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To my family and my
dear uncle Dr. Shawky Hammam
List of included papers
I Tin and Tin Alloy Coatings for Electrical Contacts- A Literature Review.Tag Hammam and Anders Kamf. Proceedings of the 31st Annual Connector & Interconnection Technology Symposium, USA, 1998, pp 105 –117.
II Friction, Wear and Electrical Properties of Tin-Coated Tin Bronze for Separable Electric Connectors.Tag Hammam. Proceedings of the 42nd Holm Conference on Electrical Contacts, USA, 1996, pp 321-330.
III Tin Coating Techniques for Copper-Base Alloys - The Effects on Friction, Wear and Electric Properties.Tag Hammam. Proceedings of the 43rd Holm Conferences on Electrical Contacts, USA, 1997, pp 201-211.
IV The Impact of Sliding Motion and Current Load on the Deterioration of Tin-Coated Contact Terminals. Tag Hammam. IEEE Transactions CPMT, vol. 23. No. 2. June 2000, pp 278 –285.
V Heat-Treatment of Tin-Coated Copper Base Alloy and the Subsequent Effect on Friction, Wear and Electric Properties. Tag Hammam and Rolf Sundberg. Proceedings of the 20th International Conference on Electrical Contacts, Sweden, 2000, pp 291-296.
VI Friction and Electric Properties of Lubricated Tin-Coated Contact Terminals.Tag Hammam, Åsa Kassman-Rudolphi and Larz Ignberg Proceedings of the 34th Annual Connector & Interconnection Technology Symposium, USA, 2001, pp 5 – 14.
VII Vibration-Induced Deterioration of Tin-Coated Connectors Studied by Using a Force Controlled Fretting Bench Test. Tag Hammam, Åsa Kassman-Rudolphi and Per Lundström. Proceedings of the 51st Holm Conference on Electrical Contacts, USA, 2005, pp 97 –106. Selected to IEEE Transactions CPMT.
VIII Development of a Tin-Coating Alloy with Improved Fretting Properties. Tag Hammam. In manuscript.
Contents
Preface ............................................................................................................9
1. Introduction...............................................................................................101.1 Background ........................................................................................101.2 Aims of the work................................................................................11
2 Review of tin coated electrical contacts.....................................................122.1 Fundamental contact theory and the nature of contact failure ...........122.2 Properties and structure of pure tin ....................................................132.3 Formation and growth of intermetallic compounds ...........................132.4 Laboratory contact testing..................................................................142.5 Fretting corrosion ...............................................................................152.6 Corrosion properties of tin and tin alloys...........................................172.7 Tin coating techniques .......................................................................182.8 Present trends for tin-coatings............................................................21
3 Summary and discussion of important results ...........................................233.1 Experimental set-up for simulating the insertion/withdrawal ............233.2 Contact behavior during insertion/withdrawal (Papers II and IV) .....243.3 The relation between drawing force and contact resistance (Paper II)..................................................................................................................253.4 The effect of tin coating techniques on the performance (Paper III)..273.5 The effect of heat-treatment (Paper V)...............................................293.6 The conditions for transition to a high contact resistance (Paper IV) 313.7 Model for prediction of contact failure (Paper IV) ............................323.8 Vibration induced fretting (Paper VII) ...............................................35
3.8.1 Experimental set-up ....................................................................353.8.2 Test conditions............................................................................363.8.3 Result ..........................................................................................36
3.9 Properties of lubricated contact terminals (Paper VI) ........................383.10 Properties of lubricated contacts with textured surface....................383.11 Development of a tin alloy coating with improved fretting properties (Paper VIII) ..............................................................................................40
4 Summary of contribution ...........................................................................42
5 Sammanfattning (Summary in Swedish) ...................................................44
Acknowledgement ........................................................................................46
References.....................................................................................................48
9
Preface
The work presented in this thesis was carried out at Corrosion and Metals Research Institute (KIMAB), Stockholm and Uppsala University, during the years 1996-2006 under the supervision of Dr. Åsa Kassman-Rudolphi and Prof. Sture Hogmark at Uppsala University. The author has done the major part of the planning, almost all experimental work at the lab, the interpretation of the results, and finally, essentially all the writing.
The thesis consists of two parts, the first of which has the purpose of giving an introduction to tin-coated connectors. Part two is a summary of the results of the eight enclosed papers, with the exception of section 3.10 that only appears in part two.
10
1. Introduction
1.1 Background The complexity of the electronic system in vehicles is rapidly increasing and so also is the frequency of electronic faults in new cars1. Field data has shown that 30 - 60% of operational disturbance or electrical failures are attributed to connectors2. Therefore, the reliability of electronic systems in vehicles during long-term operation has become a very important issue.
The total number of contact terminals in a modern car is in the range between 2500 and 3000, while the average value of connectors per vehicle has been estimated to $126 (2002)3. The total value of contacts produced worldwide for the car and truck industries is estimated at $6.9 billions (2002) 3. Thus, the total number of automotive connectors that are produced per year worldwide is approximately 150 billions. However, despite the huge number of connectors produced every year and the potential catastrophic cost that the introduction of poor connectors might cause, connectors do not in general get the same attention as most other electrical components.
Generally speaking, the main reasons to use separable connectors are three: manufacturing, mounting, and maintenance. These reasons also determine the overall design of a connector. Tin-coated contact terminals are frequently used for automotive connectors, mainly due to their good corrosion resistance and low cost. A typical automotive connector is shown in Figure 1. The trends for automotive connectors are towards smaller dimensions, lower contact forces, higher ambient temperature, increased vibration environment and, finally, demands for improved reliability. A review of automotive connectors by Anthony Lee is presented in the book Electrical Contacts, Principles and Applications 4.
11
Figure 1. A typical automotive connector for electric connection of e.g. fuels pumps and electric fans 5.
The soft tin-plating has a highly resistive oxide, which protects the metal from further corrosion. During insertion of a tin-coated connector the top oxide layer is wiped off, and large oxide free contact spots are obtained. This results in a low contact resistance, which is maintained even in a rather aggressive atmosphere. However, tin-coated connectors need particular attention with regard to four conditions:
1. The wear of the soft tin coating caused by insertion/withdrawal, 2. Fretting corrosion caused by vibration and thermally induced
movement. 3. The formation of intermetallic compounds caused by diffusion of
copper into the tin layer during long-term operation. 4. The relatively high insertion/withdrawal force.
1.2 Aims of the work The aims of this work have been the following:
1. To characterize the friction, wear and electrical properties of tin coatings as well as to develop appropriate test methods.
2. To achieve an improved fundamental understanding of the deterioration processes of tin coated connectors caused by wear and fretting corrosion.
3. To propose improvement of pre-tinned strip material for connectors.
12
2 Review of tin coated electrical contacts
2.1 Fundamental contact theory and the nature of contact failure There are many different types of separable electrical connectors. However, they all have in common that they consist of two mating parts that should offer electrical contact. A detailed description of the fundamental contact physics is given in Electrical Contacts6.
The contact interface between two solid bodies will have a restricted number of conductive spots (a-spots) that the current will pass through. The distortion of the current flow will cause an increased electrical contact resistance and is defined as the constriction resistance. The total interfacial resistance caused by constriction and surface films resistance is defined as the contact resistance. To achieve a low contact resistance the surface oxide should be removed, and the real contact area should be large. Tin is a very soft coating and therefore a large contact area can be achieved with a relative low normal force. Moreover, the initial surface oxide is, in general, wiped off during the insertion stroke and, thus, a low initial contact resistance can easily be achieved. A contact terminal has in general a redundancy of contact spots in order to reduce the risk for catastrophic contact failure.
