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AMMRC CR 68-09(F)

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INFLUENCE OF ALLOYING ELEMENTS ON THE TOUGHNESS

CF LOW ALLOY MARTENSITIC HIGH STRENGTH STEELS

Final Technical Report July 1967 to September 1968

by

C. Vishnevsky E. A, Steigerwald

•5>

November, 1968 D D C

TRW Inc. \jj ' Cleveland, Ohio fp OfC| ^ fcft}

Contract DAAG H6-67-C-017I

This document has been approved for publication and sale; its distribution is unlimited.

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ARMY MATERIALS AND MECHANICS RESEARCH CENTER WATERTOWN. MASSAGKUSETTS 02172

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INFLUENCE OF ALLOYING ELEMENTS ON THE TOUGHNESS

OF LOW ALLOY MARTENS IT IC HIGH STRENGTH STEELS

AMMRC CR 68-09(F) Final Technical Report

July 1967 to September 1968

by

C. Vishnevsky E. A. Steigerwald

November 1968 TRW Inc.

Cleveland, Ohio

Contract DAAG W-^-C-OI?! D/A Project ICOaMjOIAS^ AMCMS Code 5025.13.S^Z

Materials Research for Specific Army Requirements

This document has been approved for publication and sale; its distribution is unlimited

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Army Materials and Mechanics Research Center Watertown, Massachusetts 02172

ABSTRACT

This study has examined the effects of various elements on the notch bend fracture toughness and Charpy impact behavior of a 0.35%C, 3%Ni, Cr-Mo-V martensitic steel having a room temperature yield strength of approximately I6O-I8O ksi. A classical approach was used in the design of alloys which permitted a direct evaluation of single element effects rather than interactions.

The elements C, Mn, Si, Cr, and Mo raised both the notch bend fracture mode transition temperature and the Charpy V-notch transition temperature (100% fibrösity criterion). In amounts above that required for deoxidation and grain refiietnent aluminum degraded the transition temperature and toughness slightly. A minimum toughness occurred at a vanadium content of 0.1%. Over the entire range of compositions examined (1.26 - 6.231) nickel decreased the transition temperature and improved toughness at the lower test temperatures.

Charpy shelf energy, Cv (max), and fracture toughness, ^max. (at 750F), did not correlate well with work hardening oxponent (n). Good agreement was obtained however when these parameters were plotted versus true frac- ture strain. At -3210F, toughness was essentially fracture strain inde- pendent suggesting that a critical strain criterion based on fracture strain is valid only when fracture is fibrous.

A comparison was made of measured K|r calculated from tensile data using the Hahn-Rosenfield model K|c =(2/3EYe*n

2)'/2, where E = Young's modulus, Y - yield strength, I" = true fracture strain, n = work hardening exponent. The results indicated tha". the increased contribution on non- ductile fracture which accompanies increases in strength and/or decreases in test cemperature in. low alloy steels can lead to large errors in the predicted toughness.

FOREWORD

This report, TRW ER 7217-1, presents the final results of a program performed by the Materials Research Department of the TRW Equipment Laboratories for the Army Materials and Mechanics Research Center under contract DAAG 46-67-C-0171. The work was conducted by C. Vishnevsky and E. A, Steigerwald. F. R. Larson was program director for AMMRC.

[ An earlier phase of this program, "Literature Survey on the Influence of Alloy Elements on the Fracture Toughness of High Strength Steels," was published in February, 1968 as a separate repor t AMMRC CR 67-13(F).

Prepared by: c. W nevsky C. Vishnevsky

E. A. "Steigrf^ald

TABLE OF CONTENTS Page

I. INTRODUCTION 1

II. MATERIALS AND PROCEDURE i»

Ml. RESULTS AND DISCUSSION 16

Effects of Alloy Additions 17

Carbon 17 Manganese 29 S i 1 i con 29 Chromium 29 Nickel 29 Molybdenum 29 Vanadium 30 Aluminum 30 Summary ..... 30

Relation of Toughness to Tensile Properties 30

Work Hardening Exponent 33 Quantitative Models '»I

IV. SUMMARY AND CONCLUSIONS *»7

V. LIST OF REFERENCES k8

VI. APPENDIX 50

I INTRODUCTION

Although nunerous high strength low alloy steels are currently utilized in applications requiring good toughness, relatively little systematic information exists on the effects of composition on tough- ness properties. The major factors contributing to this Situation are the complexity of steel compositions, heat treatment and structure, the variety of toughness tests which complicate efforts at establishing meaningful cross correlations between parameters, and in many instances an insufficient appreciation for relationships between laboratory tests and service performance.

An important problem area involving low alloy steels is gun tubes for 175 mm cannon. At the present time no compositional require- ments are specified for this application. Instead, certain yield strength, ductility, and Charpy impact test requirements are used with th^ chemical composition being the responsibility of the commercial steel producers (1,2). The normal yield strength of these steels is in the range of I6O-I8O ksi The application is particularly severe and cracks usually initiate at rifling marks after a few rounds The propagation of these cracks under cyclic stresses poses a serious prob- lem in reducing useful tube life by promoting catastroph.c failure. To date, effective methods of eliminating the early initiation of cracks have not bean developed. The problem Is therefore one of controlling the rate of crack growth and minimizing the probabi1;ty of catastrophic bri ttle failure

The gun tube problem involves crack growth in th.ck sections (i.e , undarplane strain conditions), and lends itself to analysis by linear elastic fracture mechanics. The understanding of fracture toughness is most advanced regarding crack propagation under plane strain conditions and test methods are available for measuring the plane strain fracture toughness, (K. ). In addition to defining critical crack size under tensile loading,experimental data exists for gun lube steels which sh_ that the rate of fatigue crack propagation is inversely related to the K. (3). Thus, a large value of K indicates that fatigue crack growth rate is decreased and a larger crack can be stable before the onset of rapid crack propagation.

The Charpy V-notch impact test, however, provides a less direct evaluation of toughness as required for this application in particular, it measures both energy for crack initiation and propagation in a section which is appreciably th nner thar the gun tube The presence of shear lips in broken Charpy specimens does not accurately reflect the flat, macroscopica1ly shearless, crack propagation in thick sections. It is also important to note that the Charpy impact test, which has been

ow

valuable in toughness studies of low strength steels, becomes in- creasingly insensitive as strength increases. It is thus reasonable to expect that an analysis of steel composition requirements for large gun barrels using recent fracture mechanics knowledge will provide more insight into material behavior than approaches depending on conventional toughness tests.

The present program was divided into two stages. The first con- sisted of a review of the literature on the effects of alloying elements on toughness of low alloy steels. The results of this survey have been presented in separate form CO. The conclusions of this work were as follows:

1. The low alloy matensitic steels should be used with the lowest possible carbon content consistent with harden- ability and strength level requirements.

2. Oxygen, nitrogen, phosphorus, sulfur, and the tramp elements are all detrimental to toughness and should be maintained at minimum levels.

3. Nickel is an alloying element which improves fracture toughne5S.

k. Manganese is an alloying element which has a detrimental effect on the fracture toughness of the high-strength steels.

5. Although the data are not systematic, the carbide formers, chromium, molybdenum, and vanadium, produce no consistent effect on fracture toughness when added in the normal range of compositions used in the high-strength steels. Molybdenum, however,can be beneficial for minimizing the effects of reversibla temper «mbrittlement.

6. Silicon effectively retards the tempering reactions which promote irreversible (500oF) embrittlement. When used with vacuum melting, silicon improves the fracture toughness of ^»3^0 steel at the high-strength levels.

7. Fracture toughness is improved when the desired hardenabi 1 ity is obtained by a combination of alloying elements rather than a single alloying element, provided that nickel is one of the principal additions.

8. The superposition of a secondary hardening mechanism which involves precipitation of copper, aluminum, or complex carbides in a 0,30 - 0.^0% carbon martensite does not improve the fracture toughness at the higher strength levels. At the lower strengths however (below approximately 215 ksi tensile) there may be some advantage in the use of a pre- cipitation hardening mechanism involving copper or the use of secondary hardening with complex alloy carbides.

These conclusions suggest the validity of applying certain general principles to alloy steel development from a fracture toughness viewpoint, They cannot be applied rigorously to any specific alloy composition but rather serve as guidelines for more detailed work on particular steel types.

Based on the results of the above literature review, an experi- mental program was adopted which permitted the systematic study of compositional effects on toughness in a typical gun steel having a room temperature yield strength of 160 - 180 ksi. The base steel chosen for this portion of the investigation had a nominal composition of 0.35%C, 0.65^Mn, 0.35%Si, 0.85«Cr, 3.0^Ni, 0.30^Mo, 0.10%V. The effects of all these elements together with aluminum were examined by varying individual elements around the base composition. This classical approach permitted adirect evaluation of single element effects. The objectives of this portion of the program were to define the in- fluence of the various alloying elements on fracture toughness as a function of test temperature and explore the possibility of relating standard tensile properties, as affected by composition, to fracture toughness. The program also sought to establish correlations between fracture toughness and the conventional Charpy V-notch impact test.

I MATERIALS AND PROCEDURE

A rotal of nineteen experimental heats of steel were evaluated. The list of compositions in Tabie I shows two base heats and seventeen other steels in which individual elements were varied deliberately from the base composition. Three compositional levels were obtained for C, Mn, Si, Cr, Mo, V, and Al while four levels were examined for Ni. These are listed below:

Element Weight %

c 0.28, 0-35% 0.4l Mn 0,21, 0.65--, 1.51 Si 0.06, 0.36'-, 1.38 Cr 0.51, 0.82*. 1.61 Ni 1.26, 2.83^. 4.45, Mo 0.13, 0,30*, 1.04 V <.öl. 0.10-'=, 0.28

6.23

Al 0.033';, 0.13, 0.34

;'r average of two base heats

The range of compositions examined was based primarily on harden- abllity requirements and anticipated response to heat treatment.

