

intermetallics, 81–83binary eutectic composition solder with no

intermetallics, 79–81binary peritectic solder, 83described, 78distributed compound between eutectic solder and

component metals, 89–92higher order systems, 92–93interfacial compound between eutectic solder and

component metals, 84–89limitations of, 79non-metallic systems, 92–93overview of, 49soldering applications of, 77–79ternary alloy systems, 83–92of ternary systems, 84uses of, 79in weight percentages, 78

Phase diagrams, specific alloy systemsof aluminum-zinc, 66(F)of antimony-tin, 60(F)of bismuth-tin, 56(F)of copper-tin, 85(F)of gold-antimony, 69(F)of gold-germanium, 68(F)of gold-indium, 52(F)of gold-lead-tin, 89

of gold-silicon, 67(F)of gold-silicon-tin, 69of gold-tin, 68(F)of indium-lead, 61(F)of indium-tin, 55(F)of lead-tin, 58(F)of silver-indium, 57(F)of silver-lead, 61(F)of silver-tin, 59(F)

Phase formation, 154–155Phase segregation failure, 74Phased reflow soldering, 155, 157Phases

instability of, 29rate of growth of, 81

Phonons, 226Phosphoric acid fluxes, 122Physical abrasion for surface cleaning, 129Physical properties

of Ag-96Sn solders, 53of copper-tin intermetallic compounds, 87of intermetallic compounds, 87(T)of lead-free solders, 193–196for low-expansion materials, 161(T)of selected solders, 194(T)for semiconductors, 161(T)

Physical vapor depositions, 149–150Piezoelectric ceramic elements, 31(F)Planar joints, 50Plasma, 127Plasma-assisted dry soldering (PADS), 131Plastic flow, 194, 219Plastic leaded chip carriers (PLCCs), 40Plastic yielding, 10Platinum, 147–148, 148(T)Platinum-group metals, 35Platinum-tin intermetallic compounds, specific types

PtSn4 interfacial layer, 148Polymeric materials, 172Polymers, 3Porosity. See voidsPorous ceramic materials, 151(F)Powder metallurgy materials (P/M), 161Powdered solders, 31Power cycling and joint quality degradation, 65(F)Precious metals

binary combinations for diffusion soldering, 231hallmarking regulations for, 50

Precipitation hardening, 49Precipitation-strengthened alloys, 175Preform geometry, 133–134Preforms

Alloy J for, 61disc shaped, 170dual disc, 171(F)foil, deficiencies of, 171hermetically sealed enclosures-hermetically sealed

enclosures, 44round wire, cross shaped, 171thickness of, 133




Pressure dependence of Gibbs free energy, 140Pressure in diffusion brazing, 10Pressure variation method

to minimize voids, 174for reduction of level of voids, 173for void level reduction, 174(F)

Pressure welding, 4–5of aluminum, 9and dissimilar-metal joints, 9indium in, 43joint strength of, 9(F)with standard solders, 43

Pressure welding and diffusion bonding, 8–12Primary dendrites of silver, 80Principle of Conservation of Energy, 137Printed circuit boards (PCBs)

assembly line change to lead-free solders, 197lead-free solders for, 95reflow soldering of, 36solders for low process temperature, 155, 157tin whiskers on, 196

Process control, 42Process control chart of peak reflow temperature,

42(F)Process cycle time, 33Process variability (r-chart), 42Process window for lead-free solders, 196–197Processing aspects, 30–42Product miniaturization, 151Profilometer traces, 211Progressive alloying, 98Progressive eutectic alloying, 98Progressive melting, 157Propane, 34Pulse-echo ultrasonic inspection, 235Pure metals, 14Pyroelectric elements, 34Pyrometers, 35Pythagorean theorem, 45

Q

Quality acceptance criteria, 209Quality-control testing, 210Quality of soldered joints, 12Quality of wetting and contact angle, 17Quasi-binary alloy systems, 92Quasi-ternary alloy systems, 92Quenching stages, 39

R

R-chart (range chart), 42R, RMA, RA type fluxes, 117Radiographs

of Ag-96Sn foil solders, 23(F)of joint integrity, 236–237of silicon chips, 169(F)

