q.7, ~ ~ ' 1~ . ,4

13ETURSDtA, MD 201034

U r LOTS OF MICROSTRUCTURtE, CON-11OSITION, AND

STRENGTH ON THE NIL DUCTILITY TRANSITION

(NDT) TEMPERATURE OF HY-80 STEEL

-' I byMarcel L. &,live

i

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D D C21Y

STIOTUES fl}MAR 1I 0 2EaSTRCTUESDEPArITfMENT

RESEARCH. APM' DEVELOPMENT REPORT

j;n'172~ Best Availabe C py Report 3722

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'..I! ', AC I IV I I fCt'wp-ator .~ p ,,!. REPORI SECURITY CLASSIFICATION

Naval Ship Research a-id Development Center UNCLASSIFlIED.j lthcsda, Maryland 2003412.GRU

EI:FE("IS OF MICROSTRUCTURE, COMPOSITION, AND STRENGTH ON THE NIL DUCTILITY TRANSITION (NDT)

TEMPERATURE OF IIY-80 STEELC4 " 0 NP I IV C NO T E (Tyrpr of report and irtcluv Ivr dotes)

Final Report( A tON4I S~rt ,.re,,iddle nla '.t namr)

Marcel L. Salive

*4tIPORI E()*tL 70. TOTAL NO, OF PAE Db N. OF REFS

iJmnuarv 1972 131 5Oa (ONINAC I OR GRAN't r,0 go. ORIGINATOR*S REP0ONT NUMniER(S)

h PRojrC T No SF 35.422.212 32

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Naval Ship Systems Command (NAVSIIIPS)

Steel from 22 heals of low-carhon Ni-Cr-Mo steel. IIY.80 (ASTM A543-65, and MIL-S-16216G (Ships)) was hecattrr'ete to study thie tftson the drop weight nil ductility transition (NDT) temperature of (1) commercial variationin comnposition and inclusion Content, (2) varIation in nuacrostructctire suich as prior austenitic grain size and the relativeamiount of isothermally prodlucedl ferrite or hbrinite In a tempered martensitic matrix, and (3) the observed variation in

strength obhtained after a owie-hour 1150 1: temper followed by water quench to prev'ent embr.t ~temnent 'lecoolingIrot), the temnpe ring te nipe rat tire rc

Using a tempered l00 percent martensitic microstrutc tutre as a base line, the variles founld to be most significant

were linearly related to the measuiredl Nil D~uctility Transit ion (NDT) teipmr lure, The range of the variable and iis

ceffect onl thle N DT t e nipe rat tire are as fol lows:

Variable Ra nge Effect on NDT

lPro'llocer (Coded I or 2) - 21 9 F.Prtior Aotenlic (;ralo Sie (3 to 9.5) - 19.2 F/ASTM C. S. No.Pc emit flairvi" (0 to 76 percent) 0.62 F/Percent IlainitePercent Fecrrite (0 to 69 pe rcen t) - 0.23 F/Plercent Ferrite

Percer t Pearli te (0) to 32 percent) 5.4 F/Percent Pe"atliteyield Strength1 (61.5 to 124.7 ksl) 0i.69 Ff/ksi

Iligherf order terms ;in . teractiori terms did not inicatyImprove the relationship. nor did consideration ofim.

cluton content of trc Otcel, Other thtan hydrogeni, no alloylitg (it trace elements were found to correlate with the

meisni'at NDT tempc rat tire (if this %teel.

DD 0#V 11 7 UNCLASSI FIED

_________________________L Hr W T O.F W~ RO L C W T

Nril Dth~n "iiv fansition'fnprti

Nit-O\to Sik-Id11)'-80i Steel

N'. i~t I It tire

I 'ar lit C

(rain Sue

femidei Yield

DD ~ 473-4.1473UNCLASSIFIED

DEPARTMENT OF THE NAVYNAVAL SHIP RESEARCH AND DEVELOPMENT CENTER'

BETHESDA, MD 20034

EFFECTS OF MICROSTRUCTURE, COMPOSITION, ANDSTRENGTH ON THE NIL DUCTILITY TRANSITION

(NDT) TEMPERATURE OF HY-80 STEEL

by

Marcel L. Salive

APPIROVEj[) FOR PUjBLIC( Mi LEASI3:DISTRIBUTION UNIIMITE'I)

January 1972 Report 3722

TABLE OF CONTENTS

Page

ABSTRACT. ....... I......................

ADMINISTRATIVE INFORMATION. . ........ ........

INTRODUCTION. ............................. 2

LITERATURE SURVEY. .. ............ ............... 2

GRAIN SIZE EFFECT .. ........................... 3

Ferritic Grain Size. ............................ 5

Austenitic Grain Size. ........................... 6

Microstructur . ................................ 8

Fracture Path .. ............................ 8

MICROSTRUCTURE. .................... .......... ..... 9

COMPOSITION .. ...................... ... ....... 10

STRENGTH. ...................... .. ........ 10

EXPERIMENTAL PROCEDURES . .................... .. . . ...

SUMMARY........................... . .. .. . ... .. .. . . ...

MATERIALS .. ....................... .. .. .. . ..

I IEAT TRE ATING PROCEDURES. ....................... 12

Fully Quaenched (Martensitic) .. ............ .. .............. 12

875 F Isothermal Treatment (Scmi-Bainitic) .. .............. .. .... 13

1200 F Isothermal Tre.-tnicnt (Serni-Ferritic). .. ... .. ............. 13

TESTING PROCE-DURES........ ......... ........... 13

Drop Weight NDT Tests .. ............... ........ 13

Tensile Tests. .................. .............. 14

Compression Tests .. .......................... 14

Ch.irpy Impact Tcsts .. ...... ....... ...... ....... 15

NflTALLOG RAPI IIC PROCEDURES. ... .. .................. 15

EXPERIMENTAL RE-'SULTS. .......................... 16

CIIEMICAL COM1IOSITION .. ........ .. ................ 16

METAI LOGRAPtIIC ANALYSIS. ............. ............. 17

MECIhANICAL PROPERTY TEST RESULTS........................7

DISCUSSION. .. ............................. 18

IU~~i~I~tNCEO' VARIAULLS.... .. .. .. .. ............... )

SIN'GULARLY SIGNIFICANT VARIABLES AFFECTING; NDT. .. ........... 20

RE-SULTS O1: R;.';RFSSION ANALYSIS ON NDT. ................... 21

(oniposition anI I'roduter ............. .... .......... 22

PagePrior Austenitic Grain Size I .. . . . . .22MicrOStructure......... . . . . . . . . . .2

Inclusion Content. ........................... 240trength........................ .... .... 24Summary of Factors Controlling NDT Temperature........... ...... 25

CONCLUSIONS .. ............................. 28FACTORS AFFECTING THE NDT TEMPERATURE ... ............... 28

ACKNOWLEDGMENT. .. ...................... .... 30

APPENDIX A - REAGENTS USED TO DEVELOP PRIOR AUSTENITICGRAIN BOUNDARIES. .. .....................

APPENDIX B - ETCHING TECHNIQUES FOR IDENTIFYINGISOTHERMAL PRODUCTS....... ...... ....... 55

APPEiN DIX C - 1PHOTOMICROGRAPIIS OF MICROSTRUCTURES...... ..... ... 59

APPENDIX D - CIJARPY V-NOTCH IMPACT PROPERTY CURVES. .............. 83

APPENDIX E - TIME-TEMPERATURE COOLING CURVES FOR DROP WEIGHT.SPECIMENS WH-EN TRANSFERRED FROM 1640 F FURNACETO MOLTEN SALT BATHS AT EITHER 1200 F OR 875 F. ......... 115

REFERENCES. .......................... ...... 119

LIST OF FIGURES

Figuire I -Lo'ation of Material Used in This Study .... .......... ..... 31

Figure 2 - Layout of Drop Weight Test Specimens. ........... .. ... ..... 31

Figure 3 -Location of Mechanical Property an(I Transvecrse Charpy V-Notch ImpactSpecimens Made from 'rested Drop Weight Specimens. ...... ...... 32

Figure 4 - Location of Longitudinal Charpy V'-Notchvy i~pc pcmn ade

from Special Specimen Blocks . . .. . . . . . . . . . 32

Figure 5 -- Pnotomicrogri phs of the hIcusion Con tent of Coupon X- 18 ...... I......33

Figure 6 - Photomicrogra phs of the Inclusion Con tent or Coupon Y- I............34

Figure 7 - A Plo~ t othe iftference between Calculated-ind Experimiental NDTTcmp.-rature of Material Conitaining from a Trace to 69 Perc. ntFcrrito verstis F:rrite Content . . . . . . . . . . . . . . . . . . . . .3

Page

Figure 8 -- Plots Showing the Relative Effect of Prior Austenitic Grain Size, Producer,Percent Nonmiartensitic Isothermal Transformation Produce (Bainite,Ferrite and Pearlite), and Tensile Yield Strength on the NOTTemperature of a Low-Carbon Ni-Cr-Mo Alloy SteelTempered at IO1 . ....................... 35

APPENDIXES

Figure ClI - Microstructutre (I OOOX) of Specimens from Cropped End Number MiAustenitized at 1640 F, Isothermally Treated at 875 F and 1200 F,and Then Tempered at 1 FS F.............. ........... 60

Figure C2 - Mi crost ruct tre (I 000X) of Specimens from Cropped End Number X-2Austenitized at 1640 F, Isothermally Treated at 875 F and 1200 r ,and Then Tempered at 1150 F... ........ ........... 61

Figure C3 - Miciostructure (IQOOX) of Specimens from Cropped Ends X-3 and X-4Auistenitized at 1640 F, lsothcrmafly Treated at 875 F and 1200 F,and Then Trempered at 1150 F. ......... ............. 62

F-igure C4 - Microstructure (I OOOX) of Specimens from Cropped Ends X.5 and1 X-6Austenitized at 1640 F, Isothermally Treated at 875 F and 1200 F,and Then Tempered at 1150 F ....................... 3

Figure C5 - Microstructure (I OOOX) of Specimens from Cropped Ends X-8 and X-9Austenitized at 1640 F, Isothermally Treated -,t 875 F and 12-00 F,and Th'len Tempered at 1150 F .. ..................... 64

Firt C06 Microstrucluiz (10GOOX) of Specimens from Cropped Fnds X-10 and X-1 IIAustenitizecl at 1640 1-, Isothermally Treated at 875 F and 1200 F,

an Thin Tempered at I1ISO F.. ..................... 65

Figure C7 Microstructure (I OOOX) of Specimens from Cropped End Number X-1 2Austenitized at 1 640 1-, Isothermally *rreated at 875 F and 1200 F,and Then Tempered at H150 F ...................... 66

Fipure C8 Effects of' Prior Auistenitic Grain Size on the Microstructure (.1 OQOX)of Croppedl Iinds X-1 2 and X-I 3. ..................... 67

F~igure C9 - Microst ructutre (I0OOX) (,,,Specimens from Cropped End Number X-13Austenitized at 1640 F, Isothermally Treatod at 875 F an(! 1200 F,

and Ten Temperedl at 11I F......................6

Figure CIO 0 Ef'fects of Prior Auistenitic Grain Size on the Microstructutre (I OQOX)Of Cropped Ends X. 15 and1 X- 18. .................... 9

Figure C I - Microsttuctmer (I QOOX) of Specimens from Cropped Enod Number XI 5SAusteni ti'icd at 10,10)17, Isothermally Trea tedl at 8475 F and 1200 F,.andI Then 'Tpcred ait I15 P F.. ................... ..... 70

iv

PageFigure CI 2 - Microstructure (I OOOX) of Specimens from Cropped End Number X-16.

Austenitized at 1640 F, Isothermally Treated at 875 F and 1200 F,,nd Thcn, Tempered at 1150 F ................ ...... 71

Figure C 3 - Effects of Prior Austenitic Grain Size on the Microstructure (I OOOX)of Cropped 1, r ds X- 16 and X- 17. ....................... 72

Figure C14 - Microst ructutre (I OOOX) of Specimens from Cropped End NumtC'r X-I 7Austertitized at 1640 F, Isothermally Treated at 875 F and 12,'O F,and Then Tempered at 1150 F ............ ............. 73

Figure C15 - Microst ruct tre (IOOOX) of Specimens from Cropped End Number X-18rAustenitized at 1640 F, Isothermally Treated at 875 F and 1200 F,and Then Temnered at 11.50 F ~ . . . . .. . ... ..... 74

Figure C16 - Microstructure (IOOOX) of Specimens from Cropped Ends Y-1 and Y-3Austenitized at 1640 F, Isothermally Treated at 875 F and 1200 F,and Then Tempered at H150 F. ......... I.............75

Figure C1 7 -Microst ructutre (I OQOX) of Specimens fiom Cropped End Number Y-2Austenitized at 1640A F, Isothermally Treated at 875 F and 1200 F,and Then Tempered at 1150 F............. ........... 76

Iigmirc CI9 - Mic,,ostriictrure IOQOX) of Specimens from Cropped End Numbcr Y-4Austenitized at 1640 F, Isothermally Treated at 875 F and 1200 F,and Then, Teimpered at 1150 F ............... .. ......... 77

Figtire C19- Microstructure (I GOOX) of Specimc ns from Cropped Ends Y-5 and Y-6Austenitized at 1640 F, Isothermally 'heated at 875 F and 1200 F,and Thcn Tempered at I11S0 F. ..................... 78

Figure (72() Microstructver (I OOOX) of Specimens from Crepped Ends Y-7 and Y-8Austenitized at 1640 F, Is:ntherniaIly Treated at 875 F and 1200 F,and Then Tempered at I I 5G F ...................... 79

F~igure C21 -Microstructutre (I OOOX) of Specimens from Cropped Erio. 1'-9 and Y-10.Austenitized at 1640 F, Isothermally Treated at 875 F and 1200 F,arid Then Tempered at 1150O . .................. .. .... 80

Figure C22- Microstructure (I1000X) of Specimens from Cropped End Number Y-1 IAustenitized at I1640 F, Isothermally Treated at 875 F and 1200 F,anid Then Tempered at H150 F.. ..................... 81I

Figure D)1 Transverse and Longit udinal Cliarpy \'-Nolch Impact Data forSpecimecns Austenitized at 1 64() F for 1/2 Ihour, WaterQue~wchedl andI T1emnpered at H150 for I 1 fouir. ................ 84

Figure D)2 Tranisverse Charpy V-Notch I mpaict Data for Specimens Auistenitizedit I (AO F for 1/2 Hour, IsoitierniaIl v' Treated at 87.5 Ffor 152Sef-,inds, Water Qulenchcd and Tempered at 1150 Ffor I Hot .,... .. ......... ................ 92

PageFigure D3 - Transverse and Longltudintl Charpy V-Notch Impact Data for

Specimens Austenitized at 1'640 F for 1/2 flour, IsothermallyTheated at 875 fifor 1600 Seconds, Water Quenched andTempered at 1150O F for I flour. .................... .. .. 98

Figure D4 - Transverse Cha-py V-Notch Impact Data for Specimens Austenitizedat 1 640 F for 1/2 Hour, Isotherinal~y Treated at 1200 Ffor33.50 Seconds, Water Quenched and Tempered at 1150 Ffor I Hour ..... . ................ .. .. ... 102

Figure D5 - Transverse Charpy V-Notch Impact Data for Specimens Austenitizedat 1(-40 F for 1/2 Hour, Isothermally Treated at 1200 Ffor8500 Seconds, Water Quenched and Tempered at 1150 Ffor I flour. ..... ...................... 108

Figure P,6 - Transve7se Charpy V-Notch Impact Data for Specimens Austeitzedat 2WJ() I"for I Hlour, Transferred to 1640 F for 5 MinutesMinimum, Water Quenched and Tempered at 1 100 Ffor I Hiour. ...................... .. .... .... .. .. . . ..

