Journal of Research of the National Bureau of Standards Vol. 45, No. 2, August 1950 Research Paper ,2119

Tensile Properties of Ingot Iron at Low TemperaturesBy Glenn W. Geil and Nesbit L. Carwile

True stress-strain curves obtained in tension tests at temperatures ranging from —196°to +100° C with ingot iron as annealed, normalized, quenched and tempered, hot-rolledand cold-drawn are presented. Numerous simultaneous load and diameter measurementswere made during the entire course of each test. The effects of prior thermal and mechanicaltreatment of the iron on the true stress-strain relations are discussed.

Graphs are presented showing the influence of the testing temperature, the ferrite grainsize, and the initial condition of the iron on the work-hardening characteristics, namely,strain-hardening and strain-aging. * The effects of the above factors on yield stress andultimate stress, and on true stress and true strain at maximum load and at fracture, arepresented. The twinning of iron in the tests at low temperatures is. briefly discussed.

I- Introduction

The ordinary tension test is a commonly usedmethod for determining some of the importantmechanical properties of materials, namely, propor-tional limit, yield stress, ultimate stress, breakingstress, percentage elongation, and percentage reduc-tion of area. The procedure for this test has beenstandardized and is specified in detail in ASTMstandards for the Tension Test, ASTM Standards1946 (E8-46). Stress-strain curves, when drawn,are based on the original dimensions of the specimen.It is desirable, however, to have a tension test revealmore information about the material than is indi-cated by the above-mentioned properties. This wasfirst pointed out by Ludwik [I]1 and discussed insome detail by MacGregor [2], who enumeratedmany of the advantages obtained by determining atrue stress-strain curve for the material tested intension under controlled test conditions. True stressis the load divided by the minimum area prevailingat that instant. MacGregor has shown that truestrain during a tension test may be represented byloge Ao/A, in which Ao and A represent the initialand the current minimum cross-sectional areas,respectively. If the deformation of the specimenis uniform throughout the gage length and no localcontraction occurs before fracture, and assuming thevolume of the specimen remains constant, the strainalso may be represented by loge L/Lo, in which Loand L represent the initial and current gage lengths.However, as shown in a previous paper [4], localcontraction, may even occur during the deformationbetween yield and maximum load, and thus simul-taneous load and diameter measurements during theentire tension test are essential for an accuratedetermination of the true stress-strain curve.

Many investigators [2 to 21] have determined truestress-strain curves for various steels tested intension.Several comprehensive surveys of the literature onthe mechanical properties of metals at low tempera-tures and the flow, fracture, ductility, and work-hardening characteristics have been published re-cently [22 to 28]. Several of the investigators [2, 5to 11] have reported a linear relationship between

1 Figures in brackets indicate the literature references at the end of this paper.

true stress and true strain from the point represent-ing maximum load to fracture of the specimen.Siegle [12] shows true stress-strain curves obtainedwith ingot iron tested in tension at temperaturesranging from —132° to 23° C. The diameter meas-urements were made from photographic images ofthe specimen taken during the test; although somedifficulty was encountered in obtaining good imagesof the test specimen, especially at a temperature ofabout —130° C or lower. Siegle's reported resultsdo not conform to a linear true stress-strain relation-ship from maximum load to fracture. However,only a few experimental points were determinedbetween maximum load and fracture in the tests atsubzero temperatures, and some of the irregularityin the data was ascribed to the difficulty of accu-rately measuring inferior photographic images. Gen-samer, Saibel, Kansom and Lowrie [22] state, "Itis evident that beyond that point of maximum load,the stress-strain curve in the tensile test is compli-cated by the introduction of radial and transversestress components as well as by non-uniformity ofthe axial stress. The fact that the plot of averagestress vs average strain at the minimum section is astraight line from the point of maximum load on-ward is purely fortuitous.77

In previous investigations at this Bureau [16 to 20]in which true stress-strain curves were determinedfor various steels and nonferrous metals tested intension at room temperature, the results, in general,do not conform to a linear stress-strain relation fromthe points representing maximum load to fractureof the specimens.