The steels were vacuum induction melted in an alumina lined furnace. Heats weighing approximately 40 pounds were cast into zircon ingot molds invested in steel shot to provide a rapidly solidified cast structure. The upper portion of the ingot molds was wrapped in a ceramic blanket and exothermic hot topping was applied after a solid skir had formed on ths ingot surface. The specific melting practice used is given in Table II and a sketch of the ingot casting setup is shown in Figure I. The ingots were hot worked by forging between flat dies using a 7 to 1 reduction. The forged product was approximately 2 1/2" wide x .6" thick.

The forging practice, outlined in Table III, involved appreciable redundant working to aid in breaking up the cast structure.

After forging all steels were cooled in air, normalized I hour at l650oF, and spheroidized for 12 hours at I3000F. Specimen blanks were cut from the forged plates, austenitized for 1/2-3/4 hours at 1550oF and quenched in agitated oiI.The tempering treatment consisted of a double temper of l+l hours at 800oF for all but the l.38%Si alloy which was double tempere^ for the same length of time at I0500F in order to avoid the shift in the 500oF embrittlement range which occurs in high Si steels. The choice of these tempering conditions was based on pre-

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TABLE I I

Steel Melting and Casting Procedure

Furnace - Vacuum Induction, 50^ Capacity, Alumina Lined Crucible

1. Charge electrolytic iron (A" dia, bars previously cast fro^n vacuum melted electrolytic Fe in flake form), Ni, Mo, Cr end .2I^C as graphite rod.

2 Evacuate chamber; at first sign of melting fill chamber with argon to 1/2 atm . pressure.

3. After meltdown, cool bath unt i 1 surface freezes then melt under low power to establish reference melting point tempera- ture with optical pyrometer.

*». Initiate C boil by reducing pressure gradually. Initally maintain temperature at 50oF (optical) above melting point; near end of boil superheat is approximately 100oF (optical). After pressure of 50>tt,was achieved, an additional 12-15 min. were required to reduce pressure to the 10-1 Strange.

5. After pressure of l0-15/<iis reached, add Fe-5i.

6. Backfill chamber with argon to 1/2 atm. pressure.

7. Add remainder of C, permit to dissolve, raise temperature to 125-150°F (optical) above melting point,

8. Add Fe-Mn and Al (single addition of Fe-Mn wrapped in aluminum foil).

'j. Pour at approximately 150oF (optical) superheat into zircon mold (see Figure 1). Entire mold setup was baked at 'tOO-600°F and inserted hot into furnace chamber immediately before Step 1.

10. 1 minute after solid skin forms on ingot surface, Turnace chamber is opened and exothermic hot topping added.

ZIRCON MOLD (MADE BY LOST WAX PROCESS)

16"

COARSE REFRACTORY GRIT

CROP LINE

FORGING STOCK

^^a^^ •* - - -r - i" -' - -' ^■^MM^j^—^i^ui

9"

I" THICK CERAMIC FIBER BLANKET

STEEL CAN

PACKED STEEL SHOT (APPROX. 1/16" DIAM.)

Figure 1. S^tup used for casting steel ingots.

TABLE I I I

Forging Practice

Heat Ingot at 2100oF for 2 hours; corge on 150 ton hydraulic press between preheated fiat dies, no lubri cat ion.

1. Forge ingot (appro .imately V' round , Figure 1) to 3" square; reheat (2100oF).

2. Forge 3" square to 2 1/2" square; reheat (2100oF).

3. Forge 2 1/2" square to 2" square; reheat (2000oF).

k. Forge 2" square to 1 1/2" thick x 2 plus" wide; reheat (2000CF).

5. Forge 2 l/2"(plus) width to 2 1/2"; reheat (2000oF),

6. Forge I l/2"(plus) thickness to 1", reheat (2000oF).

7. Forge 1" thickness to .650"; reheat {2000oF).

8. Forge 2 1/2" (plus) width to 2 1/2"; reheat (2000oF)

9. Repeat steps 7 and 8.

10. Cool bar in still air.

lininary heat treatment studies which indicated that the 800oF tempering temperature (and 1050oF for 1.38% Si) produced desired strength levels, small variations in hardness between steels and a reasonable assurance of avoiding both the 500oF and temper brittleness ranges.

Three types of specimens were machined from neat treated bars. The dimensions of tensile and notch bend fracture toughness specimens are shown in Figure 2. Standard Charpy V-notch impact specimens were taken from the broken halves of notch bend bars. The orientations of test specimens and directions of crack extension relative to the dimensions of the forged bars are illustrated in Figure 3.

Tensile tests were conducted at temperatures of 75 to -32I0F using a 60,000 pound capacity hydraulic testing machine at a crosshead speed of .Üi inches per minute. An air environment was used for 750F while the lower temperatures were achieved with ethyl alcohol and liquid nitrogen mixtures. In addition to the standard tensile properties such as tensile strength, .2% offset yield strength, reduction in area, and percent elongation the work hardening exponent, n, as used in the equation c =Ae (a=stress, A=constant, £=true plastic strain) was evaluated. A 1.0 inch gage length knife edge creep extensometer was adapted to provide chart records of load versus elongation during the entire tensile test. Only the portion of the stress-strain curve preceding maximum load (necking) was used to calculate n. The instantaneous stress was calculated using the instantaneous load and an area computed assuming a constant volume contribution from the plastic component of the elongation together with a slight correction because of an elastic Poisson contraction of the specimen diameter.

In most cases at least five values of stress and true plastic strain were obtained from the load-elongation curves. The work hardening co- efficient or slope of a log o-log e plot was obtained from a linear regression analysis of the data. The standard deviation in n, and linear correlation coefficient of X-Y (o-e) were generally such as to indicate excellent straight line fits, thus supporting the use of a a= AE strain hardening relationship to define n. All log o-log e data were also plotted münua'ly as a precautionary qualitative visual check of the calculated results.

Complete tensile data were not obtained for all specimens tested because occasionally at low temperatures premature brittle failures initiated under the extensometer knife edges. These failures always occurred after maximum load was reached so that acceptable yield strength, tensile strength, and n values were obtained. However, under these con- ditions, ductility data were incorrect and additional specimens were tested without the extensometer in order to generate valid results. These extra tests yielded duplicate tensile strength data but not yield strength or n values.

5/8 APPROX

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a) Tensile Specimen

i - 13 NC - 2A

THREAD

NOTCH ANGLE 60° INCLUDED; LESS THAN .005 R

W = ,920

THICKNESS, B = .480 SPAN, 2L = ^W

a0 = .365-.385

-. = .45-.55

b) Slow bend fracture toughness specimen (three point loadino)

Figure 2. Dimensions of tensile and fracture toughness specimens.

TENSILE

NOTCH BEND

CHARPY V-NOTCH

DIRECTION OF WOP.UNG

Figure 3. Locations of test specimens in relation to forged product.

Notch bend tests were performed using the specimeii illustrated in Figure 2b. The range of test temperatures and coolants were identical with those used in tensile testing. The machined notch used for generating a fatigue crack was of a straight through rather than chevron design currently recommended by ASTM (5,6). However, this notch geometry did not introduce any difficulties in obtaining reasonably straight and uniform crack fronts. Current ASTM recommenda- tions call for precracking at a rate sufficient to generate the last .050 inches of the fatigue crack in not less than 50,000 cycles (5). Evidence exists showing that at least in h;gh strength steels, this requ-rement may be too stringent (7,8). In the present study the rate of precracking was such that the last .050 inches of crack depth were produced in approximately 20,000 cycles. The ratio of minimum to nv;ximum load used in precracking was .25.

'he fracture toughness tests were conducted in three point bending. A coitinuous load displacement record was generated using a clip gage atteched to single edge razor blade knife edges spot welded to the specimen as shown in Figure V The shape of the load displacement curves could be classified into one of the three classes shown in Figt.-e 5. For each test two load values were obtained from the ioad- diS|. Iacement record. The maximum load, P , was used to compute a max'rrurn value of stress intensity factor, R , using equation I in Figure k. A load was also obtained for calculating the plain strain fracture toughness, K, , corresponding to the stress intensity for measurable crack growth under plain strain conditions. The procedure, illustrated in Figure 5, consisted of drawing a secant line OP,- with a slope 5% less than the straight line portion of the load-displacement record. The load P^, required to calculate K. , was the maximum load on the record preceaing or including P,.. Thus in the case of Class I curves Pn = P,. for Class 11 curves PQ^C, and for Class III curves P = P = P^. For Class ill curve? tRe first measurable crack extension corresponded to unstable crack growth. In the other cases some slow crack growth beyond that corresponding to K. preceded the onset cf rapid fracture and the peak load in the record occurred at a position past OP-. The load P was then used to calculate a tentative K. value, Kn using equation I.