Figure D7 Transverce Charpy V-Notch Impact Data for Specimens Austenirizedat 2000 F for I Iflour, Transferred to 1 640 F for 5 MinutesMini mumi, Isothermrally Treated at 875 P'for 1600 Seco.nds,Water Quenchied and Tempered at 1150 F for I Hlour. ............. 113

Figure El -Comparison of Actual Cooling Rates with Computed Cooling Ritesfor the Center of a 5/8 x 2 x 5 Inch Drop Weight Specimen. ........... 117

LIST OF TABLES

Table I - Producer's Reported Chemical Comporition and Mechanicil Properties .. .. .... 36

Table 2 - Actual Chcemical Composition of Coupons. .................. 38

Table 3 Inclusion Content .. ...................... 39

Tahle 4 Prior Austenitic G;rain Size. ....................... 40

Table 5 --Percen t Microconstituents Produced by Various Isothermal Treatmients .. ...... 41

Table 6 -- Results of Mechanical Property 'ests. ...... ...... .......... 42

Table 7 -- List of Variables and Correlioni Coefficients Showing I IighlyCorrelated Varibdles ...... ...... ....... ...... .... 4

'lable 8 - Results of I meai Regression Rela ting N I)T i'cmpe rat ure toPrior Austenitic Grain Size. ..... ...... ....... ..........

vi

PageTable 1) - Results of Linear Regression Relating NDT Temperature to

Prior Austenitic Grain Size, Percent Batnite, PercentFerrite and Percent Pearlite... . . . ........... ... . . ..

Table !0 - Results of linear Regression Relating NDT Temperature to Producer,Prior Austenitic Grain Size, Percent Bainite, Percent Ferrite,PerL, nt Pearlite and Tensile Yield Strength . . . . . . . . . . . . . . . •. 52

APPENDIX

Table El - Listing of a Coaputer Program Using the Instantaneous Film Coefficient Methodof Sinnotl and Shyne4 7 to Predict the Instantaneous Cooling Rate and Time-Temperature Cooling Curve for a Low-Carbon Ni-Cr-Mo Alloy SteelSpecimen 5/8 x 2 x 5 Inches When Transferred. from a NeutralSalt Bath at 16(40 F to a Neutral Salt Bath at 875 F . ....... 118

vii

ABSTRACT

Steel from 22 heats of low-carbon Ni-Cr-Mo steel, ItY-80 (ASTM A543-65, andMIL.S- I6216G (Ships)) was heat treated to study the eff~ct. on the drop weight nilductility transition (NDT) temperature of (1) commercial variation in compositionand inclusion content, (2) variation in microstructure such as prior austenitic grain sizeand the relative amouni of isothermally produced ferrite or bainite in a tempered marten-s ritic rPlatrix, and (3) the observed variation in strength obtained after a one-hour 1150 Ftemper followed by a water quench to prevent embrittlement while cooling from thetempering temperature.

Using a termpered 100 percent martensitic microstructure as a base line, th~vari-ables found to be most significant were linearly related to the measured Nil DuctilityTransition (NDT) temperature. The range of the variable and its effect on the NDTtemperature are as follows:

Variable Range Effect on NDT

Produc- " (Coded I 'r 2) -21.9 F

Prior Austenitic (3 to 9.5) -19.2 F/ASTMGrain Size G. S. No.

Percent Bainite (0 to 76 percent) 0.62 F/PercentBainite

Percent Ferrite (0 to 69 percent) - 0.23 F/PercentFerrite

Percent Pearlite (0 to 32 percent) 5.4 F/PercentPearlite

Tensile Yield (61.5 to 124.7 0.68 F/ksiStrength ksi)

Iligher order terms and interaction teems did not significantly improve the relationship,nor did consideration of inclusion content of the steel. Other than hydrogen, no alloy-ing or trace elements were found to correlate with the measured NDT temperature of

this steel.

ADMINISTRATIVE INFORMATION

This work was performed under the sponsorship of the Naval Ship Systems Command (NAVSIIIPS).as. "rt ef ResenrcLi )d ,,velopment Projec. SF 35.422.212 Program Element 62512N.

INTRODUCTION

At present there is a gap in the body ef general metallurgical kncwledge concerning the mechanisms that

control the notch toughness, as measured by the drop weight nil ductility transition (NDT)* temperature of the

low carbon, medium alloy, and quencued and tempered steels. Most prior studies have been limited to the pro-

cess of characterizing spt 'ific classes of commercial steels by their expected NDT temperature frequency distri-

bution range. No previous studies have been made of the intetaction of chemical composition, mechanical prop-

erties and microstructurf with the NDT temperature of any steel.

Studies of the notch toughness of higher alloy steels, such as those recently performed by U. S. Steel, have

concerned themselves with the fracture mcchat ,-s or critical flaw size approach that is more appropriate for such

higher strength. low toughness materials. Previous studies of the higher toughness steels, however, have aiways

conceraed themselvez with the effect of variation in a single alloying element and usually only with its effect on

strength or on sonic Charpy impac: property. In addition, these previous Atudies have been concerned only with

the effects on a normalized steel or on an oil quenched alloy steel of undetermined microstructure.

The normalized structural steels are gradually being replaced in critical, high toughness, low transition

tcirperature applications by the medium alloy steels such as IIY-80 and HY-100. Therefore, it is important that

the mechanisms controlling the interaction of composition, mechanical properties, and microstructure with NDT

temperature be determined. This is so, particularly since the NDT temperature is one of the most important

criteria for the selection of a steel for structural application under severe service conditions.

There has long been an interest in the structural use of high toughness, low transition temperature steels,

and more recently with the higher strength, quenched and tempered steels such as the ASTM A543-65 (Class I

and( Class 11), MIL-S-16216G, llY-80, HY-100, and HY-130; these steels are currentlybeing used for the shells of

pressure vessels and hulls of oceanographic vzhicles. The steel used in this study is a quenched and tempered

rnedium alloy'steel having an 85,000 psi yield strength and the following nominal percent composition: 0.16 C,

0.36 Mn, 0.015 V , 0.015 S, 0,25 Si. 2.75 Ni. 1.50 Cr and 0.35 Mo. This steel is commonly classed as a low carbon

nickel.chrme.iolybden urn alloy steel.

LITERATURE SURVEY

The impact transition temperature is some arbitrary point in the temperature range over which the fracture

cLf,=acteristics of a metal with a body centered cubic lattice, and sometimes with a hexagonal close packed lattice,

exhibit a warked loss in ductility. This change is evidenced by (I) a large drop in the c'.-rgy required to produce

fracture. (2) a marked change in the fracture surface appearance from a dull, fibrous-looking transcrystalline shear

fracture associated with high energy absorption to a bright, faceted cleavage or intercrystalline fracture associated

with low energy absorption, and (3) a decrease i the anount of lateral contraction at the bottom of the notch

used to initiate the crack, oi a decrease in iie amount of lateral expansion at the "hinge" point of th,, impact

specimen opposite the notch. The temperature rag, tver which this transition from ductile to brittle Ochavior

occurs is sensitive t( test conditions such as strain rate, notch accuity, aind residual or applied stress levels, as

well as the toughness of' the material at the time of testing; as a result tihe transition range is usually determined

by using a stanldardiied inopact specimnen and test nlethod to test the material at a series of ternperat ures.

*Thc Nt'I ik the cvitpcfattlfe of initial tfanilion from plane strain o inked-imode fracture and represents the tempera-lure at which very ,%vnatl flaw Wie% sflcc for dynamtc fractlure initiation, 'file ND)T temperature is generally nt indesahle iolho, hirply ltalin. ('arpy enrgy cvrve.

2

ThI, "transition temperature" then is defined as that temperature in the transition range at which some pecific

condition occurs such as:

1. The fracture azppearance transition temperature, FATT, is the temperature at which the specimen is

50 or 100 percent fibrous and is called the 50 percent FATT ot the 100 percent FATT, respectively.

2. TIu- fraeture temperature is the temperature at which the Charpy V.notch (Cv) impact specimen

absorbs some specific amount of energy. A common energy value used to characterize mild steel ship plate is

15 ft-lb; this cri:.erion would be referred to as the Charpy V.notch 15 ft-lb transition temperature.

3. The temperature at which the drop weight impact test specimen exhibits no ductility is the NDTtemperature, and is the temperature of initial transition from plane strain to mixed-mode fracture. The NDT

temperature represents the temperature at which very small flaw sizes suffice for dynamic fracture initiation.The drop weight NI)T temperature is not generally indexable to some specific point on the sharply falling

Charpy energy curvc.

An extensive hody of literature exists on the effects of microstructure, composition, and strength on the

impact transition temperature. The data are mainly concerned with the effects of alloy additions to plain

carbon steels, or afloy steels having 0.30 percent or more carbon, whose microstructures are frequently not

specified other thn as resulting from a normalizing treatment, and these data often ignore the austenitic grain

size of the steel prior to cooling. These studies are seldom based on the drop weight nil ductility transition

temperature.

GRAIN SIZE EFFECT

Changes in grain size have been reported to have a detrimental effect on the ductile to brittle fracture

transition temn-crature of steels. 1- 32 Some authors report the effect of change in the average grain diameter,otherz in terin; of the change in ASTM micrograin size. Some authors report grain size Iti terms of: (1) ferritic

grain size a% measured on a polished and etched section; (2) prior atistenitic grain size, measured using the

McQuid-Hih test method; (3) prior austenitic grain size as measured using an etheral picrk: acid etch to outline

the prior r tcnitic grain boundaries on a polished specimen;and (4) fracture grain size as determined hy com-

paring a men broken at low temperature and exhibiting a crystalline fracture appearance to the fracture

grain sizet ;i st of specially prepared specimens used as a standard.

The sArious methods of reporting grain size are not independent. The ASTM micrograin size is related to

the avei;,c giain diameter as follows:

1. The ferritic grain size is highly dependent upon the prior austenitic grain sL'ex and to a lesser degree

upon otlher factors 2'4-7, 13, 17-19. 22. 24, 26, 28, 30,31 such as:

a. Cooling rate from the austenitizing temperature. The slower the cooling rate the coarserthe resultant ferrite due to fewer nucleation sites and more time for these sites to grow.

b. Temperature at which transformation from austenite to ferrite, or some other product,takes place. The higher the temperature the fewer nuclealtion sites and thcrefore larger grains.

c, The amount of ferrite grain growth that takes place after transformation while cooling,tempering or stress relieving, Longer times allow growh of some grains at the expense ofsmaller or less favorably oriented neighbors.

irencei arc Iitrd on page 119.

d. The quantity and dispersion of inclusions, oxides and nitrides that may inhibit ferriticas well as austenitic grain growth.

e. l)eoxic:ation practice.

2. The fracture grain size of embrittled material is usually the same as the prior austenitic grain. 11, 12

The strong dependence on prior austenitic grain size, all other things being equal, is due to the fact that theaustenite to ferrite reaction nucleates at the prior austenite, gamma, grain boundary. The finer the prioraustenitic grain the greater the grain boundary surface area and thus the more nucleation sites available for the

gamma to ferrite transformation. Slower cooling rates or longer times at temperature allow growth of existinggrains either from existing gam:-,a or at the expense of other smaller ferrite grains.

Most of the investigators who have reported on the effects of grain size have used the Charpy V-notch im-pact test to determine the fracture transition temperature !1 3, 4, 7-11, 13-21, 23, 24, 29-31 only a few have usedthe Charpy keyhole specimen,2' 6 the drop weight impact test specimen,21, 24, 32 or the crack or tear test speci-men. 6 How the fracture transition temperature is defined has a niirked effect on the magnitude of the shift intransition temperature caused by changes in grain size. 26 Some of the definitions of fracture transition tempera-ture used in the literature are as follows:

1. For Charpy V-notch and keyhole specimcns

a. Level of impact energy absorbed was frequently used to determine the transition tempera-

ture.7, 8, 13, 14, 16, 20, 21,24 The values used ranged from 8 to 40 ft-lb, with perhaps the mostcommon value being 15 ft-lb. The IS ft-lb level is commonly used for plain carbon steels be-cause it is the level found by the National Bureau of Standards and the Ship StructuresCommittee to characterize the susceptibility of plain carbon steel ship plate to b,ittle fracture.The higher values are associated with quenched and tempered alloy steels.

b. A specific amount of fibrous fracture appearance on the broken surface of the impact

specimen is sometimes used to define the fracture transition temperature. 4' 8, 16, 21,23, 26, 27

This is usually referred to as the fibrous fracture appearance "-nsition teniperature (FATT),and 50 percent fibrous fracture is the most commonly used valje,

c. An amount of lateral expansion or contraction in the deformed area of fracture is

occasionally used to define the transition temperature.21

d. The inflection point or sharp drop in the energy absorbed versus temperature curve is

used.,, 20

e. A mean energy level such as the mean energy absorbed at +100 and -196 C has set the

transition temperature,4 as his the temperature at which the Charpy energy absorbed is one-

half the maximum' energy absorbed at upper energy shelf,24 and as "... the middle of the

temperature range over which an impact value of,10 ft-lb was obtained. *2

f. Sometimes just a visual comparison is made of energy absorption versus test tempera-

lure curves for one maItedial or condition as opposed to another. I

2. For drop weight impact test specimens, the transition teniperature is defined as the highest tempera.tre at which the crack from the brittle crack starter weld will propagate across the surface of the specimen and

around the edge on at least one side. 21' 24, 32

3. The crack; arrest or tear test transition temperature is the temperature at which a runnitng crack is

halted or arrested by the material. In a slectmen having a thermal gradient, the crack runs toward the warn

Charpy Impact Tests

All Charpy V-notch impact tests were performed in accordance with ASTM Standard E 23.64 us.ing apendulum type of impact machine with 268 ft-lb capacity and a striking velocity of 16.85 ftlsec. Energy lossesdue to friction were negligible (less than 1.5 ft-lb) over the full range and therefore were disregarded. A Tvpe A

Charpy V-notch (simple-beam) impact test specimen 0.394 x 0.394 x 2.165 in. was used for these tests.Most Charpy specimens were cut parailel to the 5 in. dimension of the drop weight test block and thus were

transverse to the major direction of plate roll (see Figure 3). Transverse Charpy specimens were selected for thisftuly so that if banding or nonmetallic inclusions were present in the test specimen, they would be oriented in

such a way as to offer the least resistance to crack propagation and so as not to act as crack arresters and thus

mask the properties being studied. However, a few coupons were large enough to provide extra material tomake a few oversize 5/8 x 2 1/4 x 5 1/4 in. specimen blanks (see Figure 4). These blanks were heat treatedand used to make a set of longitudinal, 2.165 in. dimension parallel to the major direction of plate roll,

Charpy V-notch impact specimens. Each specimen was identified before notching on both 0.394 x 0.394 endfaces with the drop weigit spccimen number, test direction and specim.en nuumber. All Charpy specimens were

notched so that the notch was perpendicular to the plate surface through-the.thickness of the plate.

The Charpy V-notch specimens were machined to closer tolerances than recommended by the ASTM inorder to minimize scatter in test results; the permissible variations used were as follows:

Adjacent sides 90 deg ±30 rin

Cross-section dimension ±0.001 in.

Length of specimen (L) ±0.002 in.

Centering of the notch (L/2) ±0.001 in.

Angle of notch 45 deg ±22 1/2 min

R1dius of notch ±0.005, - 0.002 in.