The purpose of the present investigation was todetermine the influence of the testing temperatureand various metallurgical factors on the true stress-strain curves and other mechanical properties ofingot iron tested in tension in the temperature rangeof -196° to +100° C.

II. Material and Method of Investigation

The material used in this investigation was ingotiron. The principal chemical constituents of thismaterial other than iron, as determined by chemicalanalysis, are (in percent) as follows: carbon, 0.02;manganese, 0.02; phosphorus, 0.005; sulfur, 0.018;

129

silicon, 0.002; copper, 0.10; oxygen, 0.058; nitrogen,0.002; hydrogen, 0.0005. This iron was preparedfrom a single heat and furnished by the manufac-turer in the hot-rolled and cold-drawn conditions,with the cold-drawn rods being reduced from hot-rolled rods without intermediate annealing. Thecold-drawn rods were furnished in two degrees ofcold work, namely 14- and 24-percent reductionof area. The iron was studied in the conditions(1) as annealed, (2) as normalized, (3) as quenchedand tempered, (4) as hot-rolled, and (5) as cold-drawn. The heat treatments that were made at theNational Bureau of Standards and applied exclusivelyto the hot-rolled material are summarized in table 1.Typical structures of some of these materials areshown in figure 1. The number, size, and distribu-tion of the inclusions are considered normal for ingotiron. The numerous oxide inclusions were dispersedas relatively small particles throughout the ferriticmatrix, the massive elongaged inclusions composedessentially of manganese sulfide were elongated inthe direction of rolling. Some precipitation occurredin the parent grain boundaries, especially in thosespecimens that were cooled slowly from 1,750° or1,330° F to room temperature (fig. 1, A, B and C).Iron carbide could be identified in some grain bound-aries of the specimens as annealed (fig. 1, A) and asquenched and tempered (fig. 1, B), and it also ap-peared to be present to a lesser degree in the speci-mens in each of the other conditions. It would beexpected that the relative amount of carbon retainedin solid solution should be a minimum in the speci-mens as annealed or as quenched and tempered,intermediate in the specimens as cold-drawn and amaximum in the specimens as hot-rolled or as nor-malized. The nitrogen content of this iron was

very low (0.002% as determined by a modified Allenmethod), and the amount retained in solid solutionshould be about the same for each of the initial con-ditions. Although the phenomenon of strain aginghas been attributed by some investigators to thepresence of nitrogen, it is expected that this smallamount would be ineffective in producing the markedstrain-aging observed in some of the tension testsmade at temperatures within the range of about+60° to +100° C.

TABLE 1. Heat treatment of ingot iron a

Designation

Annealed

NormalizedQuenched and

tempered

Diameterof rod

in.0.875

.875

b.500

Tem-perature

o F

1,750

1,7501,7001,330

Time

hr1

1H

Remarks

Furnace-cooled to 800° F (20hr) then air-cooled.

Air-cooled.Quenched in iced brine.Furnaced-cooled to 100° F

(45 hr).

1 All specimens were prepared from hot-rolled rods.> Specimens machined to approximately ^ in. in diameter in reduced sec-

er in shoulders priotion and % in. in diameter iy %

prior to heat treatment.

The ferrite grain size varied considerably in eachinitial condition of the iron. The average ASTMgrain size numbers are given in table 2, along withsome of the mechanical properties of the iron at roomtemperature. The ferrite grain size was found to beapproximately the same for the iron in the conditionsas hot-rolled, quenched and tempered, and cold-drawn to 14- and 24-percent reduction in area.However, the grain size of the normalized iron wasslightly larger, and the grain size of the annealed ironwas considerably larger than that of the otherconditions.

TABLE 2. Mechanical properties of ingot iron at room temperature

AnnealedNormalizedQuenched and temperedHot-rolledCold-drawn 14% reduction of

areaCold-drawn 24% reduction, of

area

HardnessRockwell B

24373132

78

86

Ferrite

ASTM No.