Range chart (r-chart), 42Raoult’s Law, 97, 98Rapid-solidification process, 66Rapidly solidified alloys, 31Rapidly solidified filler metals, 60–61Rapidly solidified processes, 31Rare earth doping

contact angle of indium solders, 228effects of, on solders, 227oxide film destabilization, 228of Sn-9Zn solders, 228

Rare earth elementseffect of additions on solder properties, 227–229function of, in nonmetallic bonding, 152implications of, for soldering technology, 229solders doped with, 227–229

Rate of dissolution. See dissolution rateRate of reaction, 78Reactive filler alloys, 152Reactive filler metals, 152Reactive gas atmospheres, 130–131Reactive ion etching, 172Reactive-metal metallizations, 181Reactive metals, 150Reactive solders, 103Reactive wetting, 16Reduced-oxide soldering activation (ROSA) process,

132Reducing atmospheres, 8, 36–37, 109–111Reducing gases, 109Reduction, 130(F)Reduction atmospheres, 34Reduction flame, 34Reference standards, 213–214Reflow soldering, 33–34, 36Reflow stage, 38Refractory metals

list of, 104oxides of, 105stability of, 105wetting problems with, 147

Reinforced solders (solder composites), 222–223Reinforcing plates, 165Relative spreading, 211Resin materials, 215Respiratory problems, 43Restoring force, 205Reverse-gas bias mode, 150Reversible chemical reactions, 105Reversible processes, 138, 139Rheological concepts, 215Richardson-Jeffes diagram. See Ellingham diagramRoll-bonding, 9Rosin, 116Rosin fluxes, 115(F), 117Rule of mixtures, 194

Index / 263



S

S-charts (standard deviation chart), 42SA type fluxes, 117Sacrificial metal, 200Safety. See health and safety issuesScanning acoustic microscopy, 170(F)Scanning acoustic microscopy (SAM), 189, 235–236Scanning acoustic techniques, 183Scarf angle, 177Scarf butt joints, 177Screen printing, 31Second Law of Thermodynamics, 138Self-alignment

of flip chip interconnection, 202(F)predictions of, 206of solder bump bonding, 201

Semiconductor die attach, 69Semiconductors, physical properties for, 161(T)Service requirements of joints, 50Service temperature range, 3Service temperatures

of brazed and soldered joints, 4of brazes and solders, 8of solid-state joining, 5

Sessile drop tests, 211(F)of composite solders, 220tests/testing, 209

Shear strengthof fluxless joints, 37(F)as a function of joint thickness, 178(F)of gold-indium joints, 231of joints, 222of joints with lead-tin eutectic solders, 54(F)of lead-free solders, 194of soldered joints, 178(F)

Shear stress, 175Shelf life and storage requirements of fluxed

components, 114Shelf life of multilayer metallic coatings, 50Silaceous dross, 66Silica, 69Silicon, 152Silicon-aluminum alloys

coefficient of thermal expansion (CTE), 163Osprey spray-forming process, 163

Silicon carbide, 152Silicon chip radiograph, 169(F)Silicon-lead-tin system, 65(F)Silicon semiconductor chips to gold-metallized pads,

66Silver

coring of, 80diffusion soldering of, 231–232dissolution rate of, in lead-tin eutectic solders, 53(F)effects of, on surface tension of tin, 193erosion of, by molten tin, 82(F)as impurity, 77oxides of, 38primary dendrites of, 80

as solder constituent, 53sulfides of, 38as wettable metallizations, 147

Silver additions, 53Silver-antimony-tin solders, specific types

Alloy J, 60Silver-antimony-tin system, 64(F)Silver-base brazes, 6Silver-base filler metals, 57Silver-coated components, 81Silver-copper-phosphorous brazes, 5Silver-copper-tin alloys

lead-free solders, 193superheats of, 126

Silver-copper-tin lead-free solders, 115Silver-copper-tin phase diagrams, 193Silver-copper-tin solders, specific types

3.5Ag-0.9Cu-95.6Sn ternary, cooling rate dependencyof, 193

3.5Ag-0.9Cu-95.6Sn ternary, melting point of, 963.8Ag-0.7-95.5Sn, stress-rupture life of, 228(F)3.8Ag-0.7Cu-95.5Cu, liquidus temperatures of, 228(F)3.8Ag-0.7Cu-95.5Cu, solidus temperatures of, 228(F)Ag-1.7Cu-93.6Sn, cobalt and iron as impurity of, 77