The load-displacement records were examined further by drawing a horizontal line at .63 P« and measuring the deviation between 0A and the load-displacement record at that point if this deviation exceeded 1/5 "he value of the deviation of line OP from 0A, K was rejected as a vali'J K. number. Furthermore, in order for K^ to Be accepted as a

1 C VJ valid K. , the ASTM recorrmendation that the specrmen thickness B must equal or exceed 2.5 (Kn/

a ) . where 0 is the yield strength, was used. St 7 ^ f

12

TO RECORDER

Equation I): K P La 1/2

B W [5.8 - 9,2i + 43.6(2.)2 - 75.3(-r+ 77.M-)

W W W W

P = load in pounds L = 1/2 total span length a = crack depth (machined notch plus fatigue crack) W = specimen width B = specimen thickness

To obtai n: K use P ; max max

KQ use P^;

" K, evaluate K and load-deflection curves to i c vi

determine if criteria for valid K, are 1c

satisfied (see text for details).

Figure k. Schematic of test setup used in notch bend tests and equation used to calculate fracture toughness.

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The crack length in cases where K. was considered valid satisfied the criterion a^Z.Zdc/o ). This is slightly less conservative than the recommended ASTMypractive of a^ 2.5 (Kn/o ), however data for high strength steels have indicated that this variation produces a negligible systematic change in K. (9).

Standard Charpy V-notch impact tests were performed using a 220 foot pound capacity impact machine having a striking velocity of 17 feet per second. Tests were conducted at temperatures ranging from -321 to 250oF. The 250oF temperature was achieved using a molten both of a low melting Bi base alloy. Baths for lower temperatures con- sisted of either water, dry ice and ethyl alcohol mixtures, or liquid nitrogen. An air environment was employed for the 75°F tests. Both impact energy and percent fibrous fracture were determined in these tests. The fibrosity readings were visual estimates made using master charts drawn to show various amounts of crystalline fracture, for a range of shear lip sizes, at hypothetical percentages of total fibrous fracture. Larson and Nunes (10) showed good agreement between visual estimates and more laborious planimetric measurements of percent fibrous fr;.:ture.

15

I I I RESULTS AND DISCUSSION

A detailed tabular and graphical presentation of test data for individual steels appears in the Appendix. The tensile test results, consisting of the tensile strength, ,1% offset yield strength, reduction in area, elongation and work hardening exponent (n), are listed in Table Al (a) of the Appendix. Table Al(b) presents the results of limited tensile tests for base heat 5, tempered at 1000oF, comparing properties of longitudinal and transverse specimens. The high values of ductility in the transverse specimens are indicative of good steel cleanliness and fogging practice. The results of the notch bend tests appear in Table Ail. This latter table gives the pertinent information required to calculate the fracture toughness K and the plane strain fracture toughnass K, . The impact energy and fracture appearance data obtained in Charpy V-notch impact testing are presented in Table AIM,

The significant tensile and notch bend test data are summarized for each heat in Figures A 1 to A 18. These show the effect of test tempera- ture on the tensile and yield strengths, reduction in area, work harden- ing exponent (n), and the stress intensity factors K and K. . The stress intensity factors incorporate the influence or an intrinsic tempera- ture effect on toughness together with its effect on the transition from plane strain to plane stress fracture as temperature increases. At -3210F the yield strength was large and K sufficiently small such that valid K, data were obtained for all steels. With rising test temperature K, increased and the yield strength decreased so that a temperature was reached at which the specimen was too thin for obtaining valid K. data. The meaningful parameter above this temperature, which varied with compo- sition, was the maximum stress intensity factor, K . At still higher

max temperatures the fracture toughness continued to increase reflecting an increase in K , a further drop in yield strength, and an increasing contribution of ductile shearing in the fracture surfaces as revealed by growing shear lips and a shrinking central flat region. For some specimens at intermediate test temperatures both K and a lower valid K, value were obtained. The two data points for such specimens are v nnected by a vertical line in the graphs. The data for the two base cohiposition steals (heats 5 and 6) are plotted together in Figure Al with single lines th igh the points to indicate the properties of a steel whose composition is the average of the two.

The transition temperature in the notch bend tests was defined as the temperature at which K /a = .h. At this stress intensity level „ B, the nominal specimen thickness of .480 inches, was equal to 3 (K /o ) ■

max ys

16

A comparison of this.ratio with the requirement that for valid K, B ^ 2.5 (KA/o ) shows that this definition of transition

I c Q y s r • temperature is a reasonable measure of the fracture mode transition from K, to K .

Ic c

The results of Charpy impact tests are plotted separately for each steel in Figures A 19 to A 27. These figures show both the variation in impact: energy and percent fibrous fracture with tempera- ture. In high strength steels Charpy impact curves are generally flat and do not show the distinct transition in energy that is common for lower'strength body-centered cubic metals. The transition temperature can be defined in various ways depending on what energy, dimensional, or fracture appearance criterion is used. In the present study the lowest temperature at which the fracture surface was completely fibrous was employed to define the transition. Above this point, commonly called the propagation transition temperature, the energy is -ssentially constant and is referred to as the Charpy shelf energy C (max). In the present study C (max) was taken as the lowest energy for 100% fibrous fracture.

Effects of Alloy Additions

A comparison of the influence of alloying elements on toughness is provided in Figures 6 to 16. The changes in notch bend fracture toughness with single element variations are shown in Figures 6 to 13. Similarly, Figures 1^ to 16 contain Charpy impact energy curves for the individual alloy series. The transition temperatures for both the notch bend and Charpy tests are identified on these curves.

The toughness properties for each alloy series are discussed in the following paragraphs:

Carbon:

Carbon was detrimental both to notch bend (Figure 6) and Charpy impact toughness (Figure l^ta). The qualitative agreement between the notch bend and Charpy curves was generally good. At very low temoeratures the relative effects of carbon were not consistent between test methods, but but this may merely reflect scatter in the very limited amount of data in this region. The fracture mode transition and propagation tran- sition temperatures increasea wi th increasing carbon content. The oeak K values and C (max) were reduced; the change w > most pronounced in the range o¥ 0.28 - 0.35'-b C

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J I I I I L_

■200 100 0 100

TEST TEMPERATURE, F

Figure 7. Effect of mdnganese on notch bend fracture touqhness.

140

130

120

. 100

üi

90 Dt o t- <_> HO < u.

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60

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1

1, TRANSITION TEMPERATURE

j I

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TEST TEMPERATURE, F

Figure 8. Effect cf silicon on notch bend fracture toughness

20

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ys

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TEST TEMPERATURE, T

Figure 9. Effect of chromium on notch bend fracture toughness,

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2.83% Ni

1.26^ Ni

1, TRANSITION TEMPERATURE (K/(7 = .'♦) ys

x_l L—J I L- -200 !00 100

TEST TEMPERATURE, F

Figure 10. Effect of nicke! on notch bend fracture toughness

22

130

120

110 "

^ - 100

^

Cf 90 o h- <_) < an u.

>- 1- 70 l/) z UJ

z 60

en en Ul 50 a: h- LO

40

30

20

T 1 1 r

.. 13^ Mo

-300

1 , TRANSITION TEMPERATURE (K/O- = .M

J I L -200 -100 0 too

TEST TEMPERATURE, F

Figure 11. Effect of molybdenum on notch bend fracture toughness

23

z

I/! 5^

o o <

130

120

no

100

90

80

70

60

in 50 UJ Q£ t- lO

^0

30

20

n—i—r -i 1 r

J L

.28^ V

<.0\l V .101 V

, TRANSITION TEMPERATURE (K/<rys = .k)

J I L

■300 -200 -100 C

TEST TEMDERATURE, 0F

100

Figure 12. Effect of vanadium on notch bend fracture toughness

2^

130

120

110

100

90

30

70

60

£ 30

k l>0

o

<

I/) z

^0

30

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/ 1, TRANSITION TEMP.

(K/O-ys = ,k) H

[-

-J—

< .033';

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Al

i.l, 1 J

-300 -200 100 100

TEST TEMPERATURE

Figure 13. Effect of aluminum on notch bend r'-acture touahness

25

50

i*0

30

20

. (a)

3 lo

>- cc 0

< CL

30

20

10

1 r

1, TRANSITION TEMPERATURE MOOl FIBROUS)

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. (b)

, TRANSITION TEMPERATURE (100» FIBROUS)

J L

28^ C _

.35S C

M% c

J L

.36% Si

■300 ■200 •100 0 100 200

TEST TEMPERATURE, F

Figure 14. Variation of Charpy V-notch impact energy with carbon (a},silicon (b) content.

26

oo in

>- o en

< Q.

30

20 -

10

30

20

10

0

30

20

10

I, TRANSITION TEMPERATURE {]00% FIBROUS)

.825' Cr

1, TRANSITION TEMPERATURE (100% FIBROUS)

J L

(c)

1, TRANSITION TEMPERATURE (\00% FIBROUS)

6.23?: Ni

J L

■300 -200 -100 (

TEST TEMPERATURE

too 200

Figure 15. Variation of Charpy V-notch impact energy with L.iromium (a), manganese (b), and nickel (c) content

27

^o -

30 -

20

10

00

I 30

< ^ 20

0

20

1, TRANSITION TEMPERATURE (100% FIBROUS)

H 1 1 1 1 h (b)

, TRANSITION TEMPERATURE (1001 FIBROUS)

(c)

1 , TRANSITION TEMPERATURE (100% FIBROUS)

033% Al

_l !_

10% V

3*«% Al

■ ' I I

-300 ■200 00 0 100 200

TEST TEMPERATURE, F

Figure 16. Variation of Charpy V-notch impact energy with molybdenum (a), vanadium (b), and aluminum (c) content

28

urinoonese :

Manganese reduced toughness at all tenperatures in both te3ts as shown in Figures 7 and i5b. This behavior was also reflected in a pronounced rise in transition temperatures with increasing manganese content.