2-33-44-54-5

4-5

4-5

Yield stress

Upper

1000 psi16.722 324.029 6

Lower

1000 psi14.720.520.524 7

b60

b72

Ultimatestress

1000 psi41.243.739.743 3

65.0

73.2

True break-ing stress a

1000 psi85.095.288.4

105.0

98.0

100.9

Reductionof area atmaximum

load

%23.723.323.623.2

5.6

4.3

Reductionof area at

initialfracture

%67.069.571.274.9

56.0

49.1

Reductionof area at

finalfracture

%72.174.476. C78.4

61.4

56.1

» At initial fracture.b No drop of beam, stress at 0.2% offset.

Cylindrical tensile specimens with a reducedsection of 2-in. gage length were used. The reducedsection was gradually tapered from each end; thediameter at the midsection of the gage length wasabout 0.003 in. less than that at the ends. A mini-mum gage diameter of 0.438 in. instead of thestandard 0.505 in. was employed to avoid fracture inthe threaded ends of the specimens during testingat low temperatures. The specimens were finishedto the final dimensions by grinding and polishingin the axial direction, to avoid circumferential toolmarks from acting as transverse notches. The ends

of the specimen were machined with %-in. X10threads and the shoulder fillets were machined to aradius of 1.5 in.

A pendulum hydraulic testing machine of 50,000-lbcapacity was used for these tests. To obtain thedesired test temperatures, the specimens, except forthose tested at room temperature, were fully im-mersed in an appropriate liquid contained in acylindrical vessel. One threaded end of the specimenpassed through the bottom of the cylindrical vesselinto the lower adapter of the testing machine.Lead washers were used at this position to seal

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against loss of the bath liquid. The outside of thevessel was covered with several layers of wool felt forthermal insulation. The felt in turn was enclosedwith a covering of thin sheet aluminum to preventcondensation of moisture on the felt. The upperand lower adapters were made as long as the clearanceof the testing machine would permit. When the testtemperature was —78° C or lower, the adjacentportions of the head and base of the testing machinewere surrounded with pulverized solid CO2 tominimize the temperature gradient in the adapters.

For obtaining test temperatures of +100° and+60° C, the bath, water with sodium chromateadded as an inhibitor to corrosion, was heated withan electrical immersion heater. For temperaturesbetween room (27° to 30° C) and —78° C, a bathmixture of equal parts by volume of carbon tetra-chloride and chloroform was cooled by the gradualaddition of small quantities of pulverized CO2.A temperature of —78° C was maintained in thecarbon tetrachloride-chloroform bath with an excessof pulverized CO2. For temperatures of —120°,— 138°, and —154° C, a bath of dichloro-difluoro-methane (Freon 12) was cooled by passing liquidnitrogen through an immersed copper coil; the bathwas stirred constantly with an electric stirrer.A chromel-alumel thermocouple was attached to thespecimens at the upper end of the reduced gagesection for measuring the test temperature. Fortemperatures of —188° and —196° C, the specimenswere fully immersed in baths of liquid air and liquidnitrogen, respectively.

The rate of loading was less than is usual in atension test to avoid local heating of the specimensduring the plastic deformation. The loading was socontrolled that the rate of contraction of the speci-mens beyond the initial yielding was maintained atapproximately 1-percent reduction of area per minute.The maintenance of this rate of contraction of thespecimen after the maximum load had been reachedwas achieved by a gradual closing of the controlvalve on the hydraulic testing machine.

A common type of reduction-of-area gage withwedge anvils was used for determining the change inthe minimum diameter of the specimen during thetension tests at room temperature. However, thisgage could not be used with the tests at other tem-peratures. A reduction-of-area gage of specialdesign, which could function satisfactorily whilelargely submerged in the bath, was constructedfor use at subzero and moderately elevated temper-atures. A detailed description of the gage and itscalibration and manipulation is given in a previouspaper [4]. Changes in diameter of a specimen canbe measured by this instrument with an accuracyof ±0.0001 in.