The nature of contact failure caused by corrosion or fretting corrosion is in general a continuously increased contact resistance followed by an intermittent very high contact resistance and finally a constant high contact resistance. The interrupts might be in the range of nano-seconds up to microseconds7. Moreover, for power connectors, the initial increasing contact resistance might cause thermal runaway and subsequently catastrophic failure.
The increase in contact resistance is mainly due to formation of non-conductive oxide films at the contact interface. A contact can work well many times and pass tests, despite 95 -99 % of the initial conductive area having been destroyed. However, the margin to failure is almost eliminated. There is also a chance that self-healing or self-repair will occur due to mechanical or electrical penetration of the oxide films. Therefore, very often when no fault is discovered it is highly likely that a contact has caused the
13
disturbance. A quite common way to “repair” a contact that has failed is to do a repeated number of insertions and withdrawals in order to break up the formed oxide films. However, there is a significant risk that the connector will fail again.
The impact of a contact failure comes often suddenly and unexpectedly. However, generally the break down of an electrical contact is not a sudden event, but the final result of a continuously long-term deterioration process.
2.2 Properties and structure of pure tin The most characteristic properties of pure ( ) tin are: low melting point, good corrosion resistance, silver-white appearance, low hardness, ductility and a relatively poor conductivity, see also Table 1. The low melting point (232 C) of tin means that, on the absolute scale, room temperature is nearly 60 % of the melting temperature. This results in a low hardness at room temperature, but also in fast recrystallisation and thereby only slight work-hardening at deformation, high creep rate and a high solid diffusion rate of impurities into the tin.
Table 1. Mechanical and electrical properties of solid tin at 20 C
Tensile strength,
(MPa)
Shearstrength,
(MPa)
Knoophardness(kg/mm)
Resistivity, (10-8 m)
Conductivity
(% IACS)
Solid tin 14.5 8 12.3 8 8 9 12.6 8 13.7
2.3 Formation and growth of intermetallic compounds A tin coating easily forms discrete intermetallic compounds with the copper substrate. The solid diffusion of the substrate material into the tin coating is mainly a function of time and temperature10, but also a high current density can promote the diffusion rate11. A conventional diffusion barrier layer, e.g. of nickel, between the substrate and the tin coating will restrict, but not prevent, the solid diffusion, since the barrier layer itself also forms intermetallic compounds. The surface structure of the intermetallic compound is uneven, and is reminiscent of cobblestones, see Figs. 2 and 3. The hardness of Cu6Sn5 is approximately 40 times the hardness of pure tin 9,while the conductivity is almost the same 12.
14
Figure 2. SEM-image, top-view of the intermetallic compound visualized by dissolving the free tin layer.
Figure 3. SEM-image, compositional contrast of pure Sn coating, heat-treated 1000 h at 150 C. Cross-section, from top to bottom; support, free tin (white), Cu6Sn5 layer (gray), Cu3Sn (dark gray) and finally the substrate (black).
The growth of Cu6Sn5 intermetallic compound is initially characterized by diffusion of copper into the tin layer and subsequent island formation. Further diffusion results in coalescence and secondary island growth in the voids, until a tight intermetallic phase covers the interface between the substrate and the tin layer. The diffusion of copper into the Cu6Sn5 phase is slower than into pure tin. However, it does occur, resulting in a Cu3Sn phase between the substrate and the Cu6Sn5 phase. Theoretically, if the heat-treatment is sufficiently long, all the Cu6Sn5 phase will be transformed to a Cu3Sn phase. Detailed descriptions of the crystal structure of tin intermetallics and crystal growth on solid surfaces can be found in A.K. Larsson (1994) 13 and J.A. Venables et al. (1984) 14 respectively.
2.4 Laboratory contact testing A large number of standards and specifications related to testing automotive connectors have been developed, e.g. by ASTM, EIA SAE, and ISO 4. One of the most recently developed standard tests is the SAE/USCAR-20 test (Field Correlated Life Test). All these tests are used to verify if a connector can be used in a vehicle. A major problem with accelerated laboratory tests is that they often involve unrealistic stresses, e.g. vibration tests under severe conditions might cause self-healing of poor contacts. Moreover, the failure
15
mechanisms that cause failure in the laboratory test might not be relevant during real operating conditions.
For fundamental studies of contact terminals, test methods are needed that address the tests to specific failure mechanisms. Therefore, bench tests with model contacts are used for fundamental studies. The two most common model contacts are a ball (i.e. hemispherical rider) on flat and crossed cylinders.
The majority of the published data on the performance of connectors presupposes the magnitude of the contact resistance as the critical parameter. The contact resistance measurement is usually performed under dry circuit conditions according to ASTM B359-80. The test includes 4-wire measurement, open-circuit voltage limited to 20 mV, and maximum current load equal to 100 mA. The measurement signal under dry circuit conditions usually becomes noisy.
Most laboratory tests are dealing with either fretting or corrosion. Å. K. Rudolphi has performed work in a test bench where the contact can be subjected to a corrosive environment as well as current load during fretting testing 15. The main conclusion from experiments with silver-coated contacts was that the mechanical degradation dominates as a contact failure mechanism, even in a severely corrosive atmosphere 16. J. Swingler et al have also studied synergetic effects caused by a combination of corrosion and vibration 17.
A major problem in all accelerated contact testing is to estimate how the lab test results apply to real operating conditions. Contact failure might occur after only a few seconds in an accelerated test, while a real contact will be operating for many years.
2.5 Fretting corrosion The main deterioration mechanism for automotive tin contacts is fretting corrosion, mainly caused by the vibration from the engine, suspension when traveling and thermally induced movement, according to J. Swingler and J. W. McBride 2. They have performed an extensive study on the deterioration of automotive connectors during actual operating conditions, in order to observe the environmental stresses during operation. Fretting results in a thick and highly resistive oxide film at the contact interface even in a mild environment. Reviews of fretting corrosion have been written by e.g. M. Antler 18 and by R. Malucci 19.
Vibrations cause externally induced forces acting on the contact and are a function of the acceleration and retardation of the moving mass. Both the magnitude of the acceleration and the mass of the contact depend to a high degree on the design of the contact 20. Fretting corrosion with subsequent
16
contact failure might occur almost instantly when the operating frequency resonates with the frequency of the connector system 21.
The displacement that can occur in a contact can be divided into three basic “fretting modes”, characterized by a linear, radial, or circumferential trajectory, defined as fretting Mode I, II and III respectively 22. The three fretting modes are depicted in Figure 2.
Mode I Linear relative displacement
Mode II Radial relative displacement
Mode III Circumferential relative displacement
Figure 2. Schematic figure of the three basic fretting modes in the case of ball-on-flat contact geometry 22.
Fretting Mode I in combination with ball-on-flat contact is the most commonly studied fretting type and is characterized by three different contact regimes, stick, partial slip and gross slip 16. During stick conditions the micro-displacement is mainly accommodated by elastic deformation, and thus relative sliding is totally prevented. Increased vibration results in an increased shear stress, that first will exceed the friction shear stress in the periphery of the contact area. This will result in slip of the periphery of the contact spot, and partial slip is achieved, see Figure 3.
Stick
Slip
-a a-a' a'r
p(r)
q(r)
Figure 3. Fretting Mode I in combination with ball-on-flat contact under partial slip conditions. The contact area is divided into a central stick area and a surrounding slip area. The distribution of pressure p(r) and tangential traction q(r) under elastic conditions is depicted 15.
17
Further increase of the vibration results in an increased slip annulus and thus a reduced stick central area. At a certain magnitude of the vibration the residual stick area fractures and sliding starts to occur over the entire contact area, i.e. gross sliding is achieved. Several studies have shown that there exists a distinctive vibration threshold value, below which the contact resistance will be stable 23, 24.