Dimensions to bottom of notch ±0.001 in.

fImish requirements 32 pin. all over

Specimens were cooled to test lemperattre using a liquid bath in an insulated container. Specimens wereremoved from the bath and centered in the test machine using self-centering tongs sirmilar ' those shown inASTM F- 23; the tongs were chilled in lhe low temperature bath with the specimens for at least 15 min and wereimmediately returned to the bath each time after they were used. Bath temperatures were controlled using adummy specimen wilh an embedded thermocouple. The values of energy absorbed by the specimen reportedherein are usualy the average of two specimens broken at the indicated temperature; individual specimen te-stilts are plotted on the energy absorbed versus test temperature curves.

METALLOGRAPHIC PROCEDURES

Meta!hlraphic specimens were cut from ile undeformed ,reas of the drop weight test specimen blankand far erough from the crack startcr bead to be sure that no alloying elements had diffused into the specimen.1.4tlgi im.ldfll arid ti.insverse specimens were prepared and mounted in Iransoptic mounting resin. After theslcifyncus werv: wouted, I 18 in. (if specimen surface was machined off in a lathe under coolani using a sharp,;iltmlng I ool. ' lhe %hlit'ac was tr ,ressively grouind using I80., 240. and 600-grit silicon carbide papers.

least 45 minutes prior to test by manually tirring tile bath and addingr,y ice or liquid nitrogen as required.

Gloves were used to handle the specimens by the heat treating loup welded to the end (5/8 X 2 in. face). Inter-pretation of the specimen fracture as to "Break" or "No Break" was as recormmended in ASTM E208. The only"No-Test" specimens encountered in these tests were a few that were cut a little thin, less than the recommended

0.62 ± 0.02 in. thickness. These specimens were not included in the measurement of the NDT temperature.

Tensile Tests

All tensile tests were performed in accordance with ASTM Standard E8-65T, Tension Testing of MetallicMaterials. A small size specimen, 0.250.in. diameter with a I-in. gage'length, proportioned to the standard Type1, 0,505-in. diameter tension test specimen was used. The specimens were taken parallel to the original direc-tion of plate roll, thus the overall "ength of the specimen was limited to 2 in since this is the size of the 1-3 dropweight specimen in that direction (see Figure 3). The 2-in. overall length was achieved by slightly reducing the

length of the threaded ends of the specimen.

The specimens were tested in hydraulically loaded universal testing machines calibiated in accordance with

ASTM Standard F 4-64, Verification of Testing Machines, using proving rings. The testing machines were round

to indicate the correct load within :r0.5 percent in the loading ranges used in these tests. Self aligning, spherical

bearing devices were used to ensure axial loading of the test specimens. Load-strain turves were recorded for allspecimens using an averaging, I-in. gage length, microformer strain indicator (extensometer), and an autographic

recording device calibrated over the range of the expected yield. Specimens were tested at a speed of 0.002

in./in./mmn or less, tip to and slightly bcyond the yield point. The 0.2 percent offset method was used to de-

termine the yield strength. The valLes reported are the average of two specimens.

Compression Tests

All compression tests were performed in accordance with ASTM Standard E' 9-61. The specimens testedwere 1/2-in. diameter by 2-in, long cylinders made with special care to ensure that the ends were parallel to each

other within 0 002 in. and perpendicular to the long axis of the specimen within 0.005 in. The specimens weretaken parallel to the original direction of plate- thus they were machined parallel to tile 2-in. dimension of thedrop weight specimen they were cut from (see Figure 3).

The specimens were tested using hydraulically loaded testing machines calibrated in accordance with ASTMStandard E 4-04, Verification of Testing Machines. The testing machines were verified using proving rings andfound to indicate the correct load within ±0.5 perceat in the loading ranges used in these tests.

All compression tests were run using a subpress with a spherical seated bearing block at the end of its shaftto ensure full contact with the end of the specimen and to ensure true axial loading, liar6ened steel bearing

blocks with smooth, ground, parallel faces were used in contact with the ends of specimen to prevent possible

danmage to the faces of the snbpress. The ends of the specimen were greased with lubriplate to eliminate end

restraints. Because of the hiape of the specimen it was not recessary to use a jig to prevent lateral buckling.IA'ad-strain curve, were recorded foi all specimens using an averaging, I-in. gage length, nicroformer strain indi-

cator (compressoliete') and an autographic recording device calibrated over the range of the expected yield.Specimens were kcted at a speed of 0.002 in./iii./min or less up to and slightly beyond the yield point. The0.2 percent offset method wa used to determine the yield strength. Tho valht,'s reported are the average of twospe'Cirn~n.

1,4

875 F Isothermal Treatment (Semi-Bainitic)

Specink.. blanks that were treated isothermally at 875 F to ultimately give a microstructure consisting ofa mixture of tempered bainite and tempered martensite were treated as follows: 1640 F - 1/2 hr - transferredto 875 F -- held 152 (or 1600) sec-BQ; -110 F- 1/2 hr; temper 1150 F - 1 hr - BQ. Additional specimenblanks available from coupons X-1 2, X-I 3, and X-i 5-X-1 8 were treated as follows: 2000 F - I hr -- transferredto 1640 F held 5 min (minimum) - transferred to 875 F - held 1600 see - BQ; -110 F -i/2 hr; 1150 F - I hr-- 13Q,

1200 F Isothe'mai Treatmenti (Semi-Ferritic)

Specimen bManks from all coupons were isothermally treated at 1200 F to ultimately give a microstructureof tempered martensite and tempered ferrite. In several cases a little pearlite was found to have been formedduring the isothermal trc.tment. These coupons were treated as follows: 1640 F - 1/2 hr - transfer'ed to1200 F - held for 3350 sec (or extra sets of specimen blocks from coupons Xl, X-2, X12, X-I 3, X. 5- Y-18 andY-2, Y-4, and Y-l I were also held for 8500 sec) -BQ;- -110 F - 1/2 hr; 1150 F-- 1 hr - BQ. There was in-sufficient material to give any specimens with a coarse austenitic grain size the 1200 F isothermal treatment.

TESTING PROCEDURES

In general an attempt was made to adhere to standard ASTM testing procedures. In a few cases it wisnecessary to make slight deviations from standard practice to take into account the specific conditions of thesetests; any such deviations will .- ,_ 'dicated below.

Drop Weight NDT Tests

All drop weight tests were performed in accordance with ASTM Method |: 208.66T, uning the type P-3specimen which is 5/8 x 2 x 5 in. Thc specimens were supported as a simple beam and subJected to an impactload of 259 ft-lb produced by dropping a 223.lb weight 1 ft 2 in.; this drop energy was found to consistentlycrck the crack starter weld bead ard to cause the specimen to consistently contact the anvil stop.

As stated previously, the specimens were prepared with one, as rolled, 2 X 5 in. face. The llardex-N crackstarter bead was deposited in the center of this face prior to heat treatment of the specimen so that the heat-affected zone created by deposition of the weld bead would not interfere with the correct measurement of theNI)T temperature of the base material. After final heat treatment, the crack startcr weld beads were notchedas required in ASTM E 208, uring a dhin abrasive disk. 'rhe combination of the isothermal treatments and thesubseclrent 1150 F temper was not found to prevent proper functioning of the crack starter bead; upon impact,

a good crazk formted in the bead in evezy case.A group of four or fihe, usually five, specimens was given a selected heat treatment and ued to determine

tl.! Nl)T tempera: tire. Since tests weie conducted on a known material with a known thermal history, the NDTtemperature was determined within 10 F by three to five tests, the usual number being four. Prior to testinge.Ich specimen was placed in a large insulated liquid, isopentane, low temperature bath with at least I in. ofliqluid all around tiz specimen. A thermocouple embedded in a dummy specimen was attached to an automaticrecorder and tisd to c nwtol the bath temperature. The bath temperature was maintained for a period of at

I.121

the table lists the identification numbers assigned to each coupon. Coupons made by the largest producer havennumber startihg with an X and coupons made by the other pi oducer have been assigned a number piefixed witha Y. To obtain the maximum use of each coupon and to make specimen location as uniform as possible for allspecimens, specimens were taken from adjacent to the original plate surfaces (see Figure 1). In fact, to minimizemachining costs, the original plate surface was made one of the 2x 5 in. surfaces of the drop weight specimenblank. The specimen blanks were 5/8 x 2 x 5 in. with the 2-in. dimension in the direction of plate roll and the5-in. dimension parallel to the plate surface and perpendicular to the direction of plate roll (see Figure 2). Thisis the size of specimen required for subsize drop weight NDT temperature tests In ASTM Method 1, 208. Eachspecimen blank was coded on the end with the appropriate number, location, test direction, and specimennumber. These specimen blocks were used for drop weight tests, after that the broken specimens were used tomake mechanical property test specimens.

HEAT TREATING PROCEDURES

The 5/8 x 2 x 5 in. drop weight specimen blanks were prepared for heat treatment by (I) welding hand-liiig loops onto one of the 5/8 x 2 in. faces, (2) depositing a brittle crack starter bead on the 2 x 5 in. surfaceof the specimen that represented the original p!ate surface of the specimen using a 3/16-in. diameter liardex Nhardfacing electrode, (3) normalizing the specimens by heating to 1650 F, holding for I hour and then air-cooling, and (4) sandblasting the specimen prior to final heat treatment.

The final heat treatment of the drop weight specimen blanks consisted of austeniti7ing in a neutral saltbath at I40 F for 1/2 hour and then either water quenching in a brine solution or quenching into an isothermalneutral salt bath, holding for a prescribed time and then brine quenching. After quenching, all rpecimens wereinmediately given a low temperature (about - 110 F) treatment in a bath made of acetone and chunks of dry ice;all specimen blanks were given at least 1/2 hour in this bath to ensure that they came down to temperature and

that all retaine'd austenite was transformed to martensite by the treatment. All specimens were subsequently

tempered in a circulating air furnace at 1150 F for I hour at temperature and then water quenched in a brine

solution (11). In a few cases, where sufficient material was available, a few sets of drop weight specimen blankswere austenitied at 2000 F for I hour in a circulating air furnace, then transferred to the 1640 F salt bath for5 minutes, minimum, to standardize the temperature differential between the isothermd bath and the specimenprior to immersion. All furnace and bath temperatures are manually controlled using potentiometers to within

+5 F. Specimens were heat treated by suspending them by their loop from a fixture that allowed them to betreated as a group. This fixture held the specimens 1 1/2 in. apart and suspended them vertically in the salt baths

to permit circulatin of the salt.

Fullv Quenched (Martensitic)

Specimen blanks that were fully quenched to form a fine prior atstenitic grain size to ultimately give ate:pered mariensitic microstructure were !ieaied as follows: 1640 F - 1/2 hr - BQ: - 110 F - 1/2 hr:tempet I 150 F - I hr - 111). Specimen blanks from coupons X-!2., X-I 3, and X-I 5--X-18 which were quenchedfrom a comrse prior aunsteniti4: size were treated is follows: 2000 F I hr - 1640 F 5 mll (minimum) I3Q,

1101 1/2 hr; temper 1150 F --1 hr - BQ.

12

under a wide variety of conditions, and tempering treatment. Increases in strength produce corresponding in-creases in impact transition temperatures unless other parameters are adjusted to compensate for the increasedstrength level. 27, 35 Conversely, a decrea'e in strength brought about by a long time temper is no guarantee thatthe impact transition temperature will improve, since for a wide variety of quenched and tempered steels a 100hour stress relief will reduce the strength and at the same time increase the transition temperature this is attri-buted to a susceptibilit) to temper embrittlement. 38

In bainitic steels, increases in strength brought about by refining the bainitic-ferrite microstructure willhave an adverse effect oin impact transition temperature unless compensation is made by refining the prior aus-tenitic grain size also. 27 The dependence of the impact transition temperature of bainiti: steels on prior aus-tenitic grain size is attributed to the fact that the upper limit on the size of the bainitic-ferrite plates is the prioraustenite grain diameter.27 Tempering bainitic steels at a temperature above the temperature at which thebainite formed improver its impact transition temporature by lowering the strcngth.4 2

EXPERIMENTAL PROCEDURES

SUMMARY

The general test program followed in this study can be summarized as follows. Test coupons were col-lected from commercially produced plates made by the two major producers of a low carbon, nickel-chromemolybdenum steel.* These coupons were selected to represent the broad range of chemical composition towhich this steel is melted in commercial practice. These coupons were slabbed to convenient size, analyzed forchemical composition, heat treated, and tested. All heat treating and testing was performed on 5/8 N 2 x 5 in.drop weight blanks heat treated so as to eliminate the effects of variation in hardenability, plate thickness, aridmill heat treatment. Isothermal heat ticatments were used to produce controlled microstructures, and a uniformtempering temperature of 1150 F, slightly above the minimum value permitted by the specification, was used.The isothermal holding times ard temperature were tentatively selected to produce approxintely 25 percent±10 percent of each nonmnartensitic product, bainite or ferrite, for an average composition. Microstructures pro-duced by a given heat treat were measured rom each coupon using microspecimens taken from the broken dropweight test specimens; mechanical properties were measured on these same test specimens.

MATERIALS

Coupons were collecte, over a period of years from plates made during actual mill runs at the plants ofthe two major producers of .;is steel and represent only a small fraction of heats melted during this period.T-he coupols were obtained after final heat treatment from the end of the plate when the plate was cropped toordered size. All LIupOns were cut from material adjacent to thi- ordered plate. The actual coupons selectedwere chosen by the investigator , the hasis of the producer's chemcal analysis to represent a fairly broad rangeoIf the chemical compositions used for this steel in commercial mill practice. A total of 27 coupons representing22 heats were collected; 1 6 coupons were from one producer and I I coupons from the other. Table 1 summna-rizes the heat numbers, mechanical properties, and chemical compositions reported by the producers; In addition,

* ornmonly caled I|Y-1X4 %eet and rnioe in accordance with AST'M Slwcificitlon A543-405 and Military Specification

MII.S-l16216G tSI II1'.

11

COMPOSITION

The effects of compositional changes in tempered martensitic and lower-bainitic microstructures on impact

properties are threefold: first, the basic impact fracture transition temperature is shifted due to the presence oftrace impurity elements; 23 , 36, 37 second, the susceptibility to embrittlement at low tempering temper,'tures,

500 to 800 F, may be affected; 10 . 37 and third, there may be susceptibility to temper embrittlement duringeither long time tempering treatments or slow cooling instead of quenching from the tempering tempera-tures.9, 10, 22, 23, 34, 36, 37, 41 In addition, the composition changes may be detrimental to impact properties ifthey increase the strength markedly or if they tend to retard softening during tempering.27 ' 35, 31

Te'ts on special high purity quenched and tempered steels have shown that certain minor impurity ele-ments, some of which are invariably present in commercial steels and for which many are not roitinely analyzed,such as P, Sb, As, Sn, and N2, are primarily responsible for embrittlement. 23' 36, 37 The high purity steels showvery low impact transition temperatures :nd very little susceptibility to embrittlement as a result of usually det-

rimental tempering treatments. Susceptibility to low tempering temperature embrittlement is increased by Mn,P, Si, Cr, As, Sn, N2 , not affected much by Ni, Cr, and Mo, and seems to be inhibited by Al.10' 37 The mostdetrimental effects were observed for P and N2 .