Silver-copper-tin system, 63(F)Silver-copper-tin ternary system, 96Silver electroplating, 151(F)Silver foil interlayers, 10Silver-gold-tin solders, specific types

3.6Ag-1.6Au-92.8Sn, costs of, 96Silver-gold-tin ternary systems

liquidus surface of, 91(F)vertical section through, 91(F)

Silver-gold-tin ternary systems phase diagrams, 90Silver-indium alloys, 232Silver-indium intermetallic compounds, specific types

AuIn2, 75Silver-indium-lead solders, specific types

Ag-80In-15Pb, melted in controlled atmosphere, 112(F)Silver-indium phase diagrams, 57(F)Silver-indium-tin solders, specific types

Ag-20In-77Sn, melting point of, 96Silver-lead phase diagrams, 61(F), 80Silver-lead solders, 6, 79–80Silver-lead solders, specific types

Ag-97.6Pb, silver-lead phase diagrams for, 80Silver-lead system diagram, 80(F)Silver-lead-tin solders, specific types

Ag-97.5Pb-1Sn, gold addition, effect of, 154(F)Ag-97.5Pb-1Sn, gold addition effects on, 154

Silver-mercury intermetallic compounds, specific typesAg2Hg3, 215

Silver oxide, 130(F)Silver particle reinforcement, 220Silver powder

in gallium-based amalgams, 216indium amalgams with, 217and mercury, diffusion reaction of, 215

Silver-tin alloys, 232




Silver-tin alloys, specific typeAg-96Sn, joints made by, 24

Silver-tin binary system, 96Silver-tin eutectic alloy, 152Silver-tin eutectic solders

3.6Ag-1.6Au-92.8Sn eutectic solders, 96effect of gold additions to, 154lead-free solders, 60reaction of, to silver, 231remelt temperature of, 232wetting effect or rare earth doping, 228–229

Silver-tin intermetallic compounds, specific typesAg3Sn, 57Ag3Sn, as continuous interfacial layer, 231–232Ag3Sn, as stoichiometric compounds, 91Ag3Sn, characteristics of, 91–92Ag3Sn, joint strength of, 232Ag3Sn, micrograph of, 92(F)Ag3Sn, rate of growth of, 92Ag3Sn, thickness of, 92(F)Au5Sn, creation of, 233

Silver-tin intermetallic phases, 52Silver-tin off-eutectic alloys

effect of bismuth additions to, 193(F)liquidus and solidus temperatures of, 193(F)

Silver-tin phase diagrams, 59(F), 81Silver-tin solders

eutectic alloys, 60vs. lead-tin solders, 60

Silver-tin solders, specific typesAg-95Sn, elongation to failure of, 228(F)Ag-95Sn, tensile strength of, 228(F)Ag-96.5Sn, 90, 193Ag-96.5Sn, contact angle of, 228(F)Ag-96.5Sn, for silver-coated components, 81Ag-96.5Sn, joint strength of, 232Ag-96.5Sn, spread area of, 228(F)Ag-96Sn, elongation, 194Ag-96Sn, gallium arsenide (GaAs) with, 134Ag-96Sn, in foil, radiograph of, 23(F)Ag-96Sn, mechanical properties of, 91(F)Ag-96Sn, properties of, 53, 90

Silver-tin systems, 231Silver-tin-zinc solders, specific types

3.5Ag-95.5Sn-1Zn, grain refinement in, 218Single phase materials, 159Solder

compatibility of, 49creep resistance of, 217deposition methods, 31–32fatigue resistance of, 217

Solder alloy familiesand melting ranges, 6(F)temperature ranges of, 6

Solder alloy systemsantimony in, 53–54bismuth in, 53gold in, 54indium and lead in, 52–53overview of, 49

silver in, 53survey of, 51–75tin in, 52zinc in, 54

Solder alloysconstituents of, 51surface tensions of, 15volatile constituents of, 6

Solder bridging, 207Solder bump bonding

process characteristics of, 202(T)process flow for, 199(T)self-aligning feature of, 201semiconductor components in, 203(F)technology characteristics of, 202(T)