The time to failure caused by fretting is prolonged if the sliding amplitude is restricted, as well as if the contact force is increased 25. In a real connector, increased contact load will also in most cases reduce the interface motion, and therefore, a high contact force is often recommended for tin-coated connectors. However, lubricants might be the most effective way of inhibiting fretting corrosion and thereby preventing the increase in the contact resistance. Lubricants prevent fretting by reducing the adhesive wear and by shielding the surface from the air and thereby reducing the rate of oxide formation 26, 27. An appropriate lubricant can extend the time to failure 100-500 times 28.
2.6 Corrosion properties of tin and tin alloys Tin is quickly covered with an oxide film, which protects it from further corrosion even in rather aggressive atmospheres. A notable advantage, beside the generally good corrosion resistance, is that tin is not prone to pitting and creep corrosion, in contrast to gold for instance. However, chlorine and chloride-containing pollutants attack the film and severe corrosion can occur. A high corrosion rate has also been reported in field tests with air containing halogenide in combination with high humidity 29.
J.L Josta et al. have studied the contact resistance of different tin coatings before and after exposure to a corrosive atmosphere and some of the results are shown in Figure 4 and Figure 5. The initial contact resistance value is given, as well as the contact resistance after the coatings have been subjected to a simulated industrial environment (1 0.3 ppm H2S and (10 3 ppm SO2). Test temperature and humidity are 23 3 C and 75 5 % respectively. The contact resistance was measured by using a moving probe, with a normal load of 1 N. The general conclusion from the corrosion tests is that pure tin or tin-lead alloys resist industrial environment better than tin alloys.
18
Sn (ele
ctrole
ss
Sn (matt
)
Sn (bri
ght)
PbSn 4
0/60 (
mat
PbSn (
reflow
)
PbSn/S
n
PbSn (
brigh
t)
Con
tact
resi
stan
ce (m
)
100
101
102
103
104106
as-received H2SSO2
CuSn 4
5/55
SnZn 6
5/35
SnZn 5
0/50
CuSnZ
n 58/2
4/18
SnNi 7
0/30
SnNi/S
n
SnNi/A
u
Con
tact
resi
stan
ce (m
)
100
101
102
103
104106
as-received H2SSO2
Figure 4. Contact resistance for tin and tin-lead coatings30.
Figure 5. Contact resistance for tin-alloys30.
2.7 Tin coating techniques Commercial tin coatings are deposited by one of three methods: hot dipping, electroplating, and electroplating with a subsequent quick fusion and solidification of the tin coating, also known as reflow tin. Tin can also be deposited by using e.g. pulse plating31 (electroplating with modulated current), chemical deposition, and glow-discharge ion-plating32, but none of these methods are used for large volume production.
The most economic way to tin coat strips is by hot dipping. Wide strips of copper-base alloy are cleaned, coated with an activated flux, and thereafter fed at a high speed into a bath of molten tin. The excess tin is removed by either mechanical wipers or air knifes. Usually, mechanical wiping is used for coating thicknesses between 0.5 and 2 m, while air knife wiping is used for a desired coating thickness above 1.5 – 2 m. If hot air is used during air wiping, the coating is usually named HALT (Hot Air Levelled Tin). Furthermore, hot dipped tinning inevitably results in a thick, hard, and uneven intermetallic compound formed between the copper substrate and the free tin layer.
For commercial electroplating of tin there are two main groups of electrolyte baths, resulting in matte and bright deposits respectively. To achieve a bright surface, organic brighteners are usually added to the electrolytic solution. Therefore the bright coating contains varying amounts
19
of co-deposited organic matter, 0,05 % to 0,25 %. The reflow tin achieves a very smooth and bright surface, since tin and tin-rich alloys exhibit extreme fluidity at just above the melting point. Besides the cosmetic appearance, a bright surface has several benefits compared with matte tin, e.g. finger marking and surface contamination by dust are reduced. Whiskers may grow slowly from the surface of electroplated pure tin and reach a length up to 1cm with a diameter of 1-5 m. They may cause short-circuiting if they bridge adjacent contacts33. The growth can be suppressed by using a nickel barrier layer34, but it is more efficient to do a quick fusion and solidification of the tin coating, i.e. reflow tin.
The electroplating process is a low-cost process that is suitable for large-volume production. Strips can be electroplated continuously but, in contrast to hot dipped tin, stamped parts can also be coated, which helps to economize the amount of tin used. However, incorporation of impurities from the electrolyte to the coating, and a large number of process variables that influence the coatings properties can give quality problems. For example, the hardness of an electroplated coating depends upon current density, bath temperature, presence of addition agents or foreign metals, pH nature of the solution, and various anions and cations present35. Finally, the thickness of the plating is normally increased at edges and corners, and is highly affected by current density and type of bath.
The main differences between the depositing techniques regarding material properties are grain size thickness of the achieved intermetallic compound, variation in coating thickness, and surface roughness.
Table 2 gives an overview of practical experience of the different deposition techniques.
20
Table 2. Overview of the practical experience of the most common tin deposition techniques, after Anders Kamf 36. A=Average B=Better W=Worse NC=Not Common, R= Recommended.
HOT TINNING EL COATING REFLOW
Properties Mech. wipe Air wipe Matte Bright
Thickness ( m)
0.5-2 1 - 3 > 4 0.8–1.5 0.8–1.5
Undercoating NC NC R R R
Thickness tolerances
W W B B A
Appearance W W A B A
Adhesion B B W W A
Porosity A A W A A
Grain size B B W W A
Whiskers B B W W B
Solderability W A A A B
Contact resistance
W A B B B
Formability B B W W A
Coefficient of friction
B W A A B
Ageing/cor.Resistance
W B A A B
Fretting corrosion
B A W W A
21
2.8 Present trends for tin-coatings The required insertion force has a significant impact on the design and the cost of a connector. During insertion of a connector without mechanical assist engagement, the insertion force arises due to the normal contact force causing sliding resistance between the contacting surfaces. In addition, depending on the design of the connector, during insertion, mechanical resistance from the connector seal and inertia lock might be superimposed on the contact friction force. The trends are towards an increasing number of contact terminals integrated into a single connector, and thus, the insertion force for the individual contact terminal has to be minimized to achieve a reasonable insertion force.
Therefore, the main trend is to reduce the insertion force by using thinner tin coatings. Reduced coating thickness will reduce the insertion force mainly due to a reduced contact area. To achieve as low an insertion force as possible, very thin electroplated tin-coatings (0.5 m to 1,0 m) have been introduced, see Figure 6. Moreover, there is an increased demand on the tin-coated surface finish as well as on reduced variation in coating thickness, i.e. increased requirements for the coating process 37.
0
1
2
3
4
5
6
0 100 200 300 400 500 600Normal load (gms)
Dra
win
g fo
rce
(N)
EL Matte 1.3-3.8 µm
EL thin Matte 1 - 2 µm
Hot Dip AW 1.3 -3.8µmReflow Tin
Hot Dip MW
EL Bright
Reflow Tin SuperThin 0.5-1 µm
Figure 6. Drawing force versus normal force. Reduced tin coating thickness will lower the friction 38. AW = Air Wiped, MW = Mechanical Wiped.
22
Until recently, usually at least some few percent of lead was added to the tin. However, due to the planned outlawing of lead in electronic products, a large number of lead-free tin alloys have been proposed. The main alternatives that have been discussed are presented in Table 3. The tin-lead has now (2005) been, or will be, replaced mainly by pure tin and also to some extent tin-silver. The pure tin is the main choice due to the low cost39.
Table 3. Tin versus other lead-free alternatives, ref. 39, if nothing else is stated.