37

Susceptibility to temper embrittlement due to either long tempering times in the region from 900 to 1100 For slow cooling from the tempering temperature is increased by C,9 ' 2 2 Mn, 3 4' 36,37 p10O 23,36,37,41 Si 36' 37

Ni, 34 Cr,34 As, 36 ,37 Sb, 36 '3 7, 39,41 and Sn,23, 36,37,39 not affected much by N2 , Bi, Cu, Co, Ge, Ga, and

Zr,36' 37 and inhibited by Mo, 10 ' 34 ' 36 , 39 and W.39 There was no evidence in studies of a quenched and tem-pered Ni-Cr-Mo steel that there was any interaction between regular alloying additions and trace elements that

exaggerated their total effect. 3 9' 41 For some hardened steels, there is evidence from the impact transitiontemperature and the fracture appearance that some embrittlement occurred during the initial tempering at anelevated (1112 to 1202 F) t7e4perattre. 1

A study of a quenched and tempered Ni-Cr-Mo steel that was not deliberately given an embrittling treat-ment shows that P, Sb, and Sn have a detrimental effect on impact transition temperature, As had little or noeffect, and Al additions were actually beneficial. 39 Phosphorus was the most damaging, and the effect of As,which had been shown to be detrimental to deliberately embrittled steel, showed little effect in the not deliber-

ately embrittled steel. The effect of Al shifted from detrimental in embriltled material to beneficial in non-embrittled material; the benefit of Al additions was attributed to a decrease in the prior gamma grain size. Othereffects reported are 1!;.'t the optimum carbon content in quenched and tempered steels is about 0.38 percent andthat this amount optimizes both NDT temperature and maximum Charpy V-notch shelfenergy.40 Sulphur isreported to have no effect on impact transition temperature but to lower the maximum Charpy shelf energy. 0

13, C, Mn, Ni, and Cr additions can be detrimental to dhe impact properties of low-carbon bainitic steelsif they reduce the austenite decomposition temperature and as i result either increase the strength excessively

or a micrstruture27or produce a mixed microstructtre of bainite and marte:site. Secondary hardening elements like Mo and Vhave a tendency to reduce impact properties of low-carbon bainites, in that they increase the hardness of theinitial microstructure and they tend to precipitate carbides which retard softening during tempering. 27

STRENGTH

The effect of strength level on impact transition temperature is complicated by the dependency of strengthon the interiction between composition, grain size, microst ructure produced by the transformation of at'stenite

10

In martenstic steels transgranular and intergranular fracture refer to austenitic grains present prior to

q.enhiig and nrjt to the ferritic grain structure developed by temperinp the martensite. 11' 12, 23, 29,36,37.4 1

The ,"lat-ve amount of transgranular and intergranular fracture developed by the impact fracture of a temperedmarte,asite depends on the following:

1. The relative temperature in the fracture transition region. The lowest .emperatures tend to givemore intergranular fracture.1I

2. The amount of temper embrittlement. Certain combinations of temperatures and long holding times,

will increase the tendency to intergranular fracture if the material is susceptible to temper embrittlemcnt. 12

3. Blue brittleness, associated with a 500 F tempering temperature. This tends to give an intergranular

fracture and shows a brittle appearance over a considerable temperature range and which slowly disappears .t

higher temperatures without ever exhibiting an abrupt transition.1 I

4. Nonembrittling tempering causes fracture to show a progression from transgranular fracture after

a 500 F temper, and a return to the transgranular fracture mode after a 1200 F temper. Tempers from 550 F

to 1200 F exhibit brittle fractures composed of a mixture of intergranular and transgranular brittle partings. 1I

5. Iligh purity tempered martensite will exhibit only tran,;granular cleavage when it breaks in a brittle

way whereas steels containing 0.04 atomic percent phosphorus or. tin showed at least some intergianular fracture

along the prior austenitic grain boundaries with the proportion increasing with grain size and with embrittling

treatments.23 Phosphorus containing steel, 0.023 weight percent in a 0.3 C, 3.0 Ni, 0.75 Cr steel, tended to

show more intergranular fracture than the same steel doped with 0.074 weight percent Sn.Summarizing, the impact fe'actoe path is characterized by the transgranular or intergranular mode relative

to the ferritic grains in plain carbon steels having a microstructure consisting of ferrite or ferrite and pearlite,

and to the lt 1Vfl fh II I I4 t ItItI lit 1ltW1lltlo qI111 tlillpfiroft 11|oI ion! l which shows some degree of temper em-

brittlement. The 'mpact properties of bainitic steels are separated into two distinct types, with the high temper-

attire bainitic structurcs behaving like ferrite.pearlite structures and low temperature bainitic structures behaving

like martensitic structures; the impact properties of the low temperature bainite 'iprove with decreasing prior

austenitic grain sizc5. 27

MICROSTRUCTURE

The general effect reported in the literature is for tempered bainitic steels to have impact properties inter-

mediate to those of tempered martensitic and tempered pearlitic steels of the same hardness, with the tempered

martensitic steel having the lowest impact transition temperature and the highest level of energy absorbed above

the transition temperature. 27 ' 33, 42 The ine bainitic structure responds to tempering like a quenched steel:

the transition temperature steadily decreases with increasing tempering temperature. 27, 3, Composition of low

carbon bafnitic steels has little effect on the impact properties if similar microstructures having the same grainsize and same hardness are compared. 27 In order to get adequate impact resistance it! bainitic or martensitic

steels., they must be tempered beyond the stage where naxinmm e:ondary hardeihing takes place- secondary

Iardening clue to the presence of other alloying elements has a deleterious effect on impact reoistance. 22

lven after tempering, a portion of upper bainite is more detrimental to the impact propertles of an otherwise

mirtensitic nicroit irctlre than is a portion of lower bainite.4 2

side of the specimen. These tests are not used extensively and apparertly are seldom used to evaluate the effectof grain size on the transition temperature.

Other parameters associated with the material, and its treatment and microstructure have been reported tohave in interaction with the grain size to change the magnitude of the shift in the impact transition temperature.These factors are: impurity level and composition changes; 5 , 7, 8, 19, 23 commercial size heats as opposed tolaboratory si7t heats or melts; 9' 31 straining or cold deformation; 5' 12, 26 thermal treatments that cause asecondary effect to overshadow grain size effects, retention of austenite,3 temper embrittlement 1 ° ' 15, 20. 23, 26and growth of grains; 2, 5, 7, 23 and microstructure changes.1. 3, 14, 17, 20, 22, 26, 27 The effect of each on theapparent shift in transition temperature with grain size (AT/AGS) may either be positive or so negative in nature

as to make grain size appear to have no effect.

As a result of these complications, each reported value of AT[AGS must be evaluated in terms of thespecimen used; the criterion defining transition temperature; the composition, including impruity levels; source

of the material, lab versus production melts; thermal treatment; straining; microstructure; and the range overwhich grain size was actually varied. If this is done, a few general observations may be made on the specificeffect of changes in grain size on the impact transition temperatiie. For convenieace, this report will alwaysrefer to grain size as the ASTM micrograin size. The following equation represents the relationship between thenumber of grains per square inch (n) and ASTM micrograin size (N) : n

Ferritic Grain Size

The ferritic grain size is commonly used in the literature to report the effects of grain size on AT/AGS.Thewe are three reasons for this: first, ferrite is the major room temperature equilibrium phase in the hypoeu-tectoid plain carbon steels used in most of the studies reported in the literature; second, ferrite with varyingamounts of carbide is the primary room temperature microconstituent of quenched and tempered'alloy steels;and third, the final ferritic grain size is readily measured using standard metallographic techniques.

Changes ir, the size of the ferrite subgrain size are reported to affect the impact transition temperature. 18, 31An increase in the sub-boundaries, a decrease in the ferritic subgrain size, raises not lowers the Charpy V-notch15 ft-lb transition temperature (20 F/per ferrite grain size number). The ferrite subgrain boundaries examinedwith an electron microscopc show a lower dislocation density than the boundarics associated with the trueferrite grain boundary. 31 While these authors were puzzled by this effect, a possible explanation must heri,inted wut. Current dislocation theory says that the subgrain structure is caused by a preferential alignment ofdislocations having like signs. As the number of dislocations increases, the subgrain size would be expected to

decrease and the strain in the grain increase. The usual effect of straining is to raise the transition temperaturefor ductile to brittle fracture. In fact, temper embrittlement has even been attributed to strain deve!oped across

prior austenitic grain boundaries as a result of the contraction of the ferrite lattice when dissolved alloying Pie-

ments are removed from solulon in ferrite by a reaction with existing minor carbide phases. 1 ' Thus it is logicalto expect that as Ihe ferri:ic subgrai size gets smaller indicating increasing dislocation densities and higher strainsin the lattice, the impact 1ransition temperatire would Increase as opposed to the usual decrease in AT/AGSassociated with a finter rue ferritic grain size.

References 2, 6. 7. 1I I, 17, 18, 24, 28, and 31 show that a decrease In the ferritic grain size of a plain

carbon steel sufficient to increase the ASTM grain size number one unit will reduce the transition temperature

22 1" on the average, tei effect ranging from an Increase or V F to a decrease of 30 I., Tie 9 F value,

Reference 31, is undoubtedly due to the normal variation in the measurement of transition ,emperature of ±10 Frepiorted by Gross2 l and to the small variation in grain si.e, less than one unit. 31 The -39 F value for AT/AGSis difficult to rationalize since the author reports neither the type of impact specimen he used nor the criterionused for determining the transition temperature.

These ATIAGS values seem to be independent of carbon content in these plain carbon steels, It is reportedthat of the factors: ferrite grain size, pearlite island size, pearlite spacing and amount of procutectoid ferrite, itsit influences pearlite carbon content; only the ferrite grain size showed any degree of correlation with the CharpyV-notch 15 ft-lb transition temperatue t1 7 The AT/AGS value's dependence on test method is seen in Reference24 where the author reports a value of-IS FIAGS based on the Charpy V-notch 15 ft-lb energy transitiontemperature and - I F/AGS based on drop weight NDT.

Austenitic Grain Size

lncreases in the prior austenitic grain size have usually been associated with increasing impact transitiontemperatures.1, 3, 4, 8-14, 16, 19, 23, 27 In fact, as temper brittleness increases in martensitic-ferrite, the fracturepaths become increasingly intergranular and follow the prior austenitic grain boundaries. 11, 12,.29, 36, 37,41The relationship between the Charpy V-notch 10 ft-lb and 50 percent FATT transition temperatures and prioraustenitic grain size was linear for hardened 0.84 and 0.95 C steels. 16

It was reported for tests on SAl 1045, 2340 and 3140 having various microstructures and tempered to thesan.: hardness, that increasing the prior austenitic grain size increased the impact transition temperature of aircooled and tempered (pearlitic) material (AT/A-yGS from -25 to -120 F) but had the opposite effect (+5 to+20 I) on quenched and tempered material. 3 fTh slight improvement in the quenched and tempered steels asaustenitic grain size increased was attributed to the increase in hardenability as the grain size increased and to theresultant increase in the amount of martensite formed during quenching. The impact l'ractu re transition temper-atuire ot' the tempered martensitic structure was always better, i. e., lower than that of the tefnpered pearliticstructures. IncreAsing gamma grain size also increased the notch sensitivity of the pearlitic microstructure.' Ina later study it was reported for 0.80 percent C plain carbon steel that as the hardness was increased, pearlitespacing decreased, the impact transition temperature went up, more so for fine than for coarse prior austeniticgrain size material; in fact, there was a minimum in the pearlite spacing (hardness) versus impact transitiontemperature relation For coarse material whereas the fincr prior gamma grain size material showed a continuousin crease.

4

Prior austenitic grain size is also reported to have a controlling effect on the mean impact energy absorbedtransition temperature of 0.12 C bainitic steels, but the effect is lessened as the strength increases with increaseddispersion of carbides. 27 The prior austenitic grain diameter, which is directly related to either the rolling finish-ing temperature or tile reheating temperature used during anstenitizing, has been shown to strongly affect theaverage Charpy energy impact transilion temperature of 70,000 psi tensile strength bainitic steel, with thetransition temperature increasing in an S shaped relationship with increasing prior austenitic grain diameter. 27

The transition temperature of fhe bainitic sleels reportedly ranged from -76 F for an austenite grain diameterof 3.9 x 110 4 in. to 140 F for a 22.3 x 10- 4 in. diameter.

Grain refiners are frequently cited as and demonstrated to have a beneficial effect on lowering the impacttransition temperature 4 5, 7 and even materials which are not slecific grain refiners, such as lead, hut whichreact to inhibit grain growth, will show a beneficial effect, especially when idded to normally coarse grained

Steels.8

Petch is reported to have shown that the Charpy V-notch 15 ft-lb transitiot temperature (7) of ferriticsteels is linearly related to the natural logarithm of the square root of the reciprocal of the mean grain diameter

(d) in the following way: T = A + B In (D 1 12). If one uses n = 2(N-l) wheren is the number of grains per

square inch and N is the ASTM grain size number, and notes 2he fact that n is proportional to d2 , one can re-write the Petch relationship in terms of the ASTM grain size and show that impact transition temperature shouldbe a direct linear function ofN, T=AI +B 1 (N).

An investigation of unhardened C1095 and C1080 plain carbon steels shows that a coarser prio, austenitic

grain size raises the impact transition temperature, regardless of the temperature at which the austenite is trans-formed to pearlite. 8 In addition, it was shown that while lower transformation temperatures raised the ultimate

tensile strength, it also lowered the impact transition temperatures by refining the microstructure. The data be-

low shows the observed range of change in Charpy V-notch impact transition temperature AT per unit change in

prior austenitic grain size, ATI-GS and demonstrates the dependence of AT/yGS on both the criteria used to

measure ATand the steel composition.

Steel Cv (10 ft-lb) CV (50 Percent FATT)

C1095 46-35 F/-tGS 17-19 FhyGS

C1080 27-29 F1'yGS 33-23 F/-GS

Studies of the effc .s of various factors on the impact transition temperature of plain carbon' steel ship

plate have shown that the higher transition temperatures associated with thicker plates are more closely related

to differences in the ferrite grain size resulting from rolling finishing temperature and cooling rate than to differ-

ences in chemical composition.1 3 The finer the grain size, the higher the ASTM grain size number, the lower

the impact transition temperature. The nagnitude of the effect of grain size on AT/AGS depends on the method

used to measure grain size; based on ferrite grain size the ATIAGS for ship plate was -14 FIAGS and based on

fracture, which may partially reflect prior austenitic grain size, 1 1, 12, 29, 36, 37, 41 AT/AGS was -9 F[AGS.! 3

In a study of the effects of impurity level on the way changes in prior austenitic grain size influenced

Charpy V-notch 100 percent FATT transition temperature of a 0.3 C, 3.0 Ni, 0.75 Cr steel, it was observed that

the relative effect of grain size was strongly influenced by both Impurity !evel and thermal treatment. 2 3

As Oil Quenched and Oil Quenched and Tempered Plus

Tempered at 1112 F 168 Hours at 842 F

hligh Purity - 4.1 F/A-tGS - 5.9 F/AyGS

Plus 0.023 Percent P* -13.6 F/AyfGS -19.8 F/A-yGS

Plus 0.074 Percent S?* --28.5 F/AyGS -42.3 F/A-,GS

These data show that the impact transition temperature of high purity alloys is affected by chanscs in gamma

grain size; that tin is more detrimental than phosphorus at equal atomic percents; and that the detrimental effect

of increasing ptior austenitic grain size is greater after an eribrittling treatment than after a simple tempering

trcatment.

"Fach, about 0.04 atomic percent.

Microstructure

The interaction of microstructure and grgin size appears to be rather straightforward; the finer the micro.structure for any given prior austenitic grain size, the lower will be the impact transition temperature. The finerthe ferrite grains in ferritic steels, 2, 4-7, 15, 18, 22, 24, 26. 28, 32 the finer the pearlite colonies and the finer theinterlaminar ferrite-carbide spacing, t , 3, 14, 16, 17 the finer the bainitic-ferrite20 , 22, 27 and the finer the

martensitic-ferrite, 3 , 22, 27 the lower will be the impact t'ansition temperature. It should be remembered that thefiner the prior austenitic grain size,2 ' 6 the lower the rolling finishing temperature, 3 the faster the cooling

rate, 2 , 6, 28 and the lower the temperature of transformation from austenite, 8' 28 the finer will be the resultantfinal grain size.