III. True Stress-Strain Curves

The true stress-strain curves obtained for speci-mens of ingot iron as annealed, normalized, quenchedand tempered, hot-rolled and cold-drawn, tested intension at temperatures ranging from —196° to

+ 100° C, are given in figures 2 to 7. A scale ofactual values of Ap/A is given at the top of each dia-gram for convenience in comparing these resultswith previous results obtained at this Bureau [16 to20], in which the data were plotted on this scale. Thetrue stress values at the initiation of fracture asobtained during test through the use of the reduction-of-area gage will be designated in this paper as thefracture stresses, whereas the points designated ascrosses are the values usually reported as the truebreaking stresses or fracture stresses. The lattervalues are obtained by dividing the load at initialfracture by the minimum area of the specimen asdetermined from diameter measurements made afterthe specimen has been fractured. Fracture of anunnotched specimen tested in tension is usuallyinitiated at its axis, and the crack propagates towardthe periphery; the nfetal at the periphery oftenextends longitudinally and contracts radially duringthe fracture. This "rim effect," varies with the com-position, structure, and condition of the metal, aswell as the temperature during the tension test.Thus, fracture stress values as determined from diam-eter measurements made after fracture of a ductilespecimen may be of little significance.

The true stress-strain curves (figs. 2 to 7) obtainedfor ingot iron specimens tested at temperaturesranging from —196° to .+ 100° C do not conform toa linear relationship for the portion of the curvesbetween the points representing maximum load andeither initial or complete fracture. This portion ofthe curves is absent for all specimens tested at —196°C, except for hot-rolled ingot iron, as fracture oc-curred before the maximum load conditions wereattained. The divergence from a linear relationshipusually increases slightly as the testing temperatureis lowered. The divergence also varies with theinitial condition of the iron and is greater for thenormalized, annealed, and hot-rolled specimens thanfor the quenched and tempered and cold-drawnspecimens.

The plastic deformation was not uniform along thegage length during the testing of the ingot iron, irre-spective of the initial conditions. Some local contrac-tion even occurred during initial yielding of the iron.At a strain value slightly above the yield, the mini-mum diameter of the specimen usually continued todecrease and the region of local contraction extendedlongitudinally along the gage length. In some speci-mens, however, local contraction occurred in twoor more regions. As the strain approached that pro-duced by the ultimate stress, the trend was for thespecimen to become nearly cylindrical throughoutits entire gage length. After the ultimate stress wasattained, the specimen started to contract locally(neck) in the usual manner, and this condition con-tinued to complete fracture. It should be pointedout, however, that the change in deformation withstress in this range, ultimate to fracture, frequentlyoccurred in a cyclic manner as illustrated by theposition of the experimental points given for the hot-rolled specimens tested at —120° and at —138° C(fig. 5).

131

The nonuniform deformation associated with pro-nounced strain-aging is manifested by the serrationsin the true stress-strain curves for some of the testsat +100° C. Oscillations of the pendulum of thetesting machine occurred during the deformationfrom yield to fracture. The magnitude of the oscilla-tions at +100° C varied both with the degree ofstrain and the initial condition of the iron. Forannealed, hot-rolled, and cold-drawn specimens, theoscillations were relatively small throughout the rangeof strain from the initial yielding to the maximum load;the magnitude generally increased slightly during theearlier portion of this range and decreased during thelatter portion to very small values at the maximumload. At strains considerably greater than that atmaximum load, the magnitude of the oscillationssuddenly increased greatly and continued in thismanner as the specimens w^re extended to fracture.The oscillations for the quenched and temperedspecimen were relatively small and were observedonly in the first half of the strain range from initialyielding to maximum load. For the normalizedspecimen, the oscillations were relatively small im-mediately following initial yielding, but increasedcontinuously as the specimen was extended to frac-ture. Only variations in the true stress-strain curvescorresponding to the larger oscillations are shown infigures 2 to 7. Smaller and less uniform oscillationsalso were observed in some of the tests at 60° C.