Item Pure Sn Sn/Cu Sn/Ag Sn/Bi Sn/Zn Au flash over Ni or Pd-Ni
Process Simple Complex Complex Complex Complex Easy
Scrap value OK OK OK Bismuth content
OK OK
Immersion plate
No Yes Yes Yes No No
Melting point ( C)
232 C 272 C 221 C 138 C 199 C 1063 C
Solderability 60/40
OK OK OK OK OK OK
Solderabilty Pb-free
OK OK OK OK ? OK
Joint reliability
OK Stressed Stressed Brittle Brittle OK
Fretting corrosion
Better than SnPb 40
Betterthan Sn / SnPb 40 ,41
Corrosion Minor Some More H2S
Some Major Minor
Tarnish Yes Yes Yes Yes Yes N/A
Whiskers Manage-able risk
Risk Risk Risk Large risk
No
Cost Low Low Medium Medium Medium High
23
3 Summary and discussion of important results
3.1 Experimental set-up for simulating the insertion/withdrawalTest equipment was developed in order to simulate the macro wear during insertion/withdrawal of a typical automotive connector, see Figure 7. The contact used consisted of a flat sample and a moving hemispherical rider with a tip radius of 2 mm. The length of the sliding stroke was 5 mm, and normal loads in the range 1-10 N was used in the experiments. The speed of the moving sample (rider) was held constant between the two reversal points. A force transducer measured the drawing force.
F=mg
Drawing force
Figure 7. The test bench to study friction, wear and electric properties of materials for separable connectors.
The contacts were subjected to electrical loading during wear testing. The four-point measuring method was used to determine the electrical contact resistance. In order to simulate the current load of a typical automotive contact a direct current (DC) source with an output in the range of 8 mA up to 20 A was utilized, and the open-circuit voltage was not limited. Thus, the measurement was not performed under dry circuit conditions.
The current and voltage drop were measured by digital multimeters with a resolution of 100 nV. The first and last 2 % of the total data sampled during one stroke were omitted, as the resistance is influenced by stop/start effects during reversing of the motion. Thereafter the mean and standard deviation
24
of the drawing force and the voltage drop for every half cycle (one stroke) was calculated.
3.2 Contact behavior during insertion/withdrawal (Papers II and IV) Three different commercial tin-coatings were included in the experiments: electroplated dull tin, hot dipped air wiped tin, and reflow tin. The thickness of the coating ranged between 0.5 and 10 m, and the same test material was chosen for both the flat sample and the rider in all experiments. Finally, all experiments were performed without any lubrication, except were otherwise stated.
Three characteristic stages were observed and defined during the macro wear testing of the tin-coated samples, see Figure 8 and Figure 9. The first stage (Stage I) is characterized by plowing of the rider into the soft tin layer. The second stage (Stage II) is characterized by sliding of the rider on the hard intermetallic compound, and finally in the third stage (Stage III), the intermetallic compound is penetrated and the rider starts to plow into the underlying substrate material.
Figure 8. Results representing the variation in drawing force throughout the three main stages during macro wear.
Figure 9. Results representing the variation in the voltage drop throughout the three main stages during macro wear.
The end of Stage I is defined as the combined minimum of the mean value and standard deviation of the recorded drawing force. Stage II is characterized by a continuously increasing contact resistance through to a
25
point where it starts rising sharply to a relatively high value. Furthermore, Stage II is divided into two “sub-stages”, referred to as Stage IIa and Stage IIb. The transition point between the two is defined by a sharp rise in the magnitude of the voltage drop, see Figure 9. A general view of the different stages during macro wear is presented in Table 4.
Table 4. General view of the different defined stages during macro-wear.
Stage Contact condition Contact resistance
Friction
I The rider plows into the soft tin. Low High and unstable
IIa The rider reaches and starts to slide on the hard intermetallic compound.
Continuously increasing
Low and stable
IIb The rider continues to slide on the hard, but now, heavily oxidized intermetallic compound.
High Low and stable
III The rider penetrates the intermetallic compound and starts to plow into the substrate.
Relative low High and very unstable
The soft tin layer of the nominal 1-2 m thick tin coatings was removed after only a few cycles (2-5), while the free tin layer of the nominal 10 m thick coatings resisted at least 10 cycles. The influence of the normal load on the wear rate was not very significant within the tested range of normal loads (1-10N). The number of insertions/withdrawals before the hard intermetallic compound was penetrated depended on the thickness of the intermetallic compound, the hardness of the substrate, as well as the applied normal load.
3.3 The relation between drawing force and contact resistance (Paper II) In general, the main requirement for a plugin connector is as low as an insertion force as possible without achieving an unstable contact resistance. Thus, the relationship to study is the insertion/withdrawal force versus the contact resistance. This can be described and analyzed by using a proposed semi-empirical relationship between the electric contact resistance (RC) and the draw force (FD). The proposed semi-empirical relationship between the
26
electric contact resistance (RC ) and the draw force (FD) involves the following assumptions 42:
1. The contact area is a single circular contact spot. 2. The draw force F is assumed to be the sum of two parts: The first
part is proportional to the contact area. This part arises mainly from the adhesion force during sliding. The second part is a constant force, mainly involving plowing.
3. The draw force F is assumed to be the sum of two parts: The first part is proportional to the contact area. This part arises mainly from the adhesion force during sliding. The second part is a constant force, mainly involving plowing.
These conditions are approximately fulfilled during Stage I, i.e. the rider is plowing into the soft tin layer. The proposed model of the contact situation is illustrated in Figure 10.
Figure 10. Simplified model of the contact situation, a cylindrical rider is plowing into a homogenous thick coating.
The following nomenclature was used:
Electric contact resistance RC Plowing force FP
Resistivity Contact temperature ( K) TC
Friction shear strength Substrate temperature ( K) TS
Conducting contact diameter D Contact voltage drop U
Draw force FD
27
The draw force is equal to: PD FDF4
2 (1)
The electric contact resistance through a single circular contact spot which is according to R.Holm 43:
DRC
(2)
The diameter D can be derived from equation (1) and inserted in (2) gives:
)(2 PDC FF
R (3)
3.4 The effect of tin coating techniques on the performance (Paper III) Three different commercial tin-coatings were included in the experiments; electroplated dull, hot dipped air wiped, and reflow tin. The experimental set-up used is described in section 3.1. The mean contact resistance versus draw force during Stage I for the hot-dipped, electroplated and reflow tin are illustrated in Figs. 11, 12 and 13 respectively. In the figure the estimated relation according equation 3 is also illustrated. The correlation between experimental data and adjusted curves is good. Three parameters have been adjusted to fit the curves along the measured values: shear stress of Sn ( ),Sn resistivity ( ) and the plowing force FP.
Figure 11. The mean contact resistance versus draw force during stage-I for the hot dipped 2 and 10 m thick tin coating.
Figure 12. The mean contact resistance versus draw force during stage-I for the el.plated 2 and 10 m thick tin coating.
28
Figure 13. The mean contact resistance versus drawing force during stage-I for the reflowed 2 m thick tin coating.
The estimated shear strength and resistivity of the coatings,
Table 5, based on experiments and the model (3), indicate a small difference in material properties for the three types of coatings. The shear stress reaches the same value as the shear strength for solid tin (12.3 MPa), while the resistivity for the coatings is twice as high as for solid tin (12.6 10-8 m). In addition, results from experiments with lubricated electroplated samples are given in Table 5.
Table 5. Result of regression, equation (3) fit to experimental results. Shear stress, resistivity and plowing force for hot-dipped, electroplated dull, and reflow tin.