Because of the increasingly acicular nature of the ferrite grain size as the microstructure is changed frompolygonal ferrite to bainitic-ferrite and then to martensitic-ferrite, it becomes increasingly difficult to make achange in ferritic grain size and as a result, ferritic grain size can hardly be considered a variable.22 As a result ofthe increasing acicular nature of the microsthicture and the fracture path following the embrittled prior austenitic

grain boundary, 1, 12, 29, 36, 37, 41 one finds the tendency is for researchers to report the effects of grain size onAT/AGS in ternis of prior austenitic grain size when dealing with tempered bainitic and martensiticsteels.3, 20, 22, 27 Researchers also tend to report ferritic grain sizes when dealing with plain carbon steels wherethe structure is essentially ferrite or some mixture of ferrite and pearlite.

For 0.02 percent carbon, plain carbon and alloy steels there is reported to be a slight increase in CharpyV-notch transition temperature due to the acicular shaped ferrite grains over what might otherwise te expectedfrom simple changes in ferrite grain size alone.20 This may be due to the fact that the identical grain sizes weredeveloped by a combination of furnace and air cooling from different austenitizing temperatures, which meansthat prior gamma grain size differences or embrittlement during cooling may have caused the differences ratherthan simply the differences between polygonal and acicular ferrite grains.

For 0.32 percent carbon, plain carbon steel, it was found that of ferrite grain size, pearlite island size,pealite spacing and amount of proeutectoid ferrite as it affects pearlite carbon content, only the ferrite grain sizeshowed any degree of corielation with the Charpy V-notch 15 ft-lb tra-irition temperature. 17 As the pearlitespacing is decreased in a 0.80 percent carbon, plain carbon steel, the hardness is increased and the Charpy V-notch8 ft-lb transition temperature increased, more so for fine than for coarse prior austenitic grain size materials. 14

In tests on SAE 1045, 2340, and 3140 having various microstructures tempered to the same hardness, theCharpy V-notch transition temperature of tempered martensite was always lower than that of the temperedpearlitic structures.3

Fracture Path

The sudden decrease in impact energy shown by the truly ferritic, low carbon, little or no alloying steelsis accompanied or brought about by a marked increase in the amount of cleavage fracture and by a marked re-

duction in the plastic defmnation accompanying fracture., 7, 13, 15, 20 The cleavage fracture of the ferritegrains initiates at the ferrite or ferrite-pearlite grain boundaries. 20 The properties of the ferrite rather than thepcarlite control notch toughnes% in killed and semikilled steels. 3 1 The low temperature brittle fracture in theseferritic steels is primarily transcrystalline cleavage of the ferrite grains, though there has been reported an in.

creased tendency to intergranular brittleness with increasingly coarse grains as the oxygen content is raised

.cve 0,00. percent. 15 ' 18, 2)

8

Final polishing was pcrforrned in two stages using 60-cycle automatic vibratory polishers. The semifinal pol.

ishing howl was covered with bleached silk cloth and a slurry of Linde A, distilled water, and aerosol wetting

agent. The final polishing bowl was covered with Gamal cloth and a slurry of Linde B, distilled water, and

aerosol. All specimens were lightly etched in 2 percent Nital and repolished several times during the final stages

of polishing to ensure the removal of any smeared structure and correct development of the raicrostructure.

'ihe inclusion content of the specimens was measured ill both the longitudinal and transverse, through-the-

thickness planes in ;iccordance with ASTM Standard E 45.63, D)etermining the Inclusion Content of Steel.

Microscopic Method 11 was used to determine the length of the longest inclusion, the number and average length

of all inclus'ions over 0.005 in. long and the background rating including all inclusions less than 0.005 in. long.

The vibratory polishing technique was found to leave the inclusions intact and in tle specimen.

The prior aus',nitic grain size was measured on fully-quenched specimens taken from the quenched and

tempered fully m:irtensitic specimens in accordance with ASTM Standard E 112-63, Estimating the Average

Grain Size of Metal, using the intercept or Ileyn procedure. Tile prior austenitic grain size was revealed by the

use of the etching reagents (!esCril)ed in Appendix A. Comparisons made on several different specimens from

different melts showed that both etches indicated the same prior austenitic grain size.

File percent of isothermal transformation product resulting from the isothermal treatments was measured

fo; each treatment. The microspeciniens were each etched in three reagents. Between each of the reagents, tile

specimens were rinsed and dried. The triple etch consisted of (I) a saturated picric acid solution to develop tile

prior :nustenitic grain bound aries, (2) a one percent nital solution to develop the structure in the areas of temp-

ered martensile, and (3) a 20 percent solution of anhydrous sodium inetabisulfite for staining (darkening) the

tempred marlensile areas. The last etch leaves the areas of tile tempered bainite or ferrite white , in sharp con-

trast t) the darker areas of stained tempered martensile. Tile details of these etching solutions and how they

were used are given il Appendix 13. The percent of t ransformation product was estimated by superimposing a

grid on tile specimen and determining tlie percentage of tile grid intersections that fell over each inicroconstit-

tuent. Because of tile finely handed nature of tOC transformation product and the high magnification (50OX

'I lO)X) ireed-d to reslve the structire, extreme care was taken to systematically traverse the specimen aind

read at leal-t It areasi so as to get a representaive average value to report.

EXPERIMENTAL RESULTS

CHEMICAL COMPOSITION

I lie actu~il chlijal tciiiplsii,)ns o1 thre coulponls testedt in tils study are listed in Table 2. These alnalyses

were roade by a conitnierci;il testing laboratory on material located adjacent to the original plate surface of the

couplon. The rilaterial alialyled came from tile top 5/8 in. of the plate hut did not include any heat treating scale

or decirhnriled llh sill Iace which would have iniluenced tile analysis,

Wet cheilnicAl aa ly~es were performed for percent ('arbon (C'), manganese (NIn), phosphorus (P), silicon

(Si), nickel (Ni). ch mie (Ci ). ndvb)denII (Mo), copper (('ii), and acid solhble ard total alumninuin (Al).

Spectrographic ialysis was pelfomied tI. deterlline the( percent present of tire following trace e~cliellils:

v |r;,dlimur (V). leal (l'), ti (Si), 1riagnusiuil (Mg). cohall (Co) and titaniumii (Ti). No values are teponted for

Il'h. Sri, arnd Mg it t ilde 2 sinc they were r i t oind to he present ii delectable amounts. After several samples

lid heell iflraly/d it Vxl, 11iii1d thai1 V arid Co were plesenl in negligible amounts adl therefore, analysis for

thc'c elenllicuts , dv' t (t lirltlvd , I hi' Ip,!'I'ic rs illi',, u t ib .. !u lror ilat ftie Co was a tram p eleentiil

I i

introduced by the nickel addition. The Co was left in with the nickel since it was too difficult to remove andsince they found that it had essentially no effect on the hardenability or properties of the steel. The V was alsopresent only as a tramp element but it came from a variety of sources that went into the make up of a givenheat of steel.

Gaseous analysis was performed for oxygen (02), hydrogen (112), and nitrogen (N2). 02 and N2 are re-ported as a percent and 112 as parts per million (ppm) in Table 2.

A review of Table 2 shows that in general the range of the alloying elements found in the samples selectedfor this study almost completely cover the full composition range permitted for nickel-chrome-molybdenumsteels such as ASTM A 543 and KY-80. The only element that falls outside the combined range for the twosteels is Mn; coupons Y-8 and Y-10 have 0.49 percent Mn. This is only 0.04 percent over the 0.45 percent(0.40 upper limit plus the allowance of 0.05 percent variation over the upper limit) maximum allowed forthese steels.

METALLOGRAPHIC ANALYSIS

The results of the metallographic analysis of coupons used in this study are given in Tables 3, 4, and 5.Table 3 presents the results of the inclusion analysis of these coupons at the mid-thickness and adjacent to thesurface of each coupon in both the longitudinal and transverse kdirections on planes perpendicular to the platesurface through the thickness of the plate. The reporting method recommended in ASTM E 45-63 is used todescribe the inclusion content. The length of the longest inclusion at I OOX is reported in units equivalent to0.005 in. on the specimen along with a superscript describing whether it is grouped (g), very disconnected (vd),

or disconnected (d). The average length of all inclusions over one unit long and excluding the longest is reported

with a superscript denoting the number of inclusions averaged. The background rating A, B, C or D correspondsto the sample photographs in E 45 with A being a rather clean steel and D a rather dirty steel. Figures 5 and 6arc photomicrographs (I OOX) of representative samples and will give the reader a better understanding of the

significance or the data reported in Table 3.The prior austenitic micro-grain size of each coupon is reported in Table 4. The grain size is reported for

the as-received coupon and for the coupons after heat treatment. This table shows that the coupons austen-itized at 640 F had a slightly finer prior austenitic grain size than as-received coupons, and a slightly coarsergrain size after the 2000 F treatment.

lie percent of isothermal transformation products produced by the various isothermal treatments are

summarized for each coupon in Table 5. The large variation in percent transformation product resulting from a

given heat treatment reflects the large range in chemical composition and hardenability of these coupons. The

high hardenability, rich composition coupons had the least transformation take place for a given isothermal

treatment. 1Photomicrographs of typical areas of each coupon for each isothermal treatment are given in

Appendix C. The dark areas arc tempered martensite that has been darkened as described in Appendix 11 and

the light areas are tempered isothermal products.

MECHANICAL PROPERTY TEST RESULTS

Table 6 surnniarimes the mechanical property test results for each heat treatment; the corresponding micro-

9 structures are also included for reader convenience. The table lists the longitudinal tensile and compressive

17

properties, the NDT temperaturc, and the transverse Charpy V-notch impact properties atthe NDT temperature,120 F and + 212 F (representative of the maximum energy shelf). All values reported are the average of two

specimens. Also longitudinal Charpy data are reported for the few cases where it was measured, Appendix Dcontains the full Charpy V-notch curves for most of the coupons.* The Charpy curves show the energy absorbed,the lateral expansion at the notch after fraciure and the percent fibrous fracture appearance of the fracturesurface.

DISCUSSION

A comparison of Tables I and 2 shows that there are some significant differences between the producer'sreported chemical composition of the heats and the actual composition of the coupons. The biggest differencesare in C, Mn, P, and S with the coupon values generally higher than the producer's values. These variations areslightly more thar, what might be expected from normal variation in analytical results and from Variation causedby the differences in the position of the sample in the heat of steel, and point up the need to check the actualcomposition of coupons being evaluated in laboratory experirnenits rather than to rciy on the producer's data.Table 2 shows a good compositional range for the usual alloying elements and a fair spread for most of the restof them.

Table 3 shows the mid-thickness and surface inclusion content of tie coupons. In general, the materialadjacent to the surface is cleaner than at the mid-thickness of the coupon. Even though a good range of in-clusion sizes is represented in these samples, it must be pointed out that no large continuous inclusions werefound.

Table 4 compares the prior austenitic grain size of the as-received material at mid-thickness to the prioraustenitic, gamma grain size that resulted from the austenitizing treatments used in this study. While the as-received gamma grain size was fairly coarse, rar!ing from ASTM 8 to 4, the treatments used in this studybracketed an equally wide range in gamma grain size of ASTM 9.5 to 3.0.

The isothermal holding times referred to in Tables 5 and 6 are time in the hath, not actual holding timesat bath temperature. Appendix E gives a comparison of the actual time-temperattire history of 5/8 x 2 x 5 in.drop weight specimens when they are transferred from 1640 F into either V75 or 1200 F salt baths and corn-pares it to the cooh. g curves predicted using the method of Sinnott and Shync. 47 The mid-thickness of thedrop weight specimen is 'iihin 20 F of 875 F in about 120 seconds, a considerable portion of the 152 secondtreatment at 875 F but a reasonable fraction of the 1600 second treatment. The data indicate that the mid-thickness of the specimens is within 20 F of 1200 F in about 90 sconds, which is negligible compared to the3350 and 8500 second holdirg t ines used at i 200 v. The irportant matter concerning holding times is thatthey produced a reasonable range in the anount of microconstit nents.

Table 5 reports the aniount of bainitc produced during the 875 j: isothermal treatments and the amountof ferrite and pearlite produced by the 1200 F treatments. The balance of the austenite that was untransformedat ,he end (f the isothermal trefit ment was completely transformed to martensite by the brine-wa ter quench and

It7 : treatment. lie ainotill of hainite produced by the 875 F treatment ranged from 0 to 76 percent andv;iriC(l with the hmrdenlability Iof the particular steel and tile isothermal holding, time. As would be exPelted. the

OA few conlirwr I X-4 X-6 a (I a -Y. and Y-(,) only hil t It r or (1tnifmpy specitil nwn tested at each of the tiree indtcitedIcrm pettres. MNtI, 20 1. anm t 212 F. 1hi, wans not notugtt (al to draw it menanitngfiti (harpy data curve %o they arenot thown.

steels with the highest hardenabiliies transformed the least. A similar effect was observed for the 1200 Fisothermal treatment, the amount of ferrite ranging from 0 to 69 percent. Only four of the 1200 F treatmentsproduced measurable amounts of pearlite, ranging from 4 to 32 percent; these steels were low hardenability, lowNi, and low Mn compositions. The fact that only four samples had pearlite in the microstructure must be con-sidered during the subsequent discussion of the effects of microstructure on the NDT temperature. There are sofew points that not much confidence should be placed on the estimated effect of pearlite as opposed to the othermicroconstituents. It must be pointed out that one normally would not find pearlite in the microstructure ofthis low-carbon, Ni-Cr-Mo alloy steel unless something were grossly wrong. Any discussion of the effect ofpearlite is of academic rather than practical interest.

Table 6 gives a summary of the microstructure, tempered-bainite, -ferrite, -pearlite, and -martensite, iongi-tudinal mechanical properties, the drop weight NDT temperature, and transverse and longitudinal Charpy V-notchimpact properties resulting from the various heat treatments. Because of the uniform 1150 F temper, thestrengths reported here are more characteristic of an IY-1 00 steel rather than an HlY-80 steel. The variations instrength reflect the variety of chemistries studied and the different initiai hardnesses, the slightly different re-sponses to tempering, and the different microstructures. The mechanical property data in Table 6, the chem-istry data in Table 2, the inclusion content data in Table 3, the prior austenitic grain size data in Table 4 and themicroconstituent data in Table 5 were combined to form the set which is the basis of this analysis. In summary,the 115 data points represent the 27 different coupons and constitute the sample used in this analysis.

INDEPENDPENCE OF VARIABLES

When performing a regression analysis to determine tile functional relationship between a dependent vari-

able and more than one independent variable (in this case NDT temperat ure and such things as composition,microstructure and strength), it is obvious that the magnitude of each independent varialle must be randomlydistributed relative to the magnitude of each other independent variable; otherwise it will be difficult ift not iml-possible to apportion a part of the observed changes in tile dependent variable to one or another independentvariable. In experiments where the investigator is melting his own material he can ensure independence betweenthe independent variable he wishes to study by careful randomizzstion in statistically designed experiments. Inthis case wr have 27 coupons with a wide variation in chemical composition and mechanical properties. In thisstudy to ensure independence, or at least to ensure that there was riot a strong correlation between the variables

considered, the data making up the sample (115 data sets) were run on a correlation matrix as described in

Reference 43.The resulls from this correlation matrix are sumnarized in Table 7, showing the cross-correlated varhlra s

with a correlation cocfficient from 0.85 to 0.999. 'able 7 gives a numbered list of the 83 variables consideredand indicates which variables are related to another for three ranges of' correlation. The identificationr numberof the variable having the indicated level ot' correlation is listed next to each. For example VI-Salmple Number*

is strongly related (correlatior from 0.90) to 0.949) to variables 10 and 15 (amlolut of coppetand anolirt of'

cobalt). A look at Table 2 shows that this is in leed tile case. Producer X has much less copper in his rrwlts than

producer Y anid cobali was measured only oil coupons troin producer X.