The change in diameter of the specimen duringeach oscillation could be followed with the reduction-of-area gage. The diameter decreased rapidly dur-ing each drop of the pendulum. It remained almostconstant during the subsequent rise of the pendulumuntil the load reached a critical value and a new oscil-lation was initiated. The strain-aging and strain-hardening of the metal resulting from the rapid defor-mation that accompanied the drop of the pendulumapparently was of sufficient magnitude to preventfurther deformation until the true stress was againincreased to a new critical value by an increase ofthe applied load.

The locus of fracture curves, L, (figs. 2 to 7) aredrawn through the points representing initial fractureat the various temperatures and depict the influenceof the testing temperature on the fracture stress andstrain at fracture of the ingot iron in each of theinitial conditions. The influence of the testingtemperature on the strain at fracture varies withthe initial condition, and this may result in a largechange in the shape of these curves as illustrated bythe double reversal in curvature in those for thenormalized and hot-rolled specimens (figs. 3 and 5)in contrast to the absence of reversals in those forthe other specimens (figs. 2, 4, 6, 7). The influenceof these factors will be reported in more detail laterin the discussion of the diagrams representing theinfluence of the testing temperature on the mechani-cal properties of the iron.

IV. Influence of Temperature on the Work-Hardening Characteristics •

The most satisfactory method of studying thework-hardening of metals is by means of the rela-tion between true stress and true strain in a simpletension test in which the load and the change indiameter of the specimen are accurately measuredthroughout the entire test. As the term "work-hardening" relates to the increase of the true stress,<r, with increase in the true strain, 5, (8=loge Ao/A),the slope of the true stress-strain curve, da/d8), atany strain is a measure of the instantaneous rate ofwork-hardening at that strain. If aging or recov-ery occurs during the tension test, the total work-hardening is the combined effect of these factorsand the ordinary strain-hardening, and da/d8 repre-sents the instantaneous rate of the total work-hardening.

Several investigators [8, 10, 22, 28 to 31] havepresented data indicating that the true stress-strainrelationship for metals tested in tension is parabolicand may be expressed mathematically as follows

<r = k8m, (1)

in which k and m are constants, characteristic of thematerial. According to this relationship, the truestress and true strain values plotted on logarithmiccoordinates should conform to a straight-line graphwith the slope of the line equal to m. The coefficientm is generally designated the strain-hardening co-efficient, and the rate of work-hardening is given by

da(2)

if the relationship expressed in cq 1 is valid.It should be pointed out, however, that many of

the logarithmic true stress-strain curves, presentedas evidence confirming the above parabolic relation-ship (eq 1) were plotted on scales of equal logarith-mic cycle length for both axes, thus giving lines ofvery small slope. Any deviations from a linearrelationship are not detected easily in graphs of thiskind.

Now let any general relationship be assumed be-tween the true stress a and the true strain 5, namelycr=f(8). If <T and 8 are plotted on logarithmic co-ordinates, the slope, b, of such a curve at any pointwould be:

and as

then

, d(loge<r) da Id8b~d(ioge8y- a 1 5'

8=logeAo/A,

dS=-dA,IA.

(3)

(4)

(5)

132

Substituting in eq 3, one obtains

do- Id A , ..

h h i 5 (6)

At the upper and lower yield points and at maxi-um load (where dL=d(<rA) = 0)xxt tut; upper auu. luwtJi yiviu. p

mum load (where dL=d(<rA) = 0)

d<r_ dA

Substituting in eq 6, one obtains

8=b.

(7)

(8)

Therefore, the slope of a logarithmic true stress-strain curve at the upper and lower yield points andat the maximum load is equal to the respective strainsat these points.

By substituting b=8 in eq 3, one obtains the relation

da(9)

Thus, the rate of work-hardening at the upper andlower yield points and at the maximum load is equalto the respective true stresses at these points.

The relation expressed^ in eq 9 has been developedin a slightly different manner by Gensamer [29] andby Voce [21]. The relation, 8=m at the maximumload has been developed by Hollomon [10], althoughit was based on the true stress-strain relationshipgiven in eq 1. However, as shown in the develop-ment presented above (eq 3 to 9), the relationshipsgiven in eq 8 and 9 do not depend upon any specifiedtrue stress-strain relation.