Si I icon:

Silicon exhibited lore conplex behavior tnan carbon or r.ianganese as shown in Tiyures 8 and l^tb. The transition termeralures were raised consistently with increasino Si content. However, the Maxi- mum K values and C (nax) were obtained at t''e highest Si level (1.3o/Si) followed by the lowest {0.06/Si) rather than the intermediate (0.36/?i) level. Both in the notch bend and Charpy impact test a cross-over occurred between the 0.36'O and 1.3B^ Si curves.

Chromi urn:

The-notch bend and Charpy impact curves comparing the effects of various chromium contents appear in Figures 9 and 15a respectively. The best properties were obtained for the lowest level (0,51* Cr). V.'ith increasing chromium the transition temperatures were raised aporeciably, but the effect on toughness was much smaller. In fact, both C (max) and peak fracture toughness were slightly higher ^or the 1.612 (> alloy than the base .82% Cr level.

Nickel:

Figures 10 and 15c show the notch bend and Charpy impact curves for four nickel levels in the range 1.26 - 6.23% Ni. The largest effect of nickel occurred between 1.26 and 2.83%. The 1.26% nickel steel exhibited the highest room temperature fracture touqhness and C (max). At the higher levels of ^.^5% and 6.23% C (max) was reduced slightly below the value for 2.83% Ni. In both tests the transition temperature was reduced con- sistently with increasing nickel content, but the effect was most pro- nounced in the range of 1.26-2.83%. Low temperature tounhness was also improved consistently with increasing nickel content.

Molybdenum

Highest Charpy and fracture toughness values were obtained with the lowest molybdenum level of 0.13% as shown in Figures 11 and 16a. Impact energy and fracture toughness were considerably lower at the intermediate (0.30%) level but a further increase in molybdenum content to 1.0^% produced only a relatively small decrease in both tounhness properties. Molybdenum raised the transition temperatures in bot!; tests. The fact that toughness was not imoroved by increasing the moly- bdenum content above 0.13% suggests that the 800oF tempering temperaturo employed in this study did not introduce appreciable temper brittleness.

29

Vanac i u"

Changes in toughness and transition temperature were not con- sistent with changes in vanadtun content as illustrated in Fjgu-es 12 and i6b. Lowest fracture toughness and Cnarpy inpact energy values and highest transition temperatures were obtained for the 0.10", V steel. The toughness was improved and the transition temperatures lowered when vanadium was either virtually absent (<0-0K'V) or increased to a level of 0.28.

Ai uni num

Aluninum was not beneficial to toughness above the aver je level of 0.033-' utilized in the base composition as shown in Figures 13 and 16c. The notch bend toughness and Charpy impact energy were reduced slightly with increasing aluminum content, but the difference in tough- ness between 0.13% and 0.3^'- a lumi num appears egligible. Similarly, the transition temperature was slightly lower for 0.033.' aluminum and relatively constant in the range of 0.13% to 0.3^':'.

Summa ry:

A comparison of the influence of alloying elements on toughness as measured in the notch bend and Charpy impact tests indicates considerable qualitative agreement between these test methods. The agreement .vas particularly good between the notch bend fracture mode transition temperature at K /- = .k and the Cnarpy impact test propanation transition temperature in the sense that relative changes due to alloying, for a given series, were essentially the same. A summary of the effects of alloying elements on both transition temperatures appears in Fiaure 17 and 18.

Relation of Toughness to Tensile Properties

A full interpretation of compositional effects on toughness requires an understanding ,*" how structural and material parameters are affected by alloying 9"'J the relation of these prope-ties to toughness. Currently the interplay of the various factors involved Is not understood sufficiently w2ll to permit this ideal coupling between composition, structure, and toughness properties in all but a few instances.

A more practical approach involves attempts to relate the tensile behavior of a material with toughness. This does not disregard the effects of alloying elements on structure but merely presumes that any structural changes will also affect common material parameters such as tensile properties which in turn can be correlated with toughness.

30

z: -a-

cc >

< ^.

o t- — in t- UJ — I—

^ o < :r I- a:

-700

0 1 2 3^ 5 <

ALLOYING ELEMENT, ATOMIC /

Figure 17. Effect of alloying elements on the notch bend test fracture mode transition temperature.

31

<_> —~ t- 2 O O 2 —

I Cd > UJ

(- >- — a. a: ce (_> < £ >- <_> t-

2 t/l — C

ce LU CO o: — r) u. i— < ■:<

cc o uj O Q. —

2 CO C LU

2 < < C

I 2 3^

ALLOYING ELEMENT, ATOMIC %

Figure 18. Effect of alloying elements on Charpy impact test propagation transition temperature.

32

Work Hardening Exponent:

An important factor affecting the stress-strain behavior of a metal and thus its energy absorbing characteristics at the tip of a crack is the work hardening exponent. Presumably any changes in flow characteristics which increase the ability of the plastic zore to absorb energy will result in improved toughness.

However the question of how toughness varies with n has been controversial. According to Gensamer (11) a large n is beneficial to toughness since it would increase the area under the stress-strain curve. The experimental work of Larson and Nunes (10) on blkO steel heat treated to a w de range of strength levels supported this reason- ing in that the Charpy V-notch impact toughness increased with the logarithm of n. The parameter correlated with n was the difference in impact energy between the lowest value required for 100z fibrous fracture. C (max), and that requ-red for 0^ fibrous fracture. However, other data are available which indicate that a low n can lead to improved fracture toughness in high strength steels (12,13)- More recently, the question of whether a high or low value of strain harden- ing is beneficial to toughness has been related to the macroscopic fracture mode. Steigerwald and Hanna (IM analyzed the effect of n on toughness for various specimen thicknesses in high s.ength steels, aluminum, and titanium al'oys. For thin specimens the fracture occurred entirely by plane stress mode and the surfaces were 100% slant. In this region a high work hardening exponent was beneficial to toughness. As thickness increased the fracture surfaces contained regions of both slant (shear lips) and flat fracture. In this region a low n improved toughness. A decrease in n also shifted the transition from slant to flat fracture to larger thicknesses. This behavior was placed on an analytical basis by plotting the normalized fracture toughness parameters (K /J ) 1/2-B versus (K. /: )2|/2-B on logarithmic coordinates. Figure

C YS 1C V5 19 illustrates schematically'th!s method of presentation and shows the effect of n on the curves.

In the present study the fracture mode transition was induced by varying temperature rather than thickness and changes in the normalized toughness parameters resulted from the effects of temperature on 3 , K, , and K_. The stress intensity factor versus temperature transition^curves are analogous to the size induced transition curves described above and shown in Figure 19. The definition of a temperature induced fracture mode transition at K /a = -k provides the required deqree of normai- ization for valid comparisons between steels in terms of n.

The variation of n with alloying is shown in Figure 20. Increasing nickel and manganese content increased n appreciably. Carbon and moly- bdenum generally raised n, chromium mcieased n at al1 temperatures, while the effect of aluminum was negligible above approximately -100oF and

33

lOO'z SLANT FRACTURL

Morh

o MIXED MODE FRACTURE

FLAT FRACTURE

n

nI< n2

n n ARE WORK HARDENING EXPONENTS

Log (KIc/crys) 2 1

27rB

Figure 19. Schematic illustration of method for presenting toughness parameters (ref. ]k).

cture

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35

slight ar lower temperatures. For the vanadium ser,es n was highest for the 0.10: V alloy followed by >0 01; and 0 28 V steels. A constant n of .067 was used for the 1.38- Si alloy This value corresponded 10 the slope of the 'ogr - log-: curve at higher strains near the necking pomt. The data for the silicon series also indicated a rise in n with alloy content providing the work hardening exponent was measured at approximately the necking strain. A manual plot of log; - log-: for all steels showed that n was essentially strain independent for all but the highest silicon alloy, which exhibited a steadily increas'ng rate of work hardening with strain. The strain at maximum load for this alloy was higher at all temperatures than implied by t.e n obtained from the lineör regression analysis. This suggests that although the relation z = A-; is not applicable for this composition the slope of the logr -log-: curve in the vicinity of the necking strain is a more meaningful measure of n for comparison with other stee Is .

The temperatu'e induced fracture mode transition defined by K /r = .^4 is correlated with n in Figure 21 for the various alloy series. In general the transition temperature was reduced with decreasing n, consistent w:th the shift in fracture mode transition to larger thicknesses described earlier. The apparently inconsistent effect of vanadium on the notch bend transition curves Figure 12, acquires a rationale in terms of n in this representation. The effect of aluminum is not well defined and may be obscured by the small range of both n and transition temperature.

The effe't of nickel was opposite to that of the other elements. Increasing mckel raised n markedly yet lowered the transition tempera- ture. The dif^rent influence of nickel on the transition temperature is probably an indication that n is not a unique parameter which affects the temperature induced fracture mode transition. Instead the combined variation of alloying and test temperature may influence both micro- structural and micro-fracture characteristics in a so-far undetermined manner which overshadows the effect of n.

The influence of n on C (max) and 75CF fracture toughness is illustrated in Figures 22 and 23 while Figure 2k shows a plot of the Charpy V-notch energy difference as described earlier versus the logarithm of n. The effect of n was not clear in any of these representations. The data he in a broad band suggesting a drop rather than rise in toughness with work hardening exponent.