'Sjinrptc numbert w deriv,,It hy nmlttP in ! li_ t'upL froP Inmm rlrodltlt.e X in tile 02W hundredts and t10.1 trorm produc-r Yo% tile two 1undred,,. lor example, Inatetl-at flfl cOUpon X-1 Was identificid in the data, W tis 103 and material flItn outonY- I as 211.

Ilt

Of the individual alloying elements, V8-Amt. Cr is slightly related to VI 5-Amt. Co, related to

VSO-V7xV9-NixCr, the product of Amt. Ni times Amt. Cr, a variable representing an interaction effect, andhighly related to V79-:V8xV8-CrXCr, a second order effect in this variable itself. Table7 shows that many ofthe interaction effects and higher order effects using a given variable are strongly correlated to that variable andmy not be capable of being properly evaluated using these data. For example, the effect of total aluminum maynot be separable from tile effect of acid soluble Al in these data, VI I and V1 2. Similarly, ultimate tensilestrength, V27, tensile yield strength, V28, and compressive yield strength, V31 are so strongly correlated as tobe difficult to evaluate individually, and thus each must be regarded as just a different way of quantifying the"strength" of the sample.

Of interest in Table 7 is the failure of most of the Charpy V-notch properties at +210 F, -- 120 F and evenat the NDT temperature itself to correlate with the NDT transition temperature of this steel. It is not surprisingthat Gross2 was able to obtain only a ±50 F correlation between Charpy V-notch and drop weight NDT transi-tion temperatures for the variety of steels he was using, since the Charpy transition properties measured in thepresent study did not correlate very well with NDT temperature. Of the Charpy impact properties, only V36- 120 F Pct Fiber has a correlation coefficient greater than 0.47, and it was 0.82. Gross did not try to correlatethis with NDT. The present study was not able to relate Charpy and NDT transition temperatures even though

this steel (a) was a single grade, (b) had carefully controlled heat treatments, (c) had the drop weight specimencrack starter bead deposited before heat treatment to eliminate tle effects of a heat affected zone under thebrittle bead, and (d) that the Charpy specimens were taken from the drop weight specimen used to measure NDT.

SINGULARLY SIGNIFICANT VARIABLES AFFECTING NDT

The correlation matrix can perform an important function for the data analyst by giving an indication ofthose independent variables that by themselves have a linear relationship with the dependent variable. Thecorrelation matrix, Table-7, showed that only three variables have a fairly high linear correlation with NDTtemperature; they are:

I. V20 Prior Austenitic Grain Size, r = -- 0.78

2. V36 -- Percent Fibrous Fracture at -- 120 F, r = -0.823. V64 - Ganma (;rain Size x TYS, r -0.77

"lhe next seven best variablcs have correlation values with NDT temperature ranging between 0.4 and 0.48; therest of the variables were lower. These seven variables and their values were:

I. V33 -- Cv Energy absorbed at the NDT Temperature, r = 0.48

2. V21 - Percent llainite,r = 0.47

3. V24 Percent Martensite,r =--0.43

4. V4 -. Amount of Phosphorous, r = 0.4 I5. V37 C- (E, Fnergy Absorbed at +210 F, r = 0.41

6. V61 TixN 2 r =--0.417, V45 ('rx N2 . r = 0.40

The correlation of ,t'le of these factors with ND'T temperature is understandable. V20 Prior GammaGrain Size has been reported in the literature to be a significant variable affecting the transition tenperature of

120

tempered martensitic and bainitic steels, It, 12, 29, 36, 37, 4, 44 as has strength. 7 , 35, 38, 42, 44,'45 However, the

correlation between NDT temperature and the interaction between austenitic grain size and tensile yield strength

(TYS), V64, is more than likely due in this case to the dominating effect of changes in ASTM grain size number

on a relatively constant 106 ± 10 kst tensile yield strcngth and, as shown later, will not' pIrove to be statistically or

practically significant in the multiple variable analyOs. The Charpy V-notch fibrous fracture appearance at - 120 F,

V36, has a strong correlation with NDT temperature. Tt is logical, since - 120 F is close enough to the NDT

temperature of most of these samples for the percent fibrous fracture appearance to give a good indication of the

change in impact ductility. It is interesting to note that the Charpy V-notch energy absorbed at - 120 F, V35,

has only r -0.37 in the correlation matrix, reflecting the fact'that the producer's original mill rolling practicehas introduced a large variability in the energy absorbed and markedly reduced its level'of correlation with the

NDT lemperature.

The seven variables with correlation values between 10.40 and 0.481 are obviously not dominant factorsaffecting NDT temperature and require evaluation in terms of multiple regression to see if they still contribute

significantly to the observed variation In NDT temperature when they are included with a more dominant factor

such as prior gamma grain size. As will be subsequently shown, the inclusion of a singularly dominant variable

in a multiple regression analysis will eliminate almost all of the variables that show a markedly lower singular

correlation.

RESULTS OF REGRESSION ANALYSIS ON NDT

Almost 100 computer analyses were made using the regression analysis program XIRE and the statistical

methods described in Reference 43 to evaluate the functional relationship between the NDT temperature, V3,2'

and the other 82 variables listed in Table 7. In the initial computer analyses about 15 variables selected for their

mutual independence were selected to determine their functional relationship to NDT temperature. The CharpyV-notch impact values weit! not included since they are just another aspect of the transition from ductile to

brittle impact fracture characterized by the NDT temperature itself. One criterion for selection of significant

independent variables was that the term coefficient, Bi in the computer output, consistently demonstrate that it

was statistically significant, i.e., not likely to lie zero in several regressions with a variety of other variables. As

described in Reference 43, this requires that the term regression coefficient divided by the standard deviation of

the coefficient, Tin the output tables, be greater than some value of "t," Student "t," based on the degrees of

freedom* (DF) for that particular run andI confidence level of at least 90 percent. Because of the large sample

size (here 115 data sets) the usual value of acceptance was about ITi > 1.70 since small changes in a large value

DF do not affect the value of "I" much. Another requirement for acceptance was that additional terms should

significantly improve the functional relationship as evidenced by statistically significant improvements in the

multiple correlation coefficient R and the standard deviation S x.y, i.e., reduce the residual variance not accounted

for by the terms in the regression equation. It was noted that if an additional term or terms significantly im-

proved R it also significantly improved Sx.y; see examples 4-1 1 in Reference 43.

*)1; - Satmpl %i/t" mini0 numntbtr of Iermls In tyrcel~on, including the dependent v:.r ia htc,

'21

Composition and Producer

The results of many regression analyses can be summarized by sayirg that most of the alloying additionsand alloy interaction terms failed to indicate a consistent and statistically significant relationship to NDT temper-

ature. This is not surprising since the range, tnininmum to maximum, of each element present is a result of com-

mercial melting practice specifically aimed at trying to stay within procurement specification limits* as opposedto an artificially large range as would be the case in laboratory melts designed to explore the effect of a specific

alloy addition.

The only element that displayed a measurable effect was hydrogen; Table 2 shows that it ranged from 0 to

12 ppm with a typical value of 0 to 2 ppm. As a result, even though hy~trogen content is statistically significant,

its totnl effect is small, about 2F/ppm 112, resulting in a 0 to + 24 F change in NDT transition temperature.

The producer also turned out to be a statistically significant variable. The producer was coded as 1 forproducer X and 2 for producer Y for the regression, and the results indicate that the NDT temperature of mate-

rial made by producer Y is about 24 F less than that made by producer X, everything else being equal. In thesample of 115 data sets, 76 represent material made by producer X and 39 represent material made by produ.cer

Y. The reason underlying this observation is not simple chance and may be due to either some compositional

difference, such as amount of P or Cu which by themselves were not consistently significant, or some difference

in mill practice. In any case, the summation of differences is great enough to make the producer a significantvariable that must be included in the regression analysis equation.

Prior Austenitic Grain Size

As indicated in Table 4, the prior gamma grain size represented in this sample ranged from ASTM 9.5 to

ASTM 3, and was varied by changes in austenitizing temperature and holding time, in this study 1640 F for

1/2 hr and 2000 F for I hr, as well as by changes in composition from heat !o heat. The regression analysis

confirmed the importance of grain size reported in the literature and showed that adding a second order grain

size term did not significantly improve the correlation.

It is interesting to note that these results indicate that the Petch linear relation between the Is t-lb

Charpy V.notch transition temperature T and ferritic grain diameter d, i. e., T = A0 + A 1 In D- i:, of ferritic

,teels has a linear corollary for martensitic steels and drop weight NDT temperature. 30 , 46 As shown previously,

the Petch relationship can be expressed as a linear function of ASTM grain size (N), T = C + D (N). In terms of

NDT temperatur,' and prior austenitic grain size, it can be written NDTT = B+ + B, (ytGS); this is Il, form ofthe regression results given in Table 8. The correlation matrix and the regression analysis both indicate that this

simple linear relationship will acount for about 80 percent of the observed variability in drop weight NDT

temperature of the I IY.80 steels studied.

An additionzl factor that must le considered when discussing prior austenitic grain size is that final gamma

grain size prior to (uenching reflects the interaction of a number of variables such as:

1. Ausilenitizing lemperature. The higher the temperature above the upper critical, the coarser will bethe gainina grain sizc with controlled grain size steels tending to retain finer grain sizes until their grain coarsening

lemp,.rature is excce(led.

"|IY-,Mo %iect is nelted to ASIM A 543-65 and MI1,S- 6216G (SI III'S).

0')

2. Compositional variations. If, as in the present case, the austenitizing time and temperature are fixed,composition changes will cause the difference between the upper critical temperature (Ac 3 ) and thc austenitizingtemperature to vary slightly. Composition changes that increase this difference will result in a coarser gammagrain size. Increasing C, Mn, Ni, Cu, N2, B, Cb, and Zn or decreasing Cr, Si, M6, Ti, W, V, Sn, Al, P, and Zr willtend to increase the prior austenitic grain size if the austenitizing treatment is held constant for a hypoeutectoid

steel.48-54

3. The longer the time at the austenitizing temperature, the coarser the gamma grain size, especially attemperatures above the grain coarsening tempera'ure for controlled grain size steels.

4. Deoxidation practice or addition of grain refiners such as Al, Si, V, Ti, Mo, W, and rare earth addi-tions will affect or inhibit gamma grain growth.

Such compositional variations would account for the range in gamma grain size resulting from the 1640 Ffor 1j2 hr austenitizing treatment; Table 4 indicates that gamma grain size ranged from 7 to 9.5 for this treatment.No atempt was made to pinpoint a specific compositional variation for this difference in grain si.:e, but a quickreview of the chemistry data in Table 2 indicates that the two heats that resulted in grain size 7 (X-3 and Y-1 I)had the lowest nitrogen contents, 0.003 and 0.002 percent, respectively versus an average value of 0.007 percent.Whether or not this is the underlying cause for the difference can not really be settled from these experiments.

However, it has been reported in the literature that Al additions combine with nitrogen to form aluminumnitrides that inhibit grain growth. 52

Microstructure

The previous studies reviewed herein showed that the tempered martensitic structure generally had thelowest transition temperature and that bainitic structures are intermediate to those of the same steel treated toproduce martensite or ferrite. Hafl indicates that grain size has an cffect on transition temperature of Ohe samco(der as that of degree of hardening.:5 The data in Table 5 show that the steels tested had been treated to pro-duce from 0 to 69 percent bainite with the balance martensite and from 0 to 67 percent ferrite with the balancemartensite, except in four cases where there was some pearlite formed during the final part of tie 1200 F Ireat-

ment. Up to 32 percent pearlite was found.

Table 9 gives the regression results for the effects of percent microconstituents on NDT temperatureallowing for the effect of prior austenitic grain size, expressed as ASTM grain size numbr; 100 percent temperedmarlensitt: is the implicit baseline. Linear regression terms are used since higher order and interaction terms werenot found to statistically improve the relationship. The results show that: (I) bainite raises tie NDT transitiontemperature about one-half a degree per percent; (2) ferrite lowers the NDT transition temperature about three-fourths of a degree P per percent; (3) pearlite raises the NDT temperature about 5 F per percent; and (4) refiningthe prior austenitic grain size lowers the NDT transition temperature about 20.4 F per ",nit increase in ASTMgrain size number. Thus, one sees that changes in prior austenitic grain size expressed in terms of ASTM grainsize have an effect on tile NT)T temperature that is about an order of magnitude greater than the effect ofpercentage changes in microconstituents of the I IY-80 steel studied here,

A quest'on might arise as to whether there is a change in controlling effect on NDT temperature by tiletype of grain size, i. e., prior austcnitic or ferrltic, as tile amount of isothermally produced ferrite incteases.

This question can be answered either by metallographic examination o1 the fracture path in tle broken specimens

2.............................

or by a graphical analysis of the difference between experimental and calculated values of the NDT temperatureas a function of the ferrite content. The latter method has been selected to answer this question since such agraph can be quickly prepared from the back calculation data in the regression analysis compbter output, where.as the former would require the laborious preparation of a series of metallographic specimens.

Figure 7 is a plot of the difference between the experimentally observed NDT temperature and the calcu-lated value YDF = 1 - Y(7 using the regression results of Table 10 for those specimens which contained a meas-urable amount up to 69 percent of ferrite. This shows that there is a slight tendency for the calculated value ofthe NDT temperature to he high, but this tendency is essentially confined to those very fine prior austenitic grain,8.5 to 9.5, specimens conta ining only a trace of ferrite. Examining this figure shows that there does not appearto be a systematic difference between experimental and &c fulated value as the ferrite (and pearlite) is increasedto large amounts, until as little as 8 percent martensite remains in a matrix of 69 percent ferrite and 23 percentpearlite. See Figures Cl, CI 6 and C17 in Appendix C for the photomicrographs of th, predominately ferriteand pearlite inicrostructures.

judging by the results ef regression analysis of these data, it is fikely that prior austenitic grain size remainsthe dominant factor controlling NDT temperature of this steel, even as the amount of ferrite is increased to over50 percent and significant amounts of pearlite are found. This is confirmed also by the relative magnitude ofthe term coefficients for prior austenitic grain size and percent ferrite in Table 10, which show that a change ofone ASTM grain size number has about 100 times as much effect in determining the NDT temperature as achange of one percent ferrite, - 19 F/ASTM G. S. number versus -0.2 F/Percent Ferrite.

No attempt was made in the present study either to deliberately vary ferritic grain size or to deliberatelyvary the prior austenitic grain size before the 1200 F transformation, even though the prior gamma grain sizeranged from ASTM 7 to 9.5. While Geil et al.4 have reported some data on the relativi effects of ferfite andprior iamma grain size on the Charpy transition temperature of 0.3 C steels; they came to the conclusion thatferritic grain sic is rmore important. It must be pointed out that they only changed the gamma grain size from50 percent ASTM 4 plus 50 percent ASTM 6-7 to ASTM 7-8 with few 4-5 in three heats of an unkilled steel.Neither of these experiments has varied the prior gamma grain size enough to say with certainty that either pre-dominantly controls transition temperature in ferritic steels. A definitive set of experiments remains to be doneon the relative effects of prior austenitic grain size and ferrite grain size of the ferritic steels.