Hollomon [10] presented evidence indicating thatthe relationship (T=k8m was valid for plastic de-formation beyond the end of initial yielding to thatat the maximum load or to strains of about 0.4 forsteels tested in tension. The results obtained inthe present investigation, however, do not confirmthis relationship; the true stress-strain values plottedon logarithmic coordinates do not necessarily showa linear relationship. Some of the data obtainedin tension tests at various temperatures with thehot-rolled ingot iron are presented in figure 8. Itshould be noted that in this figure the length of alogarithmic cycle on the scale of true stress is twicethat for the scale of true strain. This method ofplotting reveals the curvature of these lines. Thedashed lines in this figure are drawn through thepoints U, representing true stress-strain values atthe maximum load, with the slopes of the lines equalto the strains at the maximum load (b=8). Thetrue stress-strain values in the immediate vicinityof the maximum load conform very closely to thesedashed lines. These lines, however, do not passthrough the points, E, representing the end of theinitial yielding.2 Moreover, the true stress-strain

2 In this investigation, the end of the nearly horizontal portion of the load-extension carve associated with the lower yield point was considered as the endof the initial yielding.

values for deformation between the end of initialyielding and maximum load cannot be representedaccurately by a straight line. Similar results wereobtained in tension tests with the ingot iron in theother initial conditions. Thus, no constant work-hardening coefficient can be determined from thesetension tests, and the work-hardening rates cannotbe derived readily from the logarithmic true stress-strain graphs.

The work-hardening rate, d<r/d8, at any strain canbe ascertained from the slope of the true stress-straincurves (figs. 2 to 7) at the designated strain, althoughsome difficulty may be encountered in accuratelydetermining the tangent line to the curve at thedesignated strain value. On the other hand, anaverage work-hardening rate, Aa/A8, over a range ofstrain A5 can be determined readily from the truestress-strain curves. This method has been used inthis investigation to determine the influence oftemperature on the work-hardening characteristicsof the ingot iron during the tension test. The resultsare summarized in figures 9 to 13. As the strain,M, at the minimum point of the true stress-straincurve after initiation of yielding varies with thetemperature of the specimen, it was deemed advisableto choose this point as the initial reference point forthe A5 values. Thus the average work-hardeningrates at the various temperatures for a specifiedstrain range at the small strains are established fromapproximately corresponding portions of the truestress-strain curves. The broken-line curve throughthe completely filled circles in each figure representsthe influence of the testing temperature on the rateof work-hardening at the maximum load, (dcr/d8=<r).Straight lines tangent to the true stress-strain curvesat the maximum load point have slopes that agreevery closely with the value of the true stress at themaximum load. These data are summarized intable 3.

All of the sets of curves in figures 9 to 13 haveseveral features in common. The curves are generallylowered as the strain of the specimen increases; thischange is greatest for the small strains and pro-gressively decreases in magnitude as the strain in-creases. At the larger strains (strains from maxi-mum load to fracture) the decrease in work-hardeningrate with increase in strain is relatively small exceptat the very low temperatures. At very low tempera-tures the curves A for the smallest strain range, Mto M +0.02 drop below some of those for the pro-gressively larger strains. This factor reflects theinfluence of the testing temperature on the initialyielding of the specimen. At room temperature orat slightly elevated temperatures, the degree ofheterogeneous deformation during the initial yieldingwas very small, and the corresponding true stress-strain curves (figs. 2 to 7) rise sharply after reachingthe minimum true stress.3 As the testing temperature

3 The minimum of the true stress-strain curve at which d<rjd8 =0, occurs at aslightly smaller strain than that at the minimum load of the lower yield point.At the latter point the slope of the true stress-strain curve must be positive andequal to the value of the true stress. The physical significance of the portion ofthe curves representing the initial yielding may be questionable, as these curvesare based on average values of the true stress and strain during the heterogeneousdeformation.

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+100 -.+60+30—30—78—120—138—154^ 1 8 8 - .—196

TABLE 3.