Shear stress,
(MPa)
Resistivity,
(10-8 m)
Plowing force FP
(N)
Hot dipped tin 12 24.1 0.58
Electroplated tin 16 26.7 0,51
Reflow tin 12 26.3 0.25
Electroplated tin, with lubrication
8.8 26.7 0.03
The plowing force for the reflow tin is approximately half as high as for hot dipped and electroplated tin. The plowing force is mainly a function of the
29
plowing stress, which is the mean pressure opposing the sliding movement of the rider, and the cross-sectional area of the resulting groove 44. Hence, the low plow force obtained for reflow tin might be caused by a small cross-sectional area of the track, due to the thin residual tin layer. A thin lubricant film lowers the required insertion force, mainly due to reduced plowing force, but also due to reduced friction shear stress.
3.5 The effect of heat-treatment (Paper V) The results of heat-treatment experiments are presented in Table 6. The table shows a schematic view of the intermetallic compound growth based on the microscopy analysis.
Table 6. The formation and growth of the intermetallic compound 45. HD=Hot dipped tin, ELP=electroplated tin. Status Duration of heat-
treatment at 173 C.
Thickness ( m) (Cu3Sn)/Cu5Sn6
Surface roughnessRa ( m)
HD ELP HD ELP HD ELP
Nucleation
Island growth
Initial (0)/0-1 0.54
Secondary island growth
Initial 0.25h 0.25H
(0)/0.3-1(0)/1
(0)/0.5-1 0.33 0.35
0.63
Coalescence and formation of
secondary phase
1h 1h (0.3)/1 (0.3)/1.0 0.31 0.44
Further growth and coalescence
32h 32h (1-1.5)/ 1.5-2
(1)/2 0.54 0.63
30
Initially, the electroplated tin had approximately 50% of the interface covered by intermetallic compound, while the hot dipped tin had a tight intermetallic compound completely covering the interface. Besides increasing the thickness of the intermetallic compound, heat treatment causes changes to the surface roughness of the intermetallic compound. An optimum smooth intermetallic compound was achieved for a certain combination of time and temperature of heat-treatment for both hot-dipped and electroplated tin.
The effect of heat-treatment was studied by using the test bench described in section 3.1. When the rider slid on the hard intermetallic compound, i.e. Stage II, the contact resistance continuously increased to a point where it started to rise sharply to a relatively high value. The highest mean contact voltage drop for a stroke (half cycle) during Stage II versus duration of the heat-treatment for the hot dipped tin coating is depicted in Figure 14.
The majority of the samples achieved a contact voltage drop above the softening voltage of tin (70 mV). However, three samples heat-treated for 1 hour at 173 C showed an exceptionally low contact resistance, below 20 mV (arrow nr1). These three samples were heat-treated at the same time. Additional tests were conducted with new samples heat-treated at the same temperature and holding time (1 hour at 173 C), but these samples did not show such a low contact resistance. Microscopy studies showed that the intermetallic compound of the “low resistive” samples heat treated 1 hour at 173 C was very smooth, compared with the intermetallic compound of a non heat-treated sample. It is probably that a unique smooth intermetallic compound was formed for these samples.
Hence, a low contact resistance can be achieved despite the rider sliding on the intermetallic compound, provided that the intermetallic compound is smooth, and thus the amount of remaining tin is minimized.
The highest mean contact voltage drop for a stroke during Stage II versus duration of the heat-treatment for the electroplated tin coating is illustrated in Figure 15. All samples reached a contact voltage above 20 mV. For the electroplated samples, the rider usually breaks through the intermetallic compound before transition to a high contact resistance occurs, especially when the intermetallic compound is not thick and homogeneous. This result in a restricted increase in the contact resistance compared to hot dipped samples, especially for non heat-treated samples. The highest contact voltage of all samples is found for the 0.25-hour heat-treated sample (arrow no. 1). Thus, a short heat-treatment can increase the contact resistance during Stage II significantly.
31
Figure 14. The highest mean contact resistance for a stroke during Stage II versus duration of the heat-treatment for the hot-dipped tin coating. I=2A.
Figure 15. The highest mean contact resistance for a stroke during Stage II versus duration of the heat-treatment for the electroplated tin coating. I=2A.
3.6 The conditions for transition to a high contact resistance (Paper IV) When the rider slides on the intermetallic compound, the contact resistance continuously increases to a point (the transition voltage) where it starts rising sharply up to a high voltage (0.5-3.5V), see Figure 9. The condition for this transition to a high contact resistance was studied. Two different normal loads were used, 2 N and 10 N, as well as current loads in the range from 8 mA to 20 A DC. Most contacts achieved a transition voltage somewhere between 70 mV and 130 mV, which corresponds to the softening and the melting voltage of tin46. It should be noted that the true transition voltage could be somewhat higher than the measured transition voltage. Hence, the lowest threshold value for transitions to a high contact-voltage drop (0.5-3.5 V) seems to be equal to the softening voltage of tin (70 mV), independent of the current load.
A characteristic contact-voltage response during fretting test, similar to the voltage drops described in this work has also been reported by A. Lee and M. Mamrick 25.
32
3.7 Model for prediction of contact failure (Paper IV) When the free-tin layer has been removed, the rider slides on a layer consisting of a mixture of uneven hard intermetallic compound and remaining tin. During this stage, Stage IIa, the contact resistance increases continuously. This is because the surface oxide that is formed during sliding is not removed efficiently as in Stage I where it was formed on a thick layer of soft tin. With a theoretical model describing the deterioration process during Stage II, it would be possible to:
1. Better understand the deterioration processes. 2. Predict the length of time prior to failure when designing a contact
terminal. 3. Design a more reliable laboratory test for contacts.
To be able to formulate a theoretical model of conditions during Stage IIa, we need to simplify the deterioration process. The model presented in this thesis makes the following assumptions:
1. The contact area initially consists of a large number of conductive a-spots.
2. The continuous increase in the contact resistance during Stage IIa is caused by oxide formation on the surface, which reduces the size and number of the conductive a-spots.
3. The wear of the contact surfaces during Stage IIa is neglected. 4. The surface (asperity) temperature of the track is constant
throughout Stage IIa. 5. The oxidation of tin can be described by a parabolic oxidation-rate
equations according to P. Kofstad 47.
According to Greenwood 48 the contact resistance of a large number of small and equal spots distributed uniformly and densely over a cluster area can be expressed approximately as follows. Nomenclature
Electrical contact resistance RC
Diameter of cluster D
Number of contact spots n
Diameter of contact a-spot d
ndDRC
11(4)
33
During sliding, the number and size of the conductive contact spots decrease, mainly due to oxidation of the intermetallic compound and the remaining tin. Probably, it is oxidation of the remaining tin, which determines the rate of deterioration. The oxidation of the surface is a function of the temperature and the time that the surface is subject to heat. Moreover, the contact spot is oxide free during the entire Stage I until the start of Stage IIa, since the surface oxide is continuously removed during Stage I. Thus, a qualified assumption is that the formation of tin oxide during Stage IIa could be approximately written as:
oxtkx (5)
where t=0 is the point when the rider reaches the intermetallic compound, which by definition is the start of Stage IIa. x represents the amount of metal transformed into oxide divided by the amount of metal available for oxidation. The oxidation constant oxk has the unit [ s/1 ]. Furthermore, given the assumption that the reduction of ( nd ) is directly related to the amount of tin oxide present, then the deterioration of the contact spots can be written as:
)( oxtkdnnd 100 (6)
Inserting equation (6) into (4) and Ohms law ( IRU CC ) gives:
)( oxC tkdnD
IU1
11
00
(7)
We now have an equation describing the increase in the contact voltage as a function of the length of time (t) that the rider is sliding on the intermetallic compound. The equation is valid as long as
10 oxtk .The next step is to determine the oxidation constant kox, based on the
experimental results. The difference in resistivity between tin and the intermetallic compound is rather small. Therefore, we can assume the resistivity to be the same as for hot-dipped tin, which is estimated at approximately 25 x 10-8 m, see Table 5. D is equal to the width of the track, and ( 00dn ) is estimated from the measured value of Uc when t=0, denoted U0 (see equation 8). We do not need to know the specific value of nand d respectively.