Inclusion Content

Examination of Table 3 and Figures 5 and 6 shows that the inclusions observed and measured in thesesamples were small and uniformly dispersed in a way that should not have affected the measured NDT tempera-ture. This was confirmed by the regression analysis, which showed that the Inclusion content of these samplesdid not correlate with the observed variations in NDT temperature of the IIY-90 steel studied.

Strength

Reference 44 .Jhmws that the strength level of IIY-80 steel, varicd by tempering at various temrcfatures forone hour and then quenching, has a strong correlation with the NDT temperature of the steel. The NDT temnoera.ture mcreases as the strength 1% increased. The tensile yield strengths measured for the 115 data sets that mallip the iafnple being analyzed have an average value of 106 ksi and a standard deviation of 10.2 ksi. Over such a

2.4'

iestricted range a linear approximation should adequately describe the effect of strength on NDT temperature,'.en though a higher order function would be required for a wider range.

Table 10 shows tile results of inear regression analysis for the effects of producer, gamma grain size, per-

cent microconstituents (bainite, ferrite and pearlite) and tensile yield strength in psi. These results show fortht s, da:a, aftcr outiting for the other variables, that tensile yield strength contributes to determining theNDT temperature only in a minor way, about 0.68 F per 1000 psi increase in tensile yield strength. This is

slightly less than tile 3 to 4 F per 1000 psi shown graphically in Reference 44 for the range in tensile yield

strength (TYS) from 100 to 120 ksi for coarse and fine grained material respectively. The reason for this

difference in AT/A TYS between the two studies is that this study covers a slightly wider range in tensile yieldstrength, from about 86 to about 126 ksi, which includes the range from 80 to 100 ksi that Reference 46 shows

has little or no effect on NDT temperature, as opposed to the 100 to 120 ksi range where tensile yield strength

has about the maximum effect on NDT temperatures.

Analysis of other computer runs shows that higher order strength terms and terms expressing an interactionbetween austenitic grain size and either strength or microconstituents do not significantly improve the correlation

for these sets of data Correlations based on ultimate tensile strength (UTS) show no statistical or practical im-

provement over those based on tensile yield strength.

The data in Reference 44 shows that strength level, as changed by variation in tempering temperature, has

a larger effect on NDT temperature than is indicated in this study. This is due to the fact that it, Reference 44

the strength was varied by tempering for one hour at temperatures ranging from 400 to 1330 F to produ~ce a

tensile yield strength ranging from 161 to 81 ksi in the tempered marlensitic structure alone; in the present study

tile tensile yield strength ranged from only 107 to 122 ksi in the 100 nercen! martensitic structure tempered atI 150 F. Reference 44 also shows that the functional relationship between strength and NDT temperature can-

not be represented over tile full range of -trength by a simple higher order polynomial and thus does not lend

itself to regression; in fact, it would appear from Reference 46 that even higher order interaction terms are re-

quired to represent the interaction between strength and microst ructure. Such a study is not warranted here

since the graphs flown in Reference 44 clearly demonstrate the complexity of this interaction.

Summary of Factors Controlling NDT Temperature

Figure 8 gives a series of graphs showing various aspects of the function relationship between tile signiti-

cant variables and Nl)T tetoperalure of I IY-80 steel given in Table 10. The range of each variable shown in these

plots is limited to approximately the range of the input data used to develop this relationship. These plots are

intended to give the reader a visual understanding of the significance of' the relationship and a feeling for the

relative effect of small cl!:!.ges in each variable. Figure 8a shows the functional relationship between N )T

temperature and prior austenitic grain size in termis of microconstitnenlts and produlcer when the tensile yield

strength is fixed at a typical value, 100 ksi. Figurc 8) shows the functional relationship between NT)T tempe..1-

ture and relative armounts of nonmartensitic microconstituents in terms of ptio austenitic grain size tor one

produccr and a fixed stiength level, I00 ksi. In Figure Xc tile relitionship belween tensile yield strength and

NI)T temperalurt is shltwm int terins of anount of nonmartensitic microconstituents and prior austclitic glaill

site; the range in tensile yield stlrength has been limited to th, range observed experimentally it Table (i.

This study has shown that under carefully controlled condition, which eliminate the possiwility ,it retained

,ustenltte and of tellper cnrittlemnent (tie to slow Cooling after tertiperiltg, tile (Iolninant factor r'iitolling. time

2.5I

NI)T temperature of a low-carbon, Ni-Cr-No steels* such as HY-80 is the Prior ausienitic grain size. The resultsin Table 10 and Figure 8 show that prior austenitic grain size has from fo'ur to 100 timtres, s much effect on NDTtemperature a% either the percent nonniartensitic isothermal products, petcent bainitejferrite and pearlite, in atempered marteirsitic matrix or the tensile yield strength. The significant Vari3ables and the relative eff-ect the)'had (in the ND? temperature of the Ni-Cr-Mo steel used in this study are:

Variable Range Effect on NDT

Producer (Coded I or 2) -21.9 FPrior Austenitic (3 to 9.5) -19.2 F/ASTM

Grain Size Grain Size Number -

PerentBanit .0 to 76 percent), 0.62 F/PercetBPercent Ferrite (0 to 69 percent)' - 0.23 _F/Percent F_Percent Pearlite (0 to 32 percent) 5.4 F/Pcrcent PTensile Yield (61.5 to 124.7 0.68 F/Ksi

Strength ksi)The regression analysis equation is expressed miathemnatically tinder Table 10.,

hestandard deviation (SVx.y,) of this correlation is 18.3 ICand the multiple correlation coefficilent (R) is 0.90with an associated degrees of freedom of 108. Higher order terms and interact ion terms did not significantlyimprove the correlation.

The value of the standard deviation for !he correlation reported here, Sx. v = 18.3 Fis close to the theo-retical limit of the best possible correlation based on the ± 10 F repeatability of the drop weight test determina-tion of tire NDT temperature and the repeatability of measurement of the supposedly fixed valued independentvariables.** For example, a repeatability of ± 1/2 in the ASTM prior gammna grain size measurements wouldindicate that the hest possible valuec of Sx..y for the regression would be about ±15S F. Because of the small con-tribution of the other variables to NI)? temperature variation, smiall errors in their measurement would not con-tribute as mujch to the magnitude of the standard deviation for the rcgression as would an error in tfie measure-inent of the prior austenitic grain size, flowever, tile effect of the errors associated with each of the independentvariables plis the variation in fihe experimental determination of the NDT temperature indicate s that the value

of the standard deviation associated with thle final repression equation given in Table 10, S.y = 18.3 F, is closeto thre theoretically best possible functional telationship between a set of variables and the NIYI temperature;lower valuies can be altained hut only by using a set of (lummy variables that would have no true significance.

The producer who melted the steel and rolled out the plate is a. significant variable when *lnalyzing thevariables affecting NI)' temperature. This is attributed to basic (liffercnices in composition and mill practice that

could niot be qutantiflcd by themselves but are lumped together in this variable.II

No chiemical composition lermns (other than amtount of hydrogen) wr, tound to consistently correlate

with NDT temnperat tires. and the low level of hydrogert found in these tests, typically 0 to 2 ppmn, results in

OASTMk A 343-65 and Mlt.-1-162l16G (Su11t'S).

*'Regre4-tion ana.lysis pr ganiituclh ai fle~ one usied to ~inalyze thewc (fain43 insure the error is associated with the depen-dent v;.rtable and thit tile independent vaerlahles ate cxact values. White thlk is not tile true ease, it is a commttonly used mtirimp-tion, 'fil more rtgrnromz technique has computational and Aiialyticai disadvanitages that are dkimsed by (Oo&% in Refreince 21,

little practical effect on NDT temperature, 2 F per ppm H2. The failure of other alloying and trace elements tosignificantly affect NDT temperature is attributed to the effect of these elements'on the prior austenitic grainsize which in itself is the dominant factor controlling NDT temperatures n.easured In these tests. In addition,the alloying and trace elements affect the microstructure and strength through their influence on quantity andtype of transformation product produced by austenite decomposition, ferrite strength, temprability, secondaryhardening and susceptibility to temper embrittlement, and through their effect on strength and microstructureexert an effect on the secondary factors that determine NDT temperature.

The steel used in these studies was relatively clean and free from inclusions. Inclusion content was notfound to be a significant variable affecting the NDT temperature.

'he range of tensile yield strength produced by this study was rather limited by the decision to use a single(I 150 F) tempering treatment for all samples; the mean value was 106 ksi and ranged from 124.7 ksi for thehighest strength teoipered martensitic structure to 61.6 ksi for the lowest strength structure containing largeamounts of tempered ferrite and pearlite and the least tempered martensite. Reference 44 indicates that therather small effect of strength on NDT temperature (0.68 F/ksi) is related to the limited range of strengths ob-tained in this study. Reference 44 indicates that there should be an interaction between strength level and mici'o.structure in their effect on the NDT temperature; this was not observed to be the case in the present study, again

probably due to the limited strength range.The order of magnitude of the effect of prior gamma grain size on the NDT temperature found in this

study, -19 Ff(-yGS), is of the same order of magnitude as that reported in the literature for both ferritic,bainitic and martensitic steels, in terms of ferritic and prior austenitic grain sizes respectively. In plain carbonsteels a unit increase in the ASTM ferritic grain size, AaGS, will reduce the transition temperature 22 F on theaverage with the change ranging from +9 to -39 F.2,6, 7, 13, 17, 18, 24, 28,31 Drop weight NDT test of aferritic steel gave - I I F/AaGS.24 Pearlitic steels 8 (C 1080 and C 1095) and steels air cooled and tempered 3

(SAIE 1045, 2340, and 3140) had ATIA-yGS ranging from - 17 to a- 125 F per unit change in prior austeniticgrain size. In a study of the effect of impurity, P and Sti level, ATIAyGS ranged from -4 to -42 F per unitchange in prior austenitic grain size in a 0.3 C, 2.0 Ni, 0.75 Cr steel with the value increasing with impurity leveland embrittlig treatrnent. 2 .1

The linear dependei,ce of transition temperature of ferritic steels on ferrite grain size, T--.A 0 +'A In D - 112,

reported by 'etch3 0' 46 and of pearlite, 0.84 and 0.95 C, steels on prior austenitic grain size reported by Grossane' Stout 16 is in agreement with the linear dependence of NDT temperature on prior austenitic (ASTM) grain

size observed for the martensite IIY.80 steel used in this study. Irvine and Pickering 27 report that the transitiontemperature of low-carbon hainitic steels increases in an S shaped relationship, third ordeia, with increasing priorgamma grain diameter as varied by reheating to various austenitizing temperatures. Replottfing their data onsmi;og graph palp'r shows that (ie transition temperature of low.carbon bainitic steels also varies in an approxi.inately linear manner withi the logarithm of the grain diameter. The results of the present study indicate that alow.carborn cmpered martensitic steel such as IIY.-80 also has a linear relation between NDT temperature andririor austenitic grain size as expressed in ASTNI grain size numbers. This relationship is not affected by appreci-abl ar;ounts of isotlherrnally produced bainite and ferrite in a structure tempered at 1150 F.

27

CONCLUSIONS

The following conclusions are based on the results of a study on the effects on the NDT temperature of:

(1) commercial variation in composition and inclusion content, (2) variation in mlcrostructure, prior austeniticgrain size and microconstituents produced by quen!hinp and by partial isothermal transformation at 1200 and875 F to produce ferrite and bainite prior to quenching, and (3) variation in strength level after an 1150 F

temper for one hour followed by a water quench to prevent embrittlement while cooling from the tempering

temperature.The steel used in these studies is a high toughness, low impact transition temperature, quenched anj

tempered Ni-Cr-Mo steel of tile type made to ASTM Standard A 543-65 and Military Specification MIL-S-1 6216G

(SHIPS). Sixteen coupons were obtained from one producer and eleven coupons from another. The averagecomposition of the steels used in this study are given below:

0.18 PercentC 0.10 Percent Cu

0.36 Percent Mn 0.032 Percen' Acid Soluble Al

0.016 Percent P 0.040 Percent Total A]

0.019 Percent S 0.002 Percent Ti

0.30 Percent Si 0.014 Percent 02

2.75 Percent Ni 0.007 Percent N2

1.51 Percent Cr .1 ppm 1120.38 Percent Mo

FACTORS AFFECTING THE NDT TEMPERATURE

The rMsultS of this study show that the NDT temperature of the low-carbon Ni-Cr-Mo steel tested is prima-

rily determined hy a single variable and that the other significant variables act in a secondary way. The conclu-

sions drawn from this study are as follows:

1. Using a tempered 100 percent martCnsitic microstructure as the baseline, the significant variaies

acting to determine the NDT temperature measured .in this study and their relative (from Table 10) effects are"

Variable Range Effect on NDT

Producer (Coded I or 2) -21.9 F

Prior Austenitic (3 ;o 0.5) -- 19.2 F/ASTMGrain Size Grain Size Number

Perce|t Bainite (0 to 76 percent) 0.62 F/Percent 13

Percent Ferrite (0 to 6) percent) -- 0.23 F/Percen; F

Percent Pearlite (0 to 32 percent) 5 A F/Percent P

Tensile Yield (61.5 t,, 124,7 0.68 F/KsiStrength ksi)

, '2S

The effect of a uini t change in prior atistenitic grain size on, NDT ternperdtture is one and one-hial f to two orders ofmagnitude greater than the effet of either a percentage change in thle amount 'of isoithermnally produced bainiteand ferrite or a 1000 psi change in tensile yieldl strength.,

NDT Temperature (F) -77 -- 22 (Producer I or 2)- 19.2,(-yfS)

+ 0.62 (Percent B)- 0.23 (PIercent F) + 5.4 (Percent P)

+0.68 (TYS In ksi)

2.A linear function of the significant terms gives the bem. i-xpression of the functional relationship be-tween these variables and thle. NDT temiperature since it was found that higher order and interaction terms didnot significantly improve the correlation.

3. No chemnical composition termns were found to cons istently correlate with measured NDT tempera-tures in a statistically significant mainner, except for the amrount of hydrogen in the steel, and the low lev' of

hydrogen found in these steels, typically 0 to 2 ppti', results ill little Practical effect onl NDT tem'perature, 2 Fper ppmn 112. The failure of alloying andi trace elements to significantly affect NDT temperature is attributed

primarily to the fact that these elements interact vwih: (a) the austenitizing treatment to control prior austen-

itic grain size; (b) the transformation product;, (c) f-tiperabilit~ 'a.fnlsrngho h tel n d ai

susceptibility ito cmhrittlemcent dluring normal teirjx ring, ais well as the fact that ranges in the amount of each

element were so restricted by the specifications and corrmmercial melting practice as to make it impossible to

dletect their relatively weak effect onV a SOUnd statltihml a.s

4. 'fhe incfo~sion :(ontent was riot found to b it significant factor affecting thle NI)T temiperattures

measured inl t hese stidies.

5. The weak correlation between tensile yield strength and NI)T tempe ratures observed in this study.

as opposed to the strong effect reporfed inl R'ferences 44 andi 45 is attributed to thle finited range of strengths

producedl by the ff50 F temper used onl all material in the present study.