Solder bumps, 205Solder coated substrates, 134Solder composites (reinforced solders), 222–223Solder drains, 206Solder elements, 52Solder flow

mechanically enhanced, 134metallurgically enhanced, 134–135

Solder foil vs. solder coated substrates, 134Solder oxides

reduction of, by atomic hydrogen, 127–128reduction of, by hydrogen, 126–127reduction rate of, in hydrogen, 128(F)self-dissolution of, 125–126

Solder pastes, 31Solder preforms. See preformsSolder reflow ovens, 174Solder reinforcement, 178Solder selection, 78Solderability

evaluation of, 207of selected metals and alloys, 121(T)

Solderability calibration standards, 212–214Solderability shelf life, 132

of gold-coated components, 148(T)of gold coating, 133

Solderability test cycle, 209(F)Solderability test measurements, 210Solderability test methods, 207–214Solderability testers, 208Solderable component surfaces, 133Soldered joints. See jointsSoldering

and brazing, distinction between, 6chemical fluxes for, 111–123design criteria, 28–30filler metal temperatures in, 5flip-chip-interconnections, 199–207fluxless, 123–137functional process of, 28–30health, safety and environmental aspects of, 42–43key parameters of, 12–28processes of, 28–45

Soldering and brazing, 3–8Soldering fumes, 43

Index / 265



Soldering iron, 129Soldering standards. See tests/testingSolders. See also specific solders

active, 152–153bulk properties of, indicative data on, 225–226for domestic water pipes, 51eutectic compositions of, 8flow enhancement, 133mechanical strengths, factors in, 223–224metallurgy of, 49–99powders, 31service temperature of, 8strengthening of, 217–222theoretical viscosity of, 19wetting of metals by, 147–157wetting of nonmetals by, 149–153

Solid-liquid interdiffusional bonding. See diffusionsoldering

Solid-liquid interfacial reactions, 24Solid-liquid reactions, 25Solid-solution strengthening, 53–54Solid solutions, 5Solid-state diffusion

tin-base intermetallic phases growth of by, 88(F)tin intermetallics by, 88

Solid-state diffusion processes, 25Solid-state joining

characteristics of, 5with gold, 43–44with indium, 43–44joint cleaning for, 5with solder constituents, 43–44temperature levels of, 5

Solid surfaces, 12Solidification casting, 32(F)Solidification shrinkage

control of voids from, 173magnitude of, 173of molten filler metals, 169of selected elements used in solders, 173(T)stress concentrations from, 50by vacancy diffusion, 173

Solids content, 118Solidus temperatures, 29, 62, 80, 228(F)Soluble coatings, 149Soluble halides, 118Solution treatment stages, 39Spherical cap, 44Spherical cap geometry, 44(F)Spontaneous chemical reaction, 137Spontaneous filtration, 223Spread characteristics on binary solder alloys, 21(F)Spread factor

and contact angle, 44–45defined, 45, 211and spread ratio and contact angle, relationship

between, 212(F)values for, 212(T)

Spread ratioand contact angle, 44–45

defined, 44, 211for a range of alloys, 212and spread factor and contact angle, relationship

between, 212(F)values for, 212(T)

Spread tests, 112–113(F)Spreading

area of, 20assessment of, 210–212atmosphere effects on, 20capillary action of, 23classical model of, 12driving force for, 84of eutectic alloys, 81irreversible nature of, 15metallization control of, 148of Pb60Sn solders, on gold-plated sample, 213(F)rate of, 22(F)vs. wetting, 49

Spreading behavior, 70(F)Spreading characteristics of lead-free solders, 197Spreading test, 212Sputter-deposited coatings, 148Sputtering, 178, 180Sputtering process, 150Stainless steels

bismuth-tin lead-free solders for, 122fluxes for, 122halogen-based fluxes for, 122nonmetallic bonding to, 152oxide layers of, 122phosphoric acid fluxes for, 122and temperature uniformity, 50thermal conductivity of, 50, 122

Standard deviation chart (s-charts), 42Standards

Scanning acoustic microscope, 183–184in soldering, 183–184visual inspection, 183x-ray inspection, 183–184