These results are in contrast with those of Larson and Nunes (10) on the effect of n in Charpy V-notch impact tests of ^3^0 steel. These authors varied n by changing the strength level through heat treatment whereas in the present study n was varied through composition changes

36

.06 j

.05

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* .28^ C

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.Ok

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0Jb.l3/Al

r .0330/ Al -\

•200 -100 0

TEMPERATURE, 0F

■200 -100 0

TEMPERATURE, 0F

Figure 21. Variation of notch bend transition temperature (at K /a =.k) with work hardeninq exponent, n max ys a K .

at the transition temperature for various steels

37

50

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LU CO

X u.

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- ABOVE 750F n WAS TAKEN AT THE 75 F VALUE.

1 1 1 1 1 1 1 1 1 1 1 1 5 6 7 8 9 10

n x 100

igure 22. Effect of work hardening exponent on the Charpy shelf energy, C (max).

38

200

150 -

100 y-

90

X ID E

80

70

60

L T 1 ' I ' 1 I T " '

U J y A y A 1— —\ L J l-

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n x 100

Figure 23- Effect of work hardening exponent on notch bend fracture toughness at 750F.

39

LTV

.10

.09

.08

.07

.06

.05

04 -

.03

.02 -

.01 0 10 20 30 ko 50

CHARPY V-NOTCH IMPACT TEST ENERGY DIFFERENCE, (Cv(nax) - ENERGY AT

0^ FIBROUS FRACTURE), FT. LBS.

Figure 2k. Effect of work hardening exponent on Charpy V-notch impact energy.

ko

and the variations m strength level were slight. The fact that K at 750F did not increase with n probably occurred because the fractures were still mixed mode and an upper shelf of purely slant fracture was not observed. Thus the expected effect of n in this region (14) would be to decrease toughness.

The ambiguous effect of n in affecting Charpy impact and notch bend toughness suggests that n, a measure of the tensile instability strain, may not be the governing fracture parameter. Instead, in this region a critical strain criterion based on fracture ductility rather than necking strain may be operating. Figures 25 and 26 show the effect of tensile ductility in notch bend tests, at 75CF and -3210F, and Charpy impact tests at the upper shelf energy and -3210F. The "upper shelf" toughness values exhibited good correlation with true fracture strain. At -3210F, where a critical stress relation- ship would probably apply, the toughness was fracture strain independent. Similar correlations of tensile ductility with C (max) have been observed for other steels (15) including the ^3^0 steel investigated by Larson and Nunes (10).

Quant i tative Models:

A thorough understanding of how material parameters influence toughness requires a separation of tneir effects. This, however, is generally difficult to achieve experimentally. An alternate approach is offered by certain models which relate fracture toughness to mechanical and structural parameters. Krafft has proposed a tensile instability fracture model which regards crack propagation as occurring by the successive rupture of small tensile specimen-like regions of size d (16). "!■'■'; initiation of unstable rupture of these ligaments is considered t.0 occur when the uniaxial tens i le necking strain (equal to n) is reached. The equation K = En ^2-dywhere E is Young's modulus, has bet.-; correlated with structure for .kSZ C-Ni-Cr-Mo steels containing different sulfur levels (17).

Hahn and Rosenfield (18) have developed a ductile fracture model which avoids the difficulty of having to.,measuce the process zone size dj. Instead the equation K. = (2/3 EYE n ) , where E = Young's modulus, Y = yield strength, c = true fracture strain and n = work hardening exponent, provides a direct relation between ordinary tensile parameters and K, . K, values calculated from tensile properties of a few steels,

1 c 1 c titanium and aluminum alloys indicated reasonably good agreement with experimentally measured values, considering that the tensile and K. data employed in the comparison were obtained from different sources.

^

60

20

"?■ •321JP

J L ■L, -L ± ■2 .k .6 .8

rDUE FRACTURE STRAIN. £ f •/n ■"•n/A

Figure 25. Effect of tens! 1 e ductility on notch bend fracture toughness of two *est tenoeracures

hi

w

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20 H

o c UPPER SHELF

(ABOVE 750F £f WAS TAKEN

AT THE 750F VALUE)

/

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"LOWER SHELF"

(-3210F]

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TRUE FRACTURE STRAIN. £f,in Ao/A

1,0

Figure 26. Effect of tensile ductility on the Charpy upper shelf energy and -32T:F energy.

^3

- *_':',e>' examination of this node! c^n be oerfomed using the K. a-d tensiie data generated on the steeis used in the cuiren: study. K .-.as computed fror tensile test results obtainea over the temperature ra^ge o* -321 to 75"F. ¥oungls -noduius is essentially independent of compos-, on for low aMoy steels; based on data for ^3^0 steei (19) it was assu~eci to varv ''nearly fro^ a 75"F value of 29 x 10^ ks i to 31 x iO3 kii at -32!:F.

A va:id comparison of calculated and measured K was possible only at low temperatures. In general the variation of calculated K with composifon within a particular alloying element series differed from the results of the actual notch bend tests Furthermore, with decreasing test temperature the ca'culated K values changed slightly in comparison with the experimental ^<i*a which exhibited a sharp drop. At -32! "F the pre- dicted values were approximately double the observed results. The curves in Figure ?7 ?re tyjical of those obtained in attempting to apply the Hahn-Rosenf ie i a mo'-el.

The large discrepancy at -321'F is believed to be the result of an increase in yield strength and change in microscopic fracture mode at low temperatures. In a fractograpnic study of ^3^0 steel Charpy specimens Bucher et a I (2Cn observed an increasing amount of quasi- cleavage fracture with increasing strength level both at room tempera- ture (by varying tempering temperature) and at a single tempering temperature (by decreasing the test temperature). Hahn and Rosenfield have compared the predicted and calculated room temperature K. values for ^3^0 steel as a function of tensile strength (18). Figure^S shows that the discrepancy tends to increase appreciably at strength levels above 2C0 ksi ultimate strength.

These observations add support to the earlier sugaestion that a critical strain criterion may not apply to fracture in low alloy steels at low temperatures Any use of the Hahn-RosenfieId model to predict K. should involve an appreciation for the microscopic fracture mode. If a portion of the fracture involves quasicleavage or intergranular crack propagation then the observed toughness will be less than predicted from the model.

^

130

120

\z 110 S.

100

^ 90 *—' cc c 1- 80 (_> < Ü.

>- 70 H-

vy^

Z 6n LU !- Z

so l^ in LU

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30

20

.13? Mo

1 . TRANSITION TEMPERATURE (K/CT * .k)

EXPERIMENTAL CURVES

J L J L J. I I I

-300 -200 100 100

TEST TEMPERATURE, F

Figure 27. Conparison of measured and calculated K (typical resul ts) . Note: Above the transition temperature experimenta'

values are generally not valid K, . J ' 1c

^5

ZOO-

ISO- \4 e KK)

^*A1rXB Various Investigators (ref. 18)

—»Colcukited curve

so #i^ 1 i

0 100 200 300

Ultimate Tensile Strengrtl, ksi

Figure 28. Calculated and measured K, values for 43^0 steel, (ref. 18) lc

k6

IV SUSMARY A;.J CONCLUSIONS

This study has examined the effects of various elements on the notch bend fracture toughness and Charpy impact behavior of a .35-'C, 3^Ni, Cr-Mo-V martensitic steel at a room temperature yield strength of approximately I6O-I8O ksi- A classical approach was used in the design of alloys which permitted an evaluation of single element effects rather than interactions.

An analysis of the toughness data was conducted in terms of the effects of alloying elements on yield strength, tensile ductility, and work hardening exponent. The elements C, Mn, Si, Cr, ard Mo raised both the notch bend fracture mode transition temperature and tne tharpy propagation transition temperature. These elements generally raised the work hardening exponent and reduced toughness. In amounts above that required for deoxidation and grain refinement, aluminum appeared to degrade the transition temperature and toughness slightly. Increasing the vanadium content from 0.1 to 0.28^ produced a significant improvement in toughness and transition temperature.

Relative changes in transition temperature in the notch bend test corresponded with shifts in the Charpy impact test propagation transition temperature.

Over the range of compositions examined (1.26 - 6.23^) mckel decreased the transition temperature and improved toughness at the lower test temperatures. For most alloy series the notch bend transition temperature was lowered as n decreased, but nickel acted in an opoosite "^nner,

Charpy shelf energy, C (max), and K at 75°^ did not correlate well with n. Good agreement was obtained however when these parameters were plotted versus true fracture strain. At -321°F toughness was essentially fracture strain independent suggesting that a critical strain criterion based on fracture strain is valid only when fracture is ori nci pal ly fi brous .

',. conparison of K calculated from tensile data using the Hafin- ;osen'icld node! with neasured values suggested that the increased contribution of non-ductile fracture which accompanies increases in strengfi and/or decreases in test temperature in low alloy steels can load to large errors in the predicted toughness.

hi

V LIST OF REFERENCES

1. Mi'itary Specification, "Tubular Parts for Cannons over jOmm, Steel Forgings for," MlL-T-10^580(MR), (1 April 1966).

2. Military Specification, "Steel Forgings, Tubular Parts for Cannon, High Yield Strength (160-180 Ksi)," Ml L-S-^HSCMR) , (23 March 1967).

3. J. F. Throop, D. P, Kendall, J. H. Underwood, and R. R. Fujczak, "Fatigue Crack Propagation in Single-Notch Tensile Specimens of Steel," Prepared for presentation to ASTM Committee E-9, Watervliet Arsenal, (March, 1967).