Temperature

°C

Rate of work-hardening

As annealed

~dT

655654606578

94

665753606578

qo

of ingot

As normalized

~dT

735655606680

100

a

726056606779

98

iron specimens a

As quenched andtempered

dc a

575052576380

100

585152586379

95

' the maximum load {1,000 psi)

As hot-rolled

d<r a

~dS~

65575560688085

100120130

6857556?687985

10?124130

As cold-drawn 14%reduction of area

~dT

7273687284

100

110

736769768298

112

As cold-drawn 24%reduction of area

d<r*dS

8078788590

107

8077778491

106

* Slope of tangent line to the true stress-strain curve at the maximum load.t> True stress at the maximum load.

is lowered, the strain during the initial yieldingincreases and the curve rises less abruptly beyondthe minimum point. At the low temperatures, therelatively flat portion of the curve is an appreciableproportion of the total strain range A8 from M toM +0.02, and this factor is the cause of the loweringof the work-hardening curves A at the low tempera-tures. The sharp rise in these curves for thetemperature range of room to +100° C is a result ofthe large increase in the magnitude of the strain-aging of the iron during the tension test.

The curves B to / of figures 9 to 13 are not appre-ciably affected by the sharpness of the rise of thetrue stress-strain curves beyond the minimum point,except curve B for the iron in the cold-drawn con-ditions. These curves are representative of theinfluence of the testing temperature on the totalwork-hardening characteristics of the iron for thedesignated strain ranges. They show a generaltrend of a decrease in the rate of total work-harden-ing with increase in temperature from —196° toabout —120° C for some curves and from —154° toabout —120° C for others. As the temperature isincreased from —120° to —30° C, the total work-hardening rates for any specified strain range eitherremains approximately constant or increases. Thework-hardening rates decrease as the temperature isincreased from —30° to about +30° C and increaseagain as the temperature is raised to +100° C.

The rate of total work-hardening of the iron atthe maximum load, irrespective of its initial condi-tion, decreases continuously as the temperature israised from —196° to about +30° C in some casesand from —154° to about +30° C in other cases, andusually increases with further increase of the tem-perature to +100° C. These curves are not acomparison of work-hardening rates at a constantstrain, as the strain at maximum load in each casevaried greatly with the testing temperature. Theyindicate the influence of the testing temperature onthe rate of work-hardening of the iron at the maxi-mum load conditions (eq 7, 8, and 9).

The portions of the total work-hardening curves(B to / in figs. 9 to 12) for temperatures rangingfrom —196° to —120° C for some curves and from— 154° to —120° C for the others manifest the influ-

ence of a decrease in strain-hardening with increasein temperature, as the total work-hardening at thesetemperatures is mainly strain-hardening. How-ever, as the temperature is increased above —120° C,the strain-aging becomes a significantly influencingfactor in increasing the total work-hardening rates,and the magnitude of this factor is as great orgreater than the decrease in the ordinary strain-hardening rates. Above about +30° C, the influenceof the strain-aging predominates, and the total work-hardening rates increase greatly.

Another factor that probably has considerableinfluence on the variation of the rate of work-hardening with temperature is the mechanism ofdeformation. Deformation by twinning (Neumannbands) occurred in the specimens tested in tension attemperatures below- —120° C, but no evidence ofthis was found in specimens tested at temperaturesabove —120° C. The extent of twinning duringthe tension test increased as the temperature waslowered from —120° to —196° C as illustrated bythe microstructures (fig. 14) for the hot-rolled ingotiron. The twinning occured to a similar extent inthe iron tested in the condition's as annealed, normal-ized, and quenched and tempered (fig. 15). Thetwinning in the iron in the conditions as cold-drawnwas restricted to the portion of the specimens adja-cent to the fracture, whereas it occurred throughoutthe entire gage length of the specimens of the ironin the other conditions (figs. 14 and 15). As twin-ning occurred in specimens deformed in tension butnot extended to fracture at the low temperatures,it is considered that this phenomenon in iron is notprimarily associated with the plastic deformationat the instant of fracture or any accompanyingshock waves.

A comparison of graphs of work-hardening ratesversus temperature (figs