Statistical process control (SPC), 42Step butt joints, 177Step height, 176, 176(F)Step-joining process, 33Step soldered flip-chip interconnects, 206–207Stirling’s Formula, 98Stoichiometric compounds, 91Storage of multilayer metallic coatings, 50Strain energy density, 226–227Strain rate, 194Strap joints, 177, 177(F)Strength

of lap joints, and rare earth doping, 229mechanical, of brittle materials, 28of pressure welded joints, 9(F)

Strengthened solders, 178, 217–222Strengths of metals, practical, 28Stress

from differential thermal expansion, 157reduction in, from differential thermal expansion, 158




reduction of, by creep, 178Stress concentrations

defined, 175and fillets, 167fillets to reduce, 176by high thermal gradients, 4of metallization layer, 151origins and magnitude of, 175from solidification shrinkage, 50

Stress cycles, 227Stress distribution, 176(F)Stress relaxation, 158Stress-rupture life

of 3.8Ag-0.7-95.5Sn solders, 228(F)of low-melting-point solders, 195(F)as ranking for creep resistance, 195of Sn-3.5Ag-0.25RE, 228(F)of Sn-3.5Ag solders, 228(F)

Stress-strain curves for dental amalgams, 215, 215(F)Strip casting, 32(F)Sub-zero service temperature, 76Sublimation, 109(T)Super Solder system, 32–33Superheat

defined, 8to dissolve native oxides, 126(T)to dissolve surface oxides, 126for indium-base solders, 75and oxygen levels for melting point, 127(T)and solder spreading, 115

Surface-area-to-volume ratio, 133Surface cleaning, 129Surface condition

of components, 12requirements of, in diffusion brazing, 10

Surface conditioning, 131–132Surface diffusion, 10Surface energy

diagram of, 12(F)of a liquid, 13of pure metals, 14of a solid, 12and surface tension, 13and surface tension, diagram of, 13(F)

Surface energy and surface tension, 12–13Surface erosion, 3SURFACE EVOLVER (software), 12, 206Surface finishes, 149Surface insulation resistance test (SIR), 119

International Electrotechnical Commission (IEC) testcoupons for, 119(F)

proposed changes to, 120test plot of, showing dendritic growth, 120(F)

Surface metallization, 4Surface mount solderability test (SMT), 207Surface-mount technology, 217Surface-opening cavities, 163Surface oxides

removal of, 37superheats to dissolve, 126

Surface roughnessof cold-rolled copper, 14(T)of components, 22–25effect of, on fracture toughness, 24(F)and fracture toughness, 23–24and spreading, 23

Surface temperature, 25Surface tensions

of binary solders, 194(T)changes to various metal alloys, 205diagram of forces of, 13(F)of lead-free solders, 193–194between liquid and vapor, 13, 14of molten filler metals, 209of solder alloys, 15solder bump self-alignment, 205between solid and liquid, 13, 14between solid and vapor, 13, 14and surface energy, diagram of, 13(F)

Surface topography, 206Surfactants, 117Surroundings, defined, 103Sweeping of trapped gas, 133Symbols and abbreviations, 243Synthetically activated fluxes, 117

T

Tantalumdiffusion bonding of, 10joining atmospheres for, 104

Tape bonding, 203Tape interconnections, 202Temperature and O/H ratio, 110(F)Temperature effects

of contact angle, 25of oxide thickness growth, 124of surface temperature, 25of viscosity, 25

Temperature gradients, 12Temperature levels

of diffusion bonding process, 9of solid-state joining, 5

Temperature limits for brazing and soldering, 4Temperature measurements, 34–35, 38Temperature/pressure curve

for diffusion bonding of gold, 43(F)for diffusion bonding of indium, 44(F)

Temperature uniformity, 50Temperatures. See also melting point

activation, of fluxes, 117active filler metals requirements, 149of boiling, 109(T)of boiling/sublimation, for selected elements, 107control of, 38in diffusion brazing, 10of eutectiferous phase transformations, 94(T)of heat treatment, 39of heat treatment for reference standards, 213

Index / 267



Temperatures (continued)peak operating, 29peak process, 33peak reflow, process control chart of, 42(F)of reactive filler metal bonding, 152for reduction of indium oxides, 126for reduction of tin oxides, 126for reduction of zinc oxides, 126of sublimation, 109(T)