A. E. A. Steigerwald and C. Vishnevsky, "Literature Survey on the Influence of Alloy Elements on the Fracture Toughness of High Strength Steels," TRW Inc., Contract OAAG ^6-67^-0171, AMMRC CR-67-13(F), (February, 1968).

5. ASTM £-2^, "Recommended Practice for Plane-Strain Fracture Toughness Testing of High Strength Metallic Materials Using a Crack-Notch Bend Specimen," (1967).

6. W, F. Brown and J. E. Srawley, Plane Strain Crack Toughness Testing of High Strength Metallic Materials, ASTM STP. No. '♦lO, (December, 19ö6),

7. E. F. Walker and M. J. May, "Single Edge Notch Specimens of Two High Strength Martensitic Steels Tested in Tension and Bending," BISRA Report MG/EB/337/67, (1967).

8. E, F. Walker and M. J. May, "Note on the Effect of Fatigue Pre- Cracking Stress on the Plane Strain Fracture Toughness of Several Martensitic Steels," presented to ASTM E-21* Subcommittee 1, (MarcK 13, 1968).

9. M. F. Amateau and E. A. Steigerwald, "Test Methods for Determining Fracture Toughness of Metallic Materials," Final Report, Contract AF 33(6!5)-3827, (May: 1967).

10. F. R. Larson and J. Nunes, "Relationships Between Energy, Fibrosity, and Temperature in Charpy Impact Tests on AISI 43^*0 Steel," Proc. ASTM, 62, 1192 (1962).

11. M. Gensamer, "Strength and Ductility - Campbell Memorial Lecture," Trans. ASM, 36_, (19^6).

48

12. W. W. Gerberich, ''Fracture Toughness of SAE 43^*0 Steel Heat- Treated to 200,000 to 300,000 psi Yield Strengths," TR32-96, Jet Propulsion Laboratory, Pasadena, California, (1961).

13. W. W. Gerberich, "Plane Strains and Energy Density in Cracked Plates," Experimental Mechanics, ^_, 335, (196M.

\h. E. A. Steigerwald and G. L. Hanna, "Influence of Work Hardening Exponent on the Fracture Toughness of High-Strength Materials," Trans. AIME, 2k2, 320, (February, 1968).

15. A. S. Tetelman and A. J. McEvily, Jr., Fracture of Structural Materials, John Wiley, New York, (1967). ~~ ~ ———

16. J. M. Krafft, "Correlation of Plane Strain C-ack Toughness with Strain Hardening Characteristics of a Low, a Medium, and a High Strength Steel," Applied Materials Research, 88, (April, 196M.

17. A. J. Birkle, R. P. Wei and G. E. Fellissier, "Analysis of Plane- Strain Fracture in a Series of O.'iSC - Ni-Cr-Mo steels with Different Sulfur Contents," ASM Trans. Quart., 59, 98l, (1966).

18. G. T. Hahn and A. R. Rosenfieid, "Source of Fracture Toughness: The Relation Between K and the ordinary Tensile Properties of Metals," Applications Related Phenomena in Titanium and its Alloys, ASTM STP. No. ^32, 5, (196«).

19. Mil. Handbook 5A, Metallic Materials and Elements for Aerospace Vehicle Structures, Table 2.3.1.1(3), Figure 2.3.1.2.^4, (February 8, 1966).

20. J. H. Bücher, J, T. Cammett, J. C. Jasper, G. W. Powell and J. W, Spertnak, "The Relationship of Microstructure to Strength and Toughness in High-Strength Steel," AFML-TR-65-60, (March, 1965)

^9

APPENDIX

50

TABLE Al (a)

Summary of Smooth T( ;ns ile Properties

Specimen (Heat No. - Tesi Tens!le .Ü Yield /Elong. % Red. Work Hard. Test No. ) Temp.0 F

75

Strength(ksi)

199.9

Strength{ksi)

177.9

1"G.L. in Area

49.7

Exponent(n)

5-1 14 .

5-2 t ä 200.9 176.4 14 45.2 .053 5-8 20 206.4 181.9 15 49.1 .056 5-7 -ho 207.6 179.6 15 47.3 .059 5-5 -too 216.9 188.6 13 ^7.3 .058 5-6 -215 229.4 - - 37.3 - 5-6D 11 226.2 201.7 12 36.6 .050 5-3 -321 256.3 230.1 - - .051 5-A I 1 251 .2 - - 18.1 5-10 ■ I 253.6 229.0 - - .056 5-11 r i 25^.5 - - 33.6 - 5-12 t t 254.0 - - 34.3 -

6-1 75 197.8 182.5 14 50.7 _ 6-2 t f 198.0 176.4 15 56.3 .052 6-8 20 200.8 178.6 15 50.7 .049 6-7 -kQ 207.5 183.9 16 51.3 .051 6-5 -100 213.4 190.9 15 48.4 .047 6-9 -140 2)4.9 187.6 15 45.5 .057 6-6 -215 223.9 - - - - 6-6D 1 i 222.2 201.5 - - .050 6-3 -321 256.3 230.1 - - .042 6-^ 1 1 251.0 - - 42.1 - 6-11 1 ! 251.0 - - ^3.9 - 6-12 t 1 252.0 - - 32.7 -

7-1 75 195.8 169.6 15 50.3 .055 7-2 11 191.6 165.8 17 48.7 .057 7-8 20 200.2 171.1 15 48.5 .063 7-7 -'40 204.5 - 16 44,9 - 7-5 -100 210,4 180.2 14 46.7 .062 7-9 -150 215.5 182.9 16 45.5 .063 7-6 -215 220.5 196.7 - .063 7-6D 11 223.9 195.3 14 41.3 -057 7-3 -321 248.1 210.5 14 26.4 .068 1-k " 251.0 - 15 37.0 -

51

TABLE Al (a) (Continued)

Specimen Heat No.- Test Tensi le .1% Yield ^Elong. / Red. Work Hard. Test No. Temp.°F

75

Strength(ksi)

200.2

Strength(ksi)

182.2

T'G.L. in Area Exponent (n)

8-! 15 54.6 .040 8-2 ii 200.7 183.7 15 57.4 .041 8-8 20 205.4 187.4 15 53.7 .044 8-7 -40 207.6 189.5 14 53.6 .045 8-5 -100 214.9 199.1 12 57.2 .043 8-9 -150 218.0 200.8 15 54.1 .048 8-6 -215 230.6 213.8 - 43.6 .047 8-3 -321 253.4 240.6 15 35.9 .029 8-4 1 t 250.6 241.8 16 42.0 .035 8-10 i; 249.3 235.0 15 31.2 .036

9-1 75 189.3 171 .4 14 59.6 .044 9-2 n 187.2 168.5 16 58.6 .045 9-8 20 195.^ 176.4 15 54.2 .046 9-7 -40 198.2 177.4 16 53.7 .050 9-5 -100 204.7 184.3 16 51.5 .046 9-9 -150 208.5 186.0 15 50.3 .053 9-6 -215 214.7 195.3 ~ 49.2 .045 9-3 -321 242.4 224.2 16 42.2 .040 3-k • 1 241.8 222.4 18 40.9 .040 9-10 1 1 242.4 - 16 44.0 -

12-1 75 181.6 165.3 15 59.5 .040 12-2 ) I 182.6 165.3 16 56.9 .042 12-8 20 189.7 171.5 16 59.3 .046 12-7 -40 192.4 173.8 17 58.9 .046 12-5 -100 197.2 178.7 16 56.1 .043 12-9 -150 206.3 - 16 55.1 - 12-6 -215 211.0 192.8 - 53.! .045 12-3 -321 234.7 221 .8 - - .043 12-4 11 237.9 - - 46.7 - 12-10 11 234.8 218.5 17 40.3 .039

52

TABLE Al (a) (Continued)

Specimen

(Heat No.- Test Tensile .2k Yield ^Elong. * Red. Work Hard. Test No. ' Temp.0F

75

Strength (ksi)

209.5

Strength (ksi)

183.0

i"G.L. in Area Exponent(n)

13-1 15 49.9 .051 13-2 11 207.6 182.0 14 51.6 ,054 13-8 20 214.2 185.3 15 49.8 .059 13-7 -40 217.6 190.9 14 48.8 .054 13-5 -100 222.7 193.6 14 47.7 .056 13-9 -150 225.7 199.6 16 49.8 .052 13-6 -215 234.3 206.2 - 45.2 .054 13-3 -321 259.1 234.9 15 28.8 .046 13-^

11 261 .0 - - -

13-10 11 260.4 235.9 - - .047

]k-\ 75 193.7 175.8 17 59.9 .041 14-2 i l 193.0 175.0 16 59.4 .041 14-8 20 197.1 179.9 16 59.8 .04! 14-7 -40 202.4 185.4 17 57.8 .042 14-5 -100 205.6 187.3 18 57.7 .045 14-9 -150 211.7 192.0 15 54.1 .046 14-6 -215 219.2 194.0 - - .036 14-6D 1 1 217.2 189.5 - - .045 14-3 -321 244.2 228.5 18 46.2 .038 14-4 1 1 243.8 227.2 16 44.3 .039

16-1 75 188.5 171.5 20 59.2 .047-'- 16-2 1 i 187.1 168.7 20 57.8 .056'- 16-8 20 194.3 176.6 18 54.9 .044-..- 16-7 -40 189.6 180.7 20 57.2 , 048 v 16-5 -100 204.7 185.3 19 50.9 .045;

16-50 11

204.5 185-0 16 52.7 .050" 16-9 -150 210.4 190.8 20 54.4 .048- 16-6 -215 216.7 193.2 - 49.2 .050- 16-3 -321 242.2 217.5 - - .053" 16-4 11

243.5 - 19 42.4

Values based on lease squares fit of logo - loge data, however, n was considered to be approximately .067 at all temperatures (see text for detailed explanation).