34
IDuDIdn
000 (8)
We can now determine the constant kox from a curve of best fit. Figure 16shows an example of where equation (7) has been fitted to the experimental result. The agreement between the estimated value of the contact voltage and the measured result is good until the voltage reaches 25 mV. This is equal to the estimated increase in the contact temperature of 80 C. Probably, above 25 mV the tin starts to be soft, with resulting growth of the contact spots as discussed earlier.
The deterioration of the contact spots during Stage IIa can be estimated by equation (6), see Figure 17. After 5 number of cycles (in Stage II) the contact voltage drop has increased from initial 1.5mV to 2mV, which corresponds to a reduction of the conductive contact spots nd with 50%. An increase of the contact temperature with 3 C (i.e. 13mV49) corresponds to a reduction of the conductive contact spots nd with as much as 94%. Hence, the degradation of the conductive contact spots (in number and size) might be very severe before the contact voltage drop starts to rise significantly.
Figure 16. Proposed model compared with the experimental result. Model data: I = 2(A),
= 25x10-8( m), D = 0.75(mm), u0 =1.38(mV), kox = 0.0163(1/s), t = sliding length /sliding speed =10(mm) number of cycles /3.33(mm/s) 3 number of cycles.
Figure 17. Remaining of normalized nd versus number of cycles for the case in Figure 16. n = number of contact spots, d = diameter of contact a-spot.
35
From equation (7) we can observe
Cox
Utk 1
lim
If we assume that no self-healing occurs, then we can predict the time it takes to achieve a high contact resistance by calculating when 1oxtk .Thus, the time taken to reach an unstable contact while the rider slides on the intermetallic compound can be expressed by:
oxkt 1 (9)
Since the oxidation constant kox is the inverted value of the length of duration of Stage IIa, it is easy to determine the constant kox from laboratory test results. Evaluation of the conducted experiment shows that increased sliding speed significant increased the oxidation constant kox.
3.8 Vibration induced fretting (Paper VII)
3.8.1 Experimental set-up The fretting test bench utilized for simulating vibration induced fretting is shown in Figure 16. Vibrations cause externally induced forces acting on the contact and are a function of the acceleration and retardation of the moving mass, i.e. the fretting test should be force controlled. An important requirement when designing the test bench was that the center of the oscillation should not move during the test. Using dead-weights the normal load levels are adjustable from 50 g to 1 kg, while the current load levels are adjustable from 0.1 mA to 20 A. During testing, both the tangential force and the displacement of the moving fixture are recorded, as well as the contact voltage drop. The contact voltage drop is measured by the 4-wires method.
36
Figure 18. Schematic drawing of the fretting test bench.
3.8.2 Test conditions The following test conditions were used:
Contact geometry: ball-on-flat, radius of the ball 2 mm. Vibration frequency: 4 Hz. Normal loads: 100, 200 or 300 g. In the following text the normal loads are approximated to 1, 2 and 3 N. Current load: 2 A DC. Ambient relative humidity: 50 %. Ambient temperature: 21–22 C.
In order to simulate insertion, one displacement cycle with amplitude of 0.5 mm was performed for all experiments before the vibration started.
3.8.3 ResultThe displacement caused by the applied tangential force constitutes of two parts, one elastic part and one non-elastic. The elastic part is a function of the stiffness of the present test rig, while the non-elastic part is caused by partial slip or gross slip of the contact spot 15. Thus, the loop of the measured tangential force versus the relative displacement reveals the true slip. Evaluation of the slip shows that the initial steep increase in contact voltage drop was efficiently stopped as soon as the slip was decreased below 1 m.
A relatively low amplitude tangential force resulted in no slip and a very low and stable contact voltage drop throughout the test, see Figure 19. At a slightly higher tangential force amplitude (e.g. 1.95 N) the contact voltage drop initially increased until it reached a level where it stabilized. The higher the tangential force amplitude, the higher the stable plateau level. However,
37
at a distinct tangential force amplitude, the contact voltage drop became unstable and the contact failed (compare e.g. 1.97 N and 1.98 N).
0
2
4
6
8
10
12
0 100 200 300 400 500 600Time (s)
Con
tact
vol
tage
dro
p (m
V)
2 N
1.98 N
1.97 N 1.95 N 1.9 N 1.7 N 1.5 1.98 N1.97 N
1.95 N
1.90 N
1.70 N
1.50 N
1.00 N0.50 N
0
20
40
60
80
100
120
0 0.5 1 1.5 2 2.5
Amplitude of the tangential force (N)
Estim
ated
rem
aini
ng c
ondu
ctiv
e ar
ea a
fter 1
00 s
(%)
Figure 19. The contact voltage drop versus time for some different tangential force amplitudes. Normal load 2 N and 5
m thick tin coating. The maximum amplitude of the tangential force that can be applied without failure is 1.97 N.
Figure 20. Slip and standard deviation (STDV) of the contact voltage drop versus time. The STDV decrease to a relative very low value as soon as the non-elastic slip (peak to peak) is below 1 m. Normal load 2 N, tangential force amplitude 1.97 N and 5 m thick tin coating.
The estimated remaining conductive area after 100 s testing for the contacts in Figure 19 is presented in Figure 20. The 104 % remaining conductive area at the amplitude of 0.5 N is due to contact area growth during the first few fretting cycles. The figure shows that the remaining conductive area starts to decrease steeply at 1.7 N, and that the distinctive threshold value to failure at an amplitude of the tangential force at 1.97 N – 1.98 N (Figure 20) is the result of a continuous decrease of the remaining conductive area.
The instability of the contact voltage drop was evaluated by analyzing the standard deviation (STDV). The STDV for every 5th second was evaluated, based on 50 recordings (Figure 21. The transition to a stable contact condition is very distinctive: the STDV of the contact voltage drop decreases to a relatively low value as soon as the non-elastic slip (peak to peak) is below 1 m. The effect is due to the fact that slip inevitably results in changes of the contact conditions and this in turn results in an increased deviation of the measured contact resistance 50.
38
0
5
10
15
20
25
30
35
40
0 100 200 300 400 500 600Time (s)
Estim
ated
non
-ela
stic
sl
ip, p
eak
to p
eak
(m
)
-0.07
-0.05
-0.03
-0.01
0.01
0.03
0.05
0.07
Stan
dard
dev
iatio
n of
co
ntac
t vol
tage
dro
p (+
/- m
V)Transition to stable condition
Slip
STDV of thecontact voltage drop
0
Figure 21. Slip and standard deviation (STDV) of the contact voltage drop versus time. The STDV decrease to a relative very low value as soon as the non-elastic slip (peak to peak) is below 1 m. Normal load 2 N, tangential force amplitude 1.97 N and 5 m thick tin coating.
3.9 Properties of lubricated contact terminals (Paper VI) For normal loads below 1 N, the insertion force can be reduced by using an appropriate lubricant mainly due to a reduced plowing force. However, for normal loads over 1 N, the ratio between insertion force and contact resistance is close to that for unlubricated contacts. Furthermore, the contact resistance can be maintained low and stable under reciprocated sliding, provided a low viscosity lubricant is used. Probably, a low viscosity lubricant is needed in order to achieve a sufficient fast flow back to protect the exposed oxide-free surface from oxidation.