Tensile strengthof Ag-95Sn solders, 228(F)of composite solders, 221(T)and dendrites arm spacing, 32(F)of joints, 222of lead-tin alloys, 81(F)of lead-tin eutectic solders, 222(F)of Sn-9Zn solders, 228(F)

Tensile stresses, 175–176Ternary systems

intermetallic phases in, 84liquidus representation in phase diagrams of, 84representation of, in phase diagrams, 84

Tests/testingacceptance criteria, 209, 231ANSI/J-STD-002 (test method), 207basic spreading test, 211bolometers (thermal imaging), 35Chadwick peel tests, 73component testing, 224dip-and-look (DNL) test, 207edge fillets examination, 168EIA/IS-86 (test method), 207fitness for purpose tests, 224IEC board test coupons, 119(F)inspection methods for voids detection, 183inspections of joint interior, 183International Electrotechnical Commission (IEC),

119(F)IPC/EIA J-STD-003A (test method), 207of mechanical properties, 224Military Standards (MIL-STD), 183, 231of palladium plated devices, 207–208quality assurance testing, 210quality-control testing, 210scanning acoustic microscopy (SAM), 170(F), 189,

235–236scanning acoustic techniques, 183sessile drop tests, 209, 211(F), 220solderability calibration standards, 212–214solderability test measurements, 210solderability test methods, 207–214solderability testers, 208spread tests, 112–113(F)spreading test, 212standards, 183–184surface insulation resistance test (SIR), 119, 120surface insulation test plots for dendritic growth,

120(F)surface mount solderability test (SMT), 207test coupons, 119(F)

test plot showing dendritic growth, 120(F)wetting balance solderability test, 208–209wetting force during solderability test cycle, 209(F)

Thallium, 191Theoretical viscosity of solders, 19Thermal activation energy, 8Thermal characteristics of fuel gases, 34(T)Thermal conductance of mechanical fasteners, 1–2Thermal conductivity

of adhesive joints, 3of aluminum and aluminum alloys, 62of carbon fibers, 223of composite solders, 221in finite element analysis, 225of lead free solders, 198limits to, 30vs. linear expansion coefficient, 159(F)and rules of mixtures, 194of soldering atmospheres, 114(T)of stainless steels, 50, 122of titanium, 160of zinc-bearing solders, 63

Thermal cyclingwithout creep, 165without fatigue fracture, 165

Thermal distortion, 33Thermal distortion parameter, 159(F)Thermal expansion

of engineering ceramics, 159mismatch in, 145of single phase materials, 159

Thermal expansion mismatcheffects of, 157in flip chip interconnection, 200stress reduction technique, 165(F)

Thermal expansion mismatch strain, 50Thermal expansivity

of aluminum and aluminum alloys, 62of bismuth solders, 225of engineering materials, 26(T)reduction of, 222–223

Thermal fatigue, 29Thermal heat capacity, 50Thermal imaging bolometers, 35Thermal properties of joints, 50Thermally conductive adhesives, 3Thermally induced distortion, 62Thermocompression bonding, 5(F)Thermocouples, 34–35Thermodynamic and diffusion-kinetic model of

metallizations, 152Thermodynamic equilibrium, 137Thermodynamic principles for analyzing oxide

reduction, 106Thermodynamics

first law of, 137–138second law of, 138–139

Thick-film formulations, 180–181Thick film metallizations, 72Thick-gap soldering, 178–179




Thick solder joints, 179Thickness of coatings, 148Thin coatings, 180Thin foil preforms

disadvantages of, 136economics of, 136shapes of, 136

Thin-gap soldering, 178–179Thin (narrow) joints

with hydrostatic pressure, 178–179in semiconductor grade clean room, 179

Threshold formation, 9Through-thickness vias, 162Time requirements of diffusion brazing, 10Tin

activation energy of, 124allotropic transformation of, 76, 77(F), 195–196allotropic transformation suppression, 196as alloy element for Au-2wt%Si solders, 68as impurity, 77and indium intermetallic compounds, 51as solder constituent, 52

Tin-base eutectic solders, 153(T)Tin-base intermetallic phases, 88(F)Tin-base solders