53

TABLE Al(a) (Continued) .

Specimen (Heat No.- Test Tensile .2^ Y.-ld %E1ong. % Red. Work Hard. Test No. Temp.0F Strength(ksi) Strength (ksi) r'G.L. in Area Exponent (n)

17-1 75 204.7 170.0 14 44.3 .072 17-2 ii 204.3 168.2 15 42.2 .077 17-8 20 210.7 175.0 IS 45.1 .080 17-7 -ko 214.9 174.7 15 44.3 .081

17-5 -100 219-6 177.1 14 42.6 .080

17-9 -150 226.8 175.6 15 39.4 .oyi 17-6 -215 227.8 186.4 - 40.7 .079 17-3 -321 257-7 216.8 - - .071 I7-* H 258.7 - 16 17.2 - 17-10 n 259.2 216.5 15 23.5 .071

l8-i 75 199.6 1/..0 16 56.6 .055 18-2 II 200.5 175.2 15 51.4 .055 18-8 20 205.2 178.1 15 52.3 .058 18-7 -1.0 211.1 181.4 15 49.5 .063 18-5 -100 214.0 183.6 14 48.3 .06! 18-9 -150 217.9 188.5 14 48.3 .061 18-6 -215 228.6 197.6 - - .061

»

18-6D " 229.0 200.2 14 45.4 .0S8

18-3 -321 253.4 221.5 14 29.0 .057 18-4 ■ 1 255.9 223.5 - - .065

19-1 75 200.8 177.4 14 51.3 .050 19-2 II 200.9 176.6 14 49.8 .053 19-8 20 203.8 179.8 14 47.3 .053 19-7 -i.0 211.6 184.9 15 46.7 .059 19-5 -100 216.7 188.9 13 45.4 .056

19-9 -150 221.7 190.7 14 43.7 .060 19-6 -215 228.4 199.5 - 43.1 .057 19-3 -321 25318 224.1 - - .058

19-2» n 256.9 - 15 31.7 -

19-10 n 255.3 225.6 14 31.7 .053

20-1 75 202.0 178.9 13 50.2 .051 20-2 n 203.6 180.5 13 48.5 .051 20-8 20 209.3 183.2 14 47.8 .057 20-7 -40 213.2 187.5 13 41.8 .056

20-5 -100 219.9 192.4 14 44.2 .056

20-9 -150 223.2 198.0 14 43.6 .051 20-6 -215 228.8 204.3 - 38.3 .050

20-3 -321 255.3 229.7 14 28.5 .042 20-k II 256.1 228.4 - - .051 20-10

11 258.5 231.9 - - .050

54

TABU Al (a) (Continued)

Specimen '.Heat No.- Test Tens!1e .2% Yield %Elong. % Red. Work Hard. Test No. } Temp.0F Strenc|th(ksi) Strength (ksi) 1"G.L. in Area Exponent (n)

21-1 75 193.7 173.1 15 55.2 .046 21-2 ii 194.6 173.7 14 53.6 .046 21-8 20 196.3 174.1 13 52.4 .049 21-7 -ko 202.3 180.2 15 50.2 .051 21-5 -100 207.2 184.4 15 50.2 .052 21-9 -150 211.6 188.6 14 46.7 .051 21-6 -215 220.7 199.6 - - .048 21-3 -321 245.9 225-5 16 36.6 .046 2]-k It 245.^ 224.9 18 40.1 .046 21-10 II 247.7 226.2 16 r8.2 .039

22-1 75 197.4 182.3 13 52.4 .036 22-2 II 198.6 182.6 15 53.0 .040 22-8 20 203.1 187.5 14 52.3 .042 22-7 -40 208.4 193.4 15 53.4 .038 22-5 -100 210.2 193.0 15 50.7 .045 22-9 -150 219.1 202.5 14 47.8 .040 22-6 -215 224.7 207.4 - - .040 22-60 II 225.7 208.9 - - .043 22-3 -321 251.2 236.9 16 35.8 .036 22-k II 252.4 237.0 18 35.5 .038

26-1 75 192.0 174.5 16 61.5 .041 26-2 M 191.5 172.8 15 58.0 .044 26-8 20 196.8 177.9 16 60.4 .045 26-7 -40 199.7 179-8 IS 57.7 .047 26-5 -100 20^.2 185.1 15 55.1 .044 26-9 -150 208.8 189.8 16 54.8 .043 26-6 -215 216.9 198.9 - 54.2 .042 26-3 -321 243.1 224.2 17 43.5 .042 26-'4 II 2i»6.0 228.9 16 45.3 .039 26-10 n 244. i» 226.2 17 40.! .040

27-1 75 186.4 169.2 15 54.0 .044 27-2 H 185.9 16:. 9 17 61,4 .045 27-8 20 189.9 169,5 17 59.5 .049 27-7 -40 196.2 174,5 17 59.5 .049 27-5 -100 199.5 180.3 15 56.2 .044 27-9 -150 208.1 - 18 55.1 - 27-6 -215 216.i 198.0 - - .047 27-60 11 209.4 188.2 15 56.6 .040 27-3 -321 237.2 217.7 18 41.3 .041 27-k II 236.6 215.5 16 45.0 .037

55

TABLE Al (a) (Continued)

Specimen (Heat No. - Test No.)

28-1 28-2 28-8 28-7 28-5 28-9 28-6 28-3 28-i» 28-10

29-1 29-2 29-8 29-7 29-5 29-9 29-6

29-3 23-k 29-10

^est Tensile 22: Y^OI/H q-n

75 It

20 -A0

-100 -150 -215 -321

75 ii

20 -ko

-100 -150 -215 -321

208.7 209.9 211.5 217.6 222.5 226.6 23'.1 259.6 262. A 261 .i.

198.A

197.3 202.8 208.8 217. 220. 225.1 251.8 251.4 252.0

I • 9

187 188 186 188 196 198 204.9 235.4 233.8

231.7

162.1 161.3 165.2 i64.9 171.0 171.3 182.9 206.3 205.2 209.8

12 14

15 16

13 14

14

15

15 15 15 15 14

13

16

14

% Red. in Area

48.0 48.0 42.0 44 44

43 44 24 35.0

48.6 49.2 49.2 44.3 45.1 42.1

47.3 32.0

23.7

Work Hard. Exponent(n)

.048

.047

.054

.060

.052

.055 ■ 053 .046 .050

.053

.076

.089

.079

.089

.088

.092

.078

.06«»

.077

.074

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TABLE All!

Summary of Charpy V -Notch Impact Properties

Specimen Test Impact Fibrous Heat No.-Test No. Temp.0F Energy(ft.lbs .) Fracture {%)

5-8 170 23 \/k 100 5-6 \ko 23 Wk 100 5-7 110 21 95 5-9 92 21 80 5-1 75 20 50 5-10 50 17 1/2 ko 5-5 20 16 \/k 30 5-2 -ko ]k 10 5-3 -100 11 l/k 15 5-1* -321 5 1/2 0

6-6 140 25 100 6-7 110 2k ]/k 100 6-10 92 23 1A 100 6-1 75 22 ]/k 95 6-9 63 21 yk 80 6-8 50 21 1/2 65 6-11 35 21 kS 6-5 20 18 1/2 35 6-2 -ko ]k 1/2 15 6-3 -100 15 5 6-^ -321 6 0

7-6 ]kQ 22 100 sc* 7-7 no 23 100 sc^ 7-1 75 22 1/2 95 7-9 63 21 75 7-8 50 19 1/2 60 7-10 35 19 \/k kS 7-5 20 17 1/2 35 7-2 -ko 13 1/2 25 7-3 -100 12 15 7-h -321 5 ]/k 0

8-7 170 30 100 8-10 155 29 1/2 100 8-6 \ko 30 yk 90 8-9 125 31 80 8-8 110 26 65 8-1 75 26 1/2 50 8-5 20 16 1/2 30 8-2 -ko 15 1/2 10 8-3 -100 13 1/2 5 B-k -321 k ]/k 0

scattered crystalline

68

TABLE AIM (Continued)

Specimen Test Impact Fibrous Heat No.-Test No. Temp. 0F Energy (ft.lbs.) Fracture {%)

9-6 140 32 100 9-1 75 30 3/4 100 9-9 63 30 i/? 100 9-7 50 28 95 9-10 35 26 1/4 90 9-5 20 25 3/4 80 9-11 5 25 65 9-8 -10 23 1/2 50 9-2 -40 19 1/4 35 9-3 -100 15 3/4 15 9-^ -321 8 0

12-6 140 40 1/2 100 12-11 110 42 100 12-1 75 40 1/4 100 12-8 63 39 3/4 95 12-7 50 37 85 12-10 35 33 75 12-5 20 31 3/4 70 12-12 5 29 1/4 55 12-9 -10 26 1/4 45 12-2 -40 21 30 12-3 -100 16 1/4 5 12-^4 -321 5 0

13-7 170 23 100 13-6 140 22 1/2 100 13-10 125 21 1/2 95 13-12 no 21 1/2 80 13-8 no 21 1/2 80 13-9 92 21 1/4 65 13-1 75 21 1/2 55 13-11 50 18 35 13-5 20 16 3/4 25 13-2 -40 17 10 13-3 -100 12 5 13-^ -32! 4 3A 0