3.10 Properties of lubricated contacts with textured surfaceOne problem with using a thin film of a low viscosity lubricant is that the lubricant might be removed as a result of sliding and/or migration out from the contact area during long-term operation. Since contact terminals are not relubricated this will eventually result in non-lubricated contact spots, with subsequent risk of fretting corrosion and failure. It was proposed to use a textured surface in order to improve the supply of lubricant into the contact spot when a very small amount of lubricant remains on the contact surface (a
39
patent approved)51. Thus, a textured surface should then extend the time when a low viscosity lubricant will protect the contact spot from oxidation.
Most steel sheets for the press forming industry have a modulated surface structure, in order to ensure an effective supply of lubricant between the sheet and the closed tool during forming. The design of the pattern including the technique to produce the rolls has almost become a science of its own. As for electrical contacts N.P. Suh has suggested a micro-modulated surface structure to trap wear particles formed during fretting52. Thus, the oxidized wear debris should be collected in tracks without doing any harm.
Experiments with reciprocating sliding of 5 mm were conducted to study the effect of a textured surface in combination with lubrication. The experimental set-up has been described in section 3.1. The contact geometry was ball on flat and only the flat sample was textured. The lubricant was applied only on the flat sample.
The following test materials were used:
Rider (hemispherical with a radius of 2 mm), CuSn4 alloy (C511) with electroplated matte tin, coating thickness 5 m. Flat normal and flat textured sample: C7026 alloy (CuNi2Si0.4) temper TM02, with electroplated matte-tin over flash nickel, coating thickness 1.5 m and 0.15 m respectively. Lubricant: Linear perfluorinated polyethers ( PFPE ) with OH end groups. Viscosity 81 cSt at 20 C .
The textured surface was produced by using Laser Beam Textured (LBT) rollers. A Scanning Electron Microscopy (SEM) image of the textured surface is presented in Figure 22.
Figure 22. SEM image of the textured surface. Maximum depths of the cavities are 4 m.
40
Experiments with a normal contact load of 2 N and with 0.5 g/m2 lubricant were performed with both a normal flat surface and a textured surface. The results were compared with experiments using well-lubricated (20 g/m2)samples with a flat surface, see Figure 23.
00.5
11.5
22.5
33.5
44.5
5
0 0.5 1 1.5 2 2.5Drawing force ( N )
Con
tact
vol
tage
dro
p ( m
V ) Normal el. tin ( 0.5 g lubricant/m2 )
Textured el. tin ( 0.5 g lubricant/m2 )
Normal el. tin, well lubricatedcondition ( 20 g lubricant/m2 )
Figure 23. Contact voltage drop versus drawing force, mean value during 5 insertion withdrawal strokes. Experiments with well-lubricated surface (20 g/m2) normal flat surfaces were made at varying contact load. Experiments with 0.5 g/m2 lubricant on normal flat and textured surface were made at 2N contact load.
The textured surface with 0.5 g/m2 lubricant showed the same performance as a well-lubricated surface (20 g/m2), while a normal flat surface and with 0.5 g/m2 lubricant showed an increased contact voltage drop. Moreover, in the present experiments a textured surface and 0.5 g/m2 lubricant was as efficient as a normal flat surface and 20 g/m2 lubricant. Thus, a textured surface of the flat part improves the supply of the lubricant to the contact area when a very small amount of lubricant remains on the contact surface.
3.11 Development of a tin alloy coating with improved fretting properties (Paper VIII) A novel method to improve the performance of pre-tinned strip material for connectors has been studied. The idea is to increase the possibility of achieving a cold-welded contact spot by adding a dope additive, and thereby, to achieve an ultimately stable connection. The invention has resulted in three approved patents 53, 54, 55.
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A cold-welded spot will restrict detrimental micro sliding and wear during vibration, and also prevent the ingress of air/moisture, which could cause corrosion, i.e. oxidation, of the contact interface56. Provided that the cold-welded contact spot can be maintained the contact will be low and stable during long-term operation. To achieve a cold-welded contact spot, the oxide layer must be disrupted and a high contact pressure be applied57. A high contact force and a thick and soft coating will, in general, promote a cold-welded contact spot, since the tin oxide film is removed during insertion. However, the trend today is towards thinner tin coatings and lower contact force and thereby the possibility to remove the oxide film during insertion is restricted.
The rate of oxidation of tin can be significantly changed by alloying small amounts of elements, according to a study performed by Boggs et al58. If an alloying element forms an oxide that is more thermodynamically stable than SnO, the alloying element is preferentially oxidized on the surface, and the formation of tin oxide is restricted. Tin with 0.1 atomic percent (abbreviated as “at. %”) zinc results in preferential oxidation of the zinc. Similar reduction in oxidation rate and inhibition of tin oxide formation were observed with tin-indium, tin-phosphorus, and tin-germanium 58.
In the present work it was shown that adding small amounts (0.1 at. %) of one of the following elements: zinc, nickel, titanium, or phosphorus, to the tin improved the possibility of achieving a cold-welded contact spot, while small amounts of indium or chromium had the opposite effect. Also 2.68 at. % (equal to 5 wt. %) silver added to the tin increased the possibility of achieving a cold-welded contact spot. Moreover, phosphorus added to any of the tested alloys (SnCr, SnAg, SnZn, and SnNi) increased the possibility of achieving a cold-welded contact spot.
A high interfacial pressure will further enhance the possibility of achieving a cold-welded contact spot. In this work, a high interfacial pressure was achieved by using a thick intermetallic compound with a thin top layer of tin. This, in combination with a preferential oxidation additive, resulted in a significantly increased possibility of achieving a cold-welded contact spot, which in turn resulted in a significantly increased electrical stability during vibration.
Finally, phosphorus is recommended as a dope-additive to tin, since it improves the ability to form a cold-weld contact spot without significantly reducing the good anti-corrosion properties of tin.
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4 Summary of contribution
The design of a connector for a specific application is, to a high degree, a matter of optimization of a number of contradictory parameters, and to be able to optimize the design it is important to know the relevant material properties. In this thesis, the friction, wear and electrical properties of tin-coated systems have been described, as well as the test and analysis methods to determine the same. From a practical point of view, the most important findings in this thesis are the following:
The electrical performance of a tin-coated contact should be discussed in terms of contact voltage drop and related to the softening voltage of tin (70mV), instead of the contact resistance. This is due to an instantaneous transition to high contact voltage occurs at a contact voltage drop corresponding to the softening voltage of tin (70mV), independent of current load (0-20A). The distinct critical threshold value of the amplitude of the tangential force, when the contact will inevitably fail, is proposed to be a design and specification parameter. The tangential force acting on the contact during vibrations is critical, as it determines whether slip will occur or not. Evaluation of the standard deviation of the contact resistance is proposed as a simple method to reveal slip during operation. For an unlubricated contact it is important to avoid almost any reciprocated slip, since it will cause fretting, and thus, a significant risk for contact failure. The insight of the deterioration of a tested contact can, in general, be significantly improved by studying the reduction of the conductive area. The relative reduction of the conductive area can be estimated by simply dividing the square of the initial contact voltage drop by the square of the present contact voltage drop.
Moreover, in order to improve the performance of pre-tinned material, the following are proposed:
When using a contact lubricant a textured surface could be used, in order to improve the supply of lubricant into the contact spot. To add an preferential oxidation additive to the tin, in order to increase the ability to form a cold-welded contact spot, and thereby achieve an
43
ultimately stable connection. Phosphorus is recommended as a dope-additive to tin, since it improves the ability to form a cold-weld contact spot without significantly reducing the good anti-corrosion properties of tin.