3.5Ag-0.9Cu-95.6Sn ternary eutectic solders, 96Alloy J, 60–61dissolution rate of platinum in, 148(T)failure modes of, 74fluxes for, 116–120and gold coatings, 133as key to lead-free solders replacement, 192silver-tin, 60

Tin-bismuth solders, specific typesSn-65Bi, 15

Tin dendrites, 77Tin-indium solders, specific types

Sn-40In, 15Tin intermetallics, 88Tin-lead solder, 31–32Tin-lead solders, specific types

61.75Sn-38.05Pb-0.2Cu(wt%), 84Sn-40Pb, hermetically sealed

enclosures, 44Tin oxides

reduction in hydrogen atmospheres, 111temperature for reduction of, 126

Tin pestdescribed, 76–77, 195–196lead-free solders, 195–196sub-zero service temperature, 76–77susceptibility of puree tin-base solders to, 195–196

Tin-silver-rare earth soldersSn-3.5Ag-0.25RE, stress-rupture life of, 228(F)

Tin-silver solders, specific typesSn-3.5Ag, stress-rupture life of, 228(F)

Tin whiskerslead-free solders, 195–196on PCBs, 196

Tin-zinc alloys, specific typeSn-9Zn eutectic, corrosion of, 96

Tin-zinc-silver solders, specific typesSn-8.5Zn-1Ag, aluminum as impurity in, 77

Tin-zinc soldersSn-9Zn eutectic alloy, 96

Tin-zinc solders, specific typesSn-9Zn, elongation to failure of, 228(F)Sn-9Zn, rare earth doping of, 228Sn-9Zn, tensile strength of, 228(F)Sn-9Zn, wetting force of, 228(F)

Tinned surfaces, 151Titanium

as additive to Au-20Sn solders, 69as additive to Au-2Si solders, 69of coefficient of thermal expansion (CTE), 160diffusion bonding of, 10joining atmospheres for, 104as low-expansion metals, 160nonmetallic bonding to, 152thermal conductivity of, 160for zinc oxide ceramic metallization, 150

Titanium alloys, 10TO-220 semiconductor die, 66(F)Tombstoning, 157Tongue and groove joints, 177Toxicity, 191Transient liquid phase (TLP) joining. See diffusion

solderingTransition reactions, 83Transuranium elements, 227Trapped gas

in adhesively bonded joints, 173described, 169–173pressure variation method for, 174(F)reduction, by vapor-phase techniques, 171sweeping of, 133, 170(F)volume reduction of, 171

Triode sputtering, 180Tungsten

of coefficient of thermal expansion (CTE), 160diffusion bonding of, 10as low-expansion metals, 160

U

Ultimate tensile strength of lead-free solders, 194Ultrahigh vacuum system (UHV), 127Ultrasonic fluxing, 129, 130Ultrasonic power and wetted area, 131(F)Ultrasonic soldering, 128–130Ultrasonic systems, 129–130Ultrasound, defined, 235Underbump metal, 200Underfill adhesive, 203Universal gas constant, 97, 124UNS K94610, 160“Unsolderable materials,” 120–122

Index / 269



V

Vacancy diffusion, 173Vacuum atmospheres

brasses and zinc metallization, 107industrial quality, 107soldering in, 107–109

Vacuum bakeout, 172Vacuum deposition process, 206Vacuum evaporation, 180Vacuum furnace, 104Vacuum joining, 174Vacuum system, 36Van der Waals forces, 15Vanadium, 104Vanadium ions, 132Vapor deposition, 31Vapor deposition coating, 234Vapor deposition technique, 148Vapor-phase techniques, 171Velocity of liquid flow, 19Ventilation, 43Viscosity

and fluid flow of solders, 18and molecular weight of metals, 19of selected conductive adhesives, 3(T)temperature effects of, 25theoretical, of solders, 19

Void content vs. joint length, 169(F)Void-free joints, 183Void levels

due to trapped gas, 174(F)vs. load applied to preforms, 134(F)and pressure variation, 174(F)