14-6 140 35 1/2 100 14-1 75 38 1/4 100 1^-7 50 33 1/2 100 H-10 35 31 1/4 100 1^-5 20 3' 3/4 95 ]k-S 5 30 1/2 90 14-8 -10 28 1/4 75 1*4-11 -25 26 1/2 65 14-2 -40 25 55 1^-3 -100 18 15 14-4 -321 5 1/4 2

69

TABLE AIM (Continued)

Specimen Test Impact Fibrous Heat No.-Test No. Temp.0F Energy (ft.lbs.) Fracture (%)

16-7 170 kl Ilk 100 16-6 ]k0 hi Mk 95 16-9 125 39 1A 85 16-8 110 36 Mk 70 16-1 75 31 Ilk 50 16-10 50 26 35 16-5 20 20 1/2 20 16-2 -ho 16 5 16-3 -100 10 Ilk 2 16-i» -321 3 0

17-9 250 21 Mk 100 17-8 212 18 1/2 90 17-7 170 18 Ilk 60 17-1 ]k0 18 Mk kS 17-6 1^0 18 kS 17-10 75 13 Ilk 20 17-11 50 \k 1/2 20 17-5 20 13 15 17-2 -ko 12 5 -.7-3 -100 10 2 ]7-k -321 3 Mk 0

18-7 170 25 1/2 100 18-6 ]k0 25 Mk 100 18-9 '25 Ik Mk 95 18-8 110 Ik Mk 70 18-1 75 23 50 18-10 50 18 Ilk 30 18-5 20 15 Ilk 20 18-2 -i.0 15 10 18-3 -100 13 1/2 5 ]S-k -321 k 0

19-7 170 21 Mk 100 19-6 140 22 Ilk 100 19-8 110 19 1/2 90 19-9 92 19 80 19-1 75 19 55 19-10 50 18 Mk k5 19-5 20 \k Mk 20 19-2 -ko 12 Mk 10 19-3 -100 11 1/2 5 \3-^ -321 5 2

70

TABLE All! (Continuec H Specimen Test Impact Fibrous

Heat No.-Test No. Temp. "F Energy (ft.lbs.) Fracture {%)

20-7 170 21 100 2C-6 140 19 1/2 100 20-8 110 20 90 20-9 92 18 1/2 60 20-1 75 18 1/2 50 20-10 20 13 1/4 35 20-5 20 14 1/2 35 20-2 -40 13 10 20-3 -100 12 5 20-4 -321 5 1 /4 0

21-7 170 27 100 21-8 110 27 100 21-10 92 26 1/2 100 21-6 75 26 1/4 95 21-1 75 25 1/4 95 21-9 50 23 1/2 75 21-5 20 20 1/2 50 21-11 -10 17 1/2 40 21-2 -40 17 25 21-3 -100 14 10 21-4 -321 5 0

22-6 140 30 1/4 100 22-! 75 29 1/2 100 22-9 63 27 1/2 100 22-7 50 28 100 22-11 35 25 i/2 95 22-10 20 29 80 22-12 5 22 1/2 70 22-8 -10 24 1/4 55 22-2 -40 21 3/4 50 22-3 -100 16 1/4 15 22-4 -321 7 1/2 2

26-6 140 34 1/2 100 26-7 ilO 34 1/4 100 26-1 75 32 1/2 100 26-9 63 32 1/2 100 26-8 50 30 1/4 95 26-10 35 30 1/4 90 26-5 20 26 3/4 80 26-2 -40 21 40 26-3 -100 16 3/4 10 26-4 -321 8 3/4 2

71

TA3LE Al 1 1 (C mti nued)

Soecinen Test Impacc Fibrous Heat No.-Test No. Temp.0F Energy (ft.lbsj Fracture {%)

27-10 170 44 3/4 100 27-6 uo 42 100 27-9 110 4! 1/2 100 27-1 75 39 1/2 85 27-7 50 Vs 60 27-11 20 26 3/4 35 27-8 -10 24 3/4 25 27-2 -itO 19 20 27-3 -100 16 1/2 10 n-k -321 4 0

28-7 170 18 100 28-6 140 17 1/2 90 28-8 no 13 70 28-10 92 16 1/2 50 28-1 75 16 1/4 40 28-11 50 14 30 28-9 20 12 20 28-2 -40 11 1/2 10 28-3 -100 10 3/4 5 28-^ -321 5 0

29-7 no 20 3/4 100 29-1 75 20 1/4 100 29-9 63 18 1/4 90 29-8 50 18 1/2 60 29-11 35 19 1/2 45 29-12 35 19 1/2 45 29-5 20 16 1/2 35 29-10 -10 15 3/4 30 29-2 -40 14 20 29-3 -100 12 1/2 10 29-'» -321 9 2

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TEST TEMPERATURE, F

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TEST TEMPERATURE, 0F

Figure A20, Charpy impact test data for heats 8 and 9.

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TEST TEMPERATURE, F

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TEST TEMPERATURE, F

Charpy impact test data for heats 19 and 2ü

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TEST TEMPERATURE, 0F

Figure A26. Charpy impact test data for heats 26 and 27.

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TEST TEMPERATURE, F

Charpy impact test data for heats 28 and 29.

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v-i urü-. Classification

DOCUMENT CONTROL DATA • R&D t 'ir-unrv .- 199*1 t'.zmujr: ,1 titlm 6odi' ol »bmtra. f »nrf ind*King «nnof*rfcm muaf t:? rr.termd whrn thm cr*r»t: report .\ . lasv:/'pj

l o«■;•'•<» -;s ?.»' ' v|T" v.' tj*.fmtr «u'-^or 2« BCPOK'' SEC_B •• C LASS riCA'iON

Unclass i fied Zt saocP

1 rigp^K- TiT^E

1 ^ f I j^^ce o» -il !oy i nq Elenents on the Toughness of Low Alloy ."jr•.:■••-, i t ; ^ "if,1-; Stre-Tth Stee Is

j« CESCOiP^ .E SO'tS Ivp, -( repc.rf »nd .nc/usii.» d«l».> 1

ri-al Tec"- ical s.eDort July 1967 - Septenber 1968 5 AUTHO« .•> ^.v: .-. rtr" r ttfti mTir tnslial!

.' i - •' - e > -s . , C. a-- E. ■•. Steig er..al d

6 REPC H-r 0*"E 7« TCTALNO OC P»<;r? 76 ^O OFREFS

i •>ve-cer 1968 99 20 | a« c a-, - ■» Ac • 3 0 z c« AV T s r 9 i OBlGtMA^OP'5 K^PO^T S^JMBE« 5

D"AG -6-67- fc PKC,F C ^ SC

>3171 AMMSC CR 68-09(F)

1--3-9

5:25.!3.S42 9b OTMER RE130,'T NO'S/ .'Anv „rbet n^mhers that rrMv be «s<itfne<*

this rvpor!)

10 *V«IL«BIL1TY LIMITATION NOTICES ... | This docunent has been approved for publication j

aid 3a'e; ! ts di st rlout Ion is un1ili ted.

Ill S'JPPL EMENT».»Y NOTES 12 SPONSORING MILITARY ACTIVITY

Amy Materials and Mechanics Research Center, Watertown, Mass. 02172

13 ABSTRACT

T*»i;, s t . !y n*? exa~ined the effect« oi various elements on the notch bend fracture toughness

and C-.irpy inpact Se-avior of a 0.35 C, 3'Ni, fr-Mo-V r-artensitic steel having a roor ten-

per;iture y! eld strength of apo rox i-la te 1 y 160-180 ksi. A classical approach was used in the

des i TI of -i llrvs .vhich permit'.ed a direct evaluation of single element effects rather than

j interact ion«,.

Tte •■'■-.■-. C, *-, Si , Cr, and Mo raised ' -h the notch bend fracture mode transition tem-

perature an d 'he Charpy V-notch transition temperature (100^ fibrosity criterion). In amount;

above that rcguired fo" decxidation and grain refinenert aluninun degraded the transition ten-

pe re t u re n^ d toughness slightly. A nininur toughness occurred at a vanadium content of O.I/,

Over the er .ire range of conpositions examined (1 26 - 6.23) nickel decreased the ransitian

tone era tu re and improved toughness at the lower test temperatures.

C h a r p y ', h e 1 f energy, C . (max), and fracture toughness Kna)< (at li't), did not correlate well

with work h ardenirg exponent (n). Good agreenert was obtained however when these param.icrs

1 were plotted versus true fracture strain. At -32}"^, toughness was essentially fracture

strain in de pendent suggesting that a critical strain criterion based on fracture strain is

valid on!y when fracture is fibrous. 1

A coniparisr n .-.as -,)de of measured ;|r calculated from tensile data u*ing the Hahn-Rosenfield

node 1 K|c - (2/ 3tVn2) 1 /2 i where E = Young's modulus, i = yield strength, Z'' = true fracture p

Strain, n = work hardening exponent. The results Indicated that the increased contribution

on non-duel lip fracture whic'1 accompanies increases in strength and/or decreases in test

tenpe rat urc i r !.).■. alloy steels can lead to larje errors in the predicted toughness.

DD FORM 1473 Unclass i fied Security Classification

Und ass i f ied

Toughness Fracture inpact Tests Transition Tenperature «Iloy ing Crack Propagation Plane Strain Plane Stress £lements Nickel Steel Gun Steeis

^ C - t *

INSTWUCTIONS

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Unclass i fied Security Classiticatton