6. The linear relationship betwci. ASTM prior austenitic grain size and NI)T temperature of thle low-

carbon Ni-Cr-Mo steel uised( iii this stutdy is similar to the litnear dependence between;. (a) prior amitenit ic grain.27diameter and the t ransitlion tempera t tire of low-cairbon bainific steels reported by Irvine and Pickering,

(b) prior justeni tic grain size and the transit ion tempera tre oif pearlitic 0.84 and 0.95 C steels reported by

(;ross andI Stoult- 26 and (c) the logarithmn of the ferritic grain dliamtnter and thle tran sition Itemperat tire of ferriticsteels reported by Petich.Y)

7. The results of these studlies indicate that thme steel producer is a statistically significant variable that1

Canl accountl for abolit 22 1: of' the observed variation in NI)T lemlperat ore, depending upon which producer

inadfc tile steel; this is attributed to a Combination of small Compositional differences, such as amounit ofi (i'o

Or which by thetmelves were niot consistently significant, and to diffre~nct'% 61~ mill practice. fIn any cawe, these

results intlicate that tile produicer is a1 significant variaible that shmoulmi !-t considered k'i m.y analy sis of variables

affecting If- NI )f icrmperaltire of' low-carbon Ni-Cr-Mo steels.

219

ACKNOWLEDGMENTThis report is based on work carried out by the author in the course of his employment with theNval Ship Research aod Development ('enter, Bethesda, Marylanl. Use of the data as a basis for a thesiswas approved by the University of Michigan and NSRDC.'Re athor gratefully acknowledges the advice and guidance given by Professor M. J. Sinnott andimembers ot his doctoral committee from the University of Michigan anu by Mr. A. R. WHiner of the Center

whose enfcouragement and support made this work possibe. Thanks are given to Mr. Daniel Embody of theBiometrical Service, U. S. Department of Agriculture, lktsvilbe, Maryland, for his advice and guidance on iheuse of reqression analysis programs. The contributions of J. liogber, R. J. Langley and R. Martin and othermembers of the Center staff are especially appreciated.

30l

Direction of RollI

Varying Top Drop WeightsThicn ssScrap

Bottos Drop Weights

1~12"

Figure I - Location of Material Used in This Study

Direction of Roll

'A I . crop burned edgits of coupon for at litt 1/2 inch oll around.k i _ 2. Stencil identification on 5/8", 2 Inch face shown on sitch.17UI 121187 121183 12UU 2I 1 Identification should Include, coupon number (12U, 13U1, tic.),

T top (T) or botto (6), anid ipecimen number (1 -10).

Figure 2 - Liyouf of Drop Welght Test Specimens

Al1

Note that type spec imnShow" as Inner Specimen Itdifferent in each cas NOTE,

I. lair.P Ie following specimens at Itis. locations shown fromeach group Of coupon$?

QMP ~ (A.)?1 longitudinal madified S-914 9.s;t specimens.(a.)?2 longitudinal S-12914 compr.,sskn specimens.)7U!I 12UY2(C.-) 10 transverse 8-328-4 Charpy V-natd. Specimens.coupn I(See shseaf No. 2 For additional InstructionsCouon regarding the Clsrpy Spec mans .)2. Identify each specimen with the (llowing.

(A,) Coupon No. Vl 2UT4 For eitample)(8I.) Directionj (L. o, T)str (C ) Specimen No. 1-.4 (or 6 it necessary)) k x --- -- - k. ')3.

Identify tensile and Charity spCIm ns on both ends by stomping.I2Ukj/4. Identify compressin %pacimans by scribing an the sideCOUON3blUB only. II

Cou~pon 4 5. IF any Cisarpy speciffens , tois .Foune edba(if avalabbeo) they shul oame Fro unupor wel bnda2Figure 3a - Lccatjon of resq Specimlens te hudcm rmCUOIad2

Weld debit t aveyhad

I~~hoiuerc 3brp V-oc Impactk Spcien inou oaccoprdaNnceSpcic

Figure ~ ~ ~ 2 3N oain~f ehnclP ote drecins s Cofrp Vnotch adseinietificacShencsmens MademrrsnioTtsoedcorppnWeiahtawayCfromItS

Figrc bLoDeati oif Lyout~lna of iarpy V-Notcht Specimsens;

Figure - oato o)fltti r Mechncal na Prop ry oTnsee hryV-Notch Imp11ppeccenSeiesMde rom Testedl Dpc Weigh Slocks en

F~igure 5a - Mid-Thickness, Figure 5b -MNid-Thicknes%,Longitudinal Transverse

IReproduced frombest availablecop.

Figure 5c Adjacent to Figure 5d - Adjacent teSurface, Longitudinal Surrace, Transvcrse

Figure 5 M~otomnicrographs of the Inclusion Content of Coupon X-l18

331

Figure 6ai Nfid-'11iicknem. F igure 61) Mid-Tbitckness.Loingitudinal Transverse

Reprducedro7.

1.itv (Ic Adiaccii Io Hu re 6d IN A Ja1-i tolklit fac(, LI.Iugituitial ktir lace, Transverw

FigmeC 0 I'Itttoicrograiphso 1)11 116101 Cousn (intent of Colupon YdI

DATA CALCULATED USING THE RESULTS FROM40 TABLE 10, Sx. y 18. 3F, R O.90

30

20 . 22. 8% P

Figure 7 - A Plot of the Difference between Calculated and10 P F Experimental NDT Temperature of Material Containinglb 2 10 O 40 5 o 0 from a Trace to 69 Percent Ferrite

cc 0 10.7% P versus Ferrite Contento0O0

1 -20 O32. 1% P

-30

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Figure 8a - unctional Relationship between NIT "- .Temperature and Prior Austenitic (;rain Size in -

Terms of Microconstituents and Producer .. i-2- FERRITE

0 25 50 75 100PRODUCER X (CODED I) NONMARTENSITIC PRODUCT. PERCENT

-50 MICROSTRUCT URE20 B + FU I Figure 8b -- Iun, .ional Relationship between NI*

-.-. - -- 100% M 'temperature and Miroc'nsfituenis in 'I'erm o--- ' 800' M Prior Austenitic Grain Size

-I00 . . . ASIM G. S. 3

z -150 . 20% B 4 8M

20% F 80 M Figure 8 -P Plots Showing the Relative Effect of' Prior-200 ASTM y G. S. 8 Ausenitic Grain Size, Producer Percent Nonmnarten-

sitic Isothermal Transfornwiton Produce (Bainite,

60 80 Im 120 140 Ferrite and Pearli Ie), and Tensile Yield St rengl hTENSILE YI1D SIRTIHT, Ksi on the NI)T Temperat tre of" Low-Carhon

Ni.Cr.Mo Alloy Steel TemperedIigure Xe€ - lto n~ l m at Relag ionsti h btIwecn NI )TTemptraftire and le , imle Yicld Sirength in 'lerins at 1150 V

of Mictmwonvtie o int a iMor Aumflenitit(0;rii Sile

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TABLE 4

Prior Austenitic Grain Size

SPECIMEN AS RECEIVED MATERIAL AUSTENITIZING TREATMENTNUMBER MID-THICKNESS 1640 F-1/2 HR-W&Q. 2000 F-i HR-W*O.

X-i1 68.X- 2 68X- 3 5 7X- 4 69.

4 8.5X-6 78X-8 .X-9i 6 8.5X-101 4 9.5X-11 6 8.5X-12, 8 8.5 3X-13 6 83X-18 5 9.5 5X-15, 78.5X-16 5 9.5 6X-17 6 8 05 6

V-I1 78.Y- 2 7 9.5Y- 3 785Y- 4 5 9.5V- 5 6 8.5Y- 6 5 9.45Y- 7 58.Y- 8 7 8.5V- 9 5 8.5Y-10 7 8.5V-11 4 7

*AFTER AUSTENITIZING AND QUENCHING, -120 F I HR.TEMPERED AT 1150 F I HR THEN WATER OUENCHFr,.

.10

TABLE 5

Percent Microconsiituents Produced by Various Isotherma! Trctments

5.STENIT IZING

TEMPERATURE 1640 F 168O F 1680 F 1610 F 2CD0 FTINE 1/2 HR 1/2 HR 1/2 HR 1/2 HR 71 HR

ISOTHERMALTEMPERATURE 87S F 87S F 1200 F 1200 F 875 F

TIME 152 SEC 1600 SEC 3350 SEC 8500 SEC 1600 SEC

MICRO-CONSTITUENTS BAINITE BAINITE FERRITE PEARLITE FERRITE PEARLITE BAINITE

COUPON NUMBER

X- L 29 S4 31 0 67 11X- 2 it 54S TRACE 0 2 0x- 3 21 31 0X2 11 0X- 5 1 0 01-6S 0 1 0X-S 1' 5 0

X- 9 7 26 0x-10 1 TRACT. 0X-11 0 0 01-12 I SO TRACE 0 1 0 63X-13 I' 60 TRACE 0 TRACE 0 67X-1S TRACE 21 0 0 0 0 56X-IS 1 27 0 0 0 0 58X-16 0 16 TRACE 0 TRACE 0 58X-17 0 32 0 0 TRACE 0 30

T- 1 53 so 32V- 2 9 41 22 0 S3 4v- 3 76 69 23y- 4 1 39 7 0 21 0y- S 3 17 0T-.6 1 12 0V- 7 1 3 0Y-8 1 1 0y- 9 6 2 0Y-10 14 1 0Y-11 27 63 3 0 'S 0

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b. la inites will exhibit ragged grain boundaies, with this becoming mrne pronu'nced at. tbe lowertransformation temperatures. in general, it may also be said that the bainites will not exhii~it the large flatblocky areas of the ferrite but will be smaller and peppered with darker areas of martensite or precipitated

carbides; this becomes more pronounced is the transformation temperature is lowered.

3. A smeared structure or an oxidized surface will make proper evaluation of thc results of this tech-nlique imipossible.

APPENDIX C

PHOTOMICROGRAPHS

OF

MICROSTRUCTURES

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APPENDIX D

CHARPY V-NOTCH IMPACT

PROPERTY CURVES

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APPENDIX E

TIME-TEMPERATURE COOLING C URVES FOR DROP WEIGHT SPECIMENS WHEN

TRANSFERRED FROM VS40 F FURNACE TO MOLTEN SALT BATHS

AT EITHER 1200 F OR 875 F

INTRODUCTION

i lthe actual tirne-temperature history wias desired for tile low-carbon Ni-Cr-Mo steel (5/8 x 2 x 5 in.)dryop weight impact specimens used in thesi. tests when they were transferred from a high temperature salt6ath at I (AO F to a salt bath at either 1200 F or 875 F. In order to determine this thermnal history a speci-mi~en was instrumented with a thermocouple at its center as recomimended by Sinnott and Shyne4 and the

t I me-temperature histoiy of the specimen and the bath were recorded at leas,' every ive seconds on a strip

chart recorder. The experimental cooling curve is compared to thle curve calculate(] using the method of

Sinnostt wl'd SIbyne.

SPECIMEN PREPARATION

A O.OX8'.inch dianmcter hole was drilled from one end, 2 1/2 inches deep into thle center of the specimen.

11he hole Was countersunk and tappedI for 1(1 1-28 threads to a depth of 3/ , . A 7 inch length of' 1/4

i high presmure steel tubingy was screwed tightly in to thle hole. An insulatedI 28-gauge Cliromel-AI ualel

thermoIIcollplc was inserted in thle titbe atd forced against thle specimen at the hot torn of the well. Another

ther mocouple was positioned in t he low te mperi t tre salt bath adjacent to thle speci men. Both thermocouples

%C-c at miched(t t a multichatnel recordler to alternately record tile te muperm totres indica ted by both thermo-

':pe' icb spcci mien was sand blasted between each rin to re move all scaile and salt and ret urn it to a

rupreseitative Condition.

HEAT TREATMENT

'lite spec: r;.en was suspenided Verftically in a1 high teltupera lure salt ba th aind allowed to come to 1 640 V,

it wa% held at 1 (,.1 1: 115 I1- for 301 min:utes. mininmum. 'file ±1 1: imha temperat tinre tolerance was used

hvere and for tile low temuperat nrt bat hs hecaluse it re presen ted the tole ranice oif the commercial temperat lire

diilroeemetr used in tim test aii pposed to thle closer ±5 F mlanulal control used for thle actual specinwi lieait

V rritft Pronm 4 tol S Inches (), the thermocouple tuhinp, was kept below lte surf'ace of tile mlolienl vilt

1 iomm/.- fliat losses from'i thme tilhe

1Ifhe p imen W, lt. TVtee(t iio no (on ,i4 bathm to thle low emperatumte Imitb within tine ort two"m4~. file w.;V, rpncmt. W4d ietcmlv Iint center olf tile low teniperatorte bath and mildly

ir Crdnea the th moit ' d i he hmth teipeiratte. NMul1ctipl is Wmw imde to allow fori

d AT 0rrmtl!,0 li theli 1"ot;1 'prime h ei runs.

RESULTS

The results of the cooling rate studies ate plotted In Figure El and are compared to the salt-bath cool-

ing iate predicted itsing thle instantaneous film coefficient method of Sinnot and Shyne4 7 with the following

Specimen Area 0,200 S(l ft

Specimen Weighit 1.771 lb

Mean Specific Heat,1640O to 1200 F 0.127 BTU/lb/F1640O to 875 F 0.128 BTU/lbf F

Table ElI is a listing of thle computer program used to predict cooling rates.

The variable names used in this program and their significance are given below:

ITO Temperature of low !timperature bath, 875, 900 or 1200 F.

ITS Initial temperature of specimen, 1640 F

[IF Instantaneous film cocfficicnt based on thc difference between thc average

surface temnperaturi; ot thle specimen and the bath during the fline inter-

val THETA.

QTI I Average cooling rate in B3TU/see when cooling from MIS I to MTS2.

THETA Time to cool fromt a tenipcrature MTSI to MTS2, seconds

StIIMTI I Total elapsed time to cool from ITS to MTS2, seconds.

1-.,V ~ *: 16si !I indicates reasonable agreement hetwcen experimental and predicted

tcooling rates. The low bath temperature experiments were run with an initial hath temperature of about

900 1: rather than 875 1: Iecatuse of the characteristics of' the commercial salt mix that was availahle in the

sUlf 1b3t11 when these lets were run.

The data show that it takes about 125 seconds for the center of thc specimen to cool fromt 1640O F to

1210 1; and about 140 second% to cool from IW F4) to 885 1- when a drop wcight specimn! is transfe rred

in a 12WX 1: and an 875 F bath respectively.

lI#Wed $inl 11.01.1 !tomr 1 S. NSirrt Cof (OIf lion,

11I0

BATH TEMPERATURESiYMBOL INITIAL FINAL

o 1625F 12006a 1630F 119OF" _ 1630F 1210F

1400 x 1635F 1190F

u -1 16406 12006'.200

0 20 40 60 80 O 120 140ELAPSED TIME, SECONDS

Figure Ela - Cooling Rate from 1640 F to 1200 F(Average Cp 0. 127 BTU/Ib/F)

.. BATH TEMPERATUREGSYMBOL INITIAL FINAL

Vo 16406 910Fi~ o a0 1633F 905Ft 1400 ---- 1640F 91OF

0 -no 16406 875F~ 1200 \0

'0 0

; - --- c"- - - o

X C

0 20 40 60 80 100 IM 140[LAPS[D TIME, SECONDS

if;l.,re l h -- cooling late from 1640 IV to 875 F

(Ave% % ,%; Cp 0. 128 IlI'U/Ih/F)

1l'guic I -I (.omparison of Actual Ccoling Rates with Computed

Cooling Rates l'or the Ceiter of a 5/8 x 2 x 5 Inch

Drop Weight Specimen

117

6W

U 4'

4'M

inca Cf %

PeNL w L;z 1: H~

CC C' 2I4 e% F

% I C>) I 2

1-- 20 41C4W C -- .

-O (/ (L i 1

*1 Z : HO-.4ma 114O 9-'c'4 H -

a) V- V4 -4V- V4'4r4P4 4Cj 0%4-4Hr.4 . .

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