Void size, maximum acceptable, 183–184Voids

and carbon-fiber-loaded solders, 223causes of, 169control of, 173and diffusion bonding, 9–10effect of, on joint mechanical integrity, 29with fluxed paste solder, 172–173formation of, through gas entrapment, 25as function of component size, 169and gap width, 15incorporation of, 168inspection methods for detection of, 183in joint gap, 169(F), 170(F)joint interfaces as source of, 28in joints, 49Kirkendall voids, 71, 234in phase segregation, 74relationships of, to flux conditions, 172–173

Volatile organic compounds (VOCs), 118environmental considerations from, 196environmental considerations of, 115

Volatilization, 107Volume contraction

cracks from, 67of silver-tin alloys, 232

Volume diffusion, 10Volume freezing, 173Volume of spherical cap, 44

W

Water-soluble fluxes, 39, 40, 118Water vapor

bakeout temperature of, 36in hydrogen reducing atmosphere, 37ingress into hermetically sealed package, 185(F)as source of oxidation, 36

Water vapor desorption and oxygen partial pressure,108

Wave soldering, 33, 114Wave soldering machines, 77Weight fraction of constituents, 96Welding, 4Welding process, 4Wet back process, 206Wet plating

for application of solderable metallizations, 148metallization techniques, 180on nonmetallic components, 149–150

Wettability of parent materials, 49Wetted area and ultrasonic power, 131(F)Wetting

along surface valleys, 22–23assessment of, 207–210classical model of, 12, 14difficulties of silica in, 69driving force for, 16effects of, by rare earth doping, 228–229and fillet formation, 6of metals by solders, 147–157of nonmetals by solders, 149–153problems with, 17, 147rate of, 18refractory metals problems with, 147with refractory parent material, 147

Wetting and contact angle, 13–18Wetting and spreading, 20, 135Wetting and spreading characteristics, 220Wetting angle, 14(F), 212Wetting area, 13–14Wetting balance, 208(F)Wetting balance solderability test, 208–209Wetting balance tests, 197Wetting behavior, 210(F)Wetting characteristics, 197Wetting equation, 13, 18

for binary metal systems, 15Wetting force

for commercial fluxes, 214(F)of Sn-9Zn solders, 228(F)during solderability test cycle, 209(F)time of, to reach acceptance value, 197

Wetting front, 23Wetting properties, 221(T)




Wetting rate, 198(F)Wide-gap joints

brazing of, 179in ceramic components, 158composite solders and, 221

Wire bonding, 203Wire cross preform, 171(F), 172(F)Work-hardened alloys, 49Work hardening, 49WS type fluxes, 118

X

X-bar chart, 42X-radiography, 172(F)X-ray inspection and defects, 204X-ray systems

fine-focus, 189microfocus, 236–237

X-ray techniques, 183X-rays, cracks and orientation of, 236

Y

Yield strengthof composite solders, 220(F)of lead-free solders, 194

Young’s equation, 13, 15Young’s modulus, 223, 225

Z

Z-axis control, 205Zinc

as additive to indium, 135as impurity of Au-2Si solders, 77as solder constituent, 54

Zinc alloysAl-94Zn solders, 54with aluminum, 54

galvanic corrosion of, 64Zinc-aluminum alloys, specific type

90% Zn, 7% Al filler metal, 136Zinc-aluminum-magnesium-gallium alloys, specific

typeZn-4Al-3Mg-3.2Ga quaternary, high-melting-point

solders, 197–198Zn-4Al-3Mg-3.2Ga quaternary, high thermal

conductivity of, 197–198Zinc-base alloys, 64Zinc-base solders

high-volume contraction of, 64as lead-free solders, 66limitations of, 50stress concentrations of, 64wetting additives for, 147

Zinc-bearing alloys, 67(F)Zinc-bearing solders, 61–64

advantages of, 63Al-94Zn eutectic solders, 62for aluminum and aluminum alloys, 61cadmium alloy of, 63common alloys of, 62–63fluxes for, 65limitations of, 64and reduced pressure atmospheres, 62, 65solidus temperature range of, 62thermal conductivity of, 63

Zinc chloride, 117Zinc-containing alloys, 195Zinc metallization, 107Zinc oxide ceramic

metallizations of, 150metallized with titanium, 150metallized with zirconium, 150

Zinc oxidesstability of, 125temperature for reduction of, 126

Zinc solders, 116Zirconium

joining atmospheres for, 104for zinc oxide ceramic metallization, 150

Index / 271



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