44
5 Sammanfattning (Summary in Swedish)
Mängden elektronik i fordon ökar snabbt och så också antalet fel i elektroniska system i nya bilar. Fältstudier har visat att 30 – 60 % av felen är relaterade till problem med elektriska kontakter. Syftet med detta arbete har varit följande: att karakterisera egenskaperna hos tennbeläggningar för kontakter, att få en ökad förståelse av nedbrytningen av en kontakt orsakad av nötning och fretting (vibrationsinducerad korrosion), samt att föreslå förbättringar av tennkontakter för fordon.
Alla plugg-in kontakter har det gemensamt att det är två mekaniska delar som trycks an mot varandra med en viss kontaktkraft och elektrisk ström ska kunna ledas över förbindelsen. Hur väl strömmen kan passera bestäms till stor del av hur många och hur stora kontaktpunkter som kan leda ström, och om det finns en isolerande oxid på ytan. Tenn är både mjukt och korrosionsbeständigt och en tennbelagd kontakt har därför som regel en mycket god ledningsförmåga när den är helt ny. Detta tillsammans med ett relativt lågt pris gör att tennkontakter används i stor utsträckning i fordon. För att förstå svårigheten med att göra en tillförlitlig kontakt som ska placeras i ett fordon behövs en insikt i hur en kontakt fungerar i praktiken.
Den kraft som krävs för att sticka i en kontakt har en avgörande betydelse för utformningen och kostnaden för kontakten. En låg instickskraft kan erhållas genom en låg kontaktkraft och/eller en tunn tenn beläggning, men detta leder i sin tur till en ökad risk för fretting. Bortnötning av ytoxiden på kontakten under insticket är nödvändigt för att skrapa bort det icke ledande oxidskiktet. En tun tenn beläggning kan dock nötas bort efter bara några få instick och utdrag, och kontaktsiftet kommer då att glida på en blandning av kvarvarande tenn och den hårda intermetalliska fasen som alltid bildas i gränsytan mellan substrat och tenn. Den kraftigt reducerade nötningen kommer att leda till att mer oxid formas vid glidning än vad som nöts bort, och därmed kommer kontaktresistansen att öka.
Nedbrytningen av de ledande kontaktpunkterna i en kontakt på grund av korrosion eller fretting är vanligtvis en kontinuerlig process som pågår under lång tid utan att funktionen hos kontakten synbart försämras. En elektronikkontakt kan fungera väl och passera tester trots att 95 - 99 % av den initialt ledande kontaktytan är förstörd. När i stort sett all ledande kontaktyta är borta uppstår det tillfälliga avbrott och till sist havererar kontakten. En havererad kontakt kan dock åter bli ledande då oxiden kan
45
brytas upp av till exempel mekaniska vibrationer. Det kan därför vara väldigt svårt att identifiera en felaktig kontakt som har orsakat ett tillfälligt avbrott. Ett vanligt sätt att ”reparera” en kontakt är att göra ett antal instick/utdrag för att bryta upp oxiden, men risken är stor att kontakten havererar igen. Således, även om ett kontakthaveri oftast uppstår plötsligt och oväntat så har det oftast föregåtts av en kontinuerlig nedbrytning under lång tid.
Vibrationer är ett stort problem för tennkontakter, och vibrationer och mekanisk rörelser till exempel på grund av temperaturväxlingar är vanligt förekommande i fordon. För varje kontakt finns det ett tröskelvärde på vibrationsnivån när mikroglidning mellan kontaktytorna börjar ske. Under denna nivå är kontakten stabil, men om vibrationerna överstiger tröskelnivån kommer kontakten förr eller senare att få en mycket hög kontaktresistans. I detta arbete visas att mikroglidning i en kontakt orsakar en förhöjd brusnivå på spänningsfallet över kontakten vilket kan detekteras med konventionella instrument. Således kan man relativt enkelt detektera mikroglidning i en kontakt och metoden skulle kunna vara lämplig för att verifiera att en kontakt inte kommer att drabbas av nötningsoxidation under drift.
Smörjmedel används ibland i tennkontakter för att förhindra fretting. För detta krävs ett smörjmedel låg viskositet. Den förmodade huvudorsaken är att vid mikrorörelser mellan kontaktytorna kommer ren metallyta bli exponerad mot luft, och med ett lågvisköst smörjmedel kommer smörjmedlet att snabbt täcka upp och skydda den exponerade oxidfria ytan från oxidation. En tunn lågviskös smörjfilm riskerar dock att försvinna antingen på grund av mekanisk avstrykning eller att smörjmedlet kryper iväg (migrerar) under drift. För att förbättra tillförseln av smörjmedel till kontaktpunkten provades en texturerad kontaktyta i kombination med smörjmedel. Texturering av karosseriplåt är väl etablerad inom fordonsindustrin och samma teknik användes i detta arbete. Arbetet visar att en mer effektiv tillförsel av smörjmedel till kontaktpunkten kan fås med en texturerad yta jämfört med en slät yta vid de små mängder smörjmedel som vanligtvis används i kontakter. Ett patent har erhållits inom detta område.
Inom detta arbete har dessutom utvecklats en tenn legering med förbättrat motstånd mot fretting. Genom att tillsätta mycket små mängder av additiv till tenn så kan man förändra oxidegenskaperna och därmed öka möjligheterna att erhålla en kallsvets vid insticket av kontakten, vilket i sin tur ger en ytterst stabil kontakt. Vid inte alltför höga nivåer på vibrationsnivån kommer kallsvetsen dessutom att bli förstärkt. Totalt har tre patent erhållits inom denna del av arbetet.
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Acknowledgement
This work has mainly been done at the Corrosion and Metals Researchinstitute, KIMAB, (former SIMR) and the Materials Science Division, The Ångström Laboratory, Uppsala University, Sweden. All along the work there has been a close cooperation with Outokumpu Copper. The work has been funded by Outokumpu Copper, Vinnova, KIMAB, ABB Corporate Research, AB Volvo Technological Development Metals Laboratory, and Ford Motor Company, which are hereby gratefully acknowledged.
I would like to thank all colleagues and co-workers, who have been involved in this work. Particularly, I would like to express my deep gratitude to the following friends and colleagues:
My supervisor at Uppsala University, Dr. Åsa Kassman Rudolphi, for invaluable discussions and encouragement during ten years, and also for carefully reading the manuscript to this thesis.
Prof. Sture Hogmark at Uppsala University, for giving me the opportunity to do this thesis at the Materials Science Division, and for interesting art discussions over lunches.
Prof. Staffan Jacobson at Uppsala University, for stimulating discussions and for linguistic advices.
Rolf Sundberg, at Outokumpu Fabrication Technologies AB (OFTA), R&D and my former supervisor at KIMAB, Erik Schedin, for initiating the research on electrical contacts at KIMAB.
Anders Kamf at Outokumpu Copper, for continuous encouraging support and for keeping me informed about the latest news in the field.
Larz Ignberg at OFTA, R&D for supporting fundamental research during many years.
Per Lundström, at OFTA, R&D, for interesting discussions and for organizing some of the experimental work performed at Outokumpu Copper.
Ulrik Palmqvist, at KIMAB and OFTA, R&D for stimulating collaboration.
Anders Mårtensson at the library of KIMAB, for all help to find old books and papers all over the world, and also for continuously informing me about new papers he sensed that I might be interested in.
Margareta Nylén at KIMAB, for all discussions about tin and tin platings. Joakim Lindblom at JL Integration AB, for writing the computer
programs to the fretting test bench.
47
Christer Eggertson, Olle Läth and Anders Krusell at the work shop atKIMAB, for always being helpful with manufacturing test equipments
and test samples. Finally, I would also like to thank family and friends for all continuous
support, and especially, I would like to thank my dear uncle Dr. Shawky Hammam and my brother Tarik Hammam for all encouragement during the course of this thesis work.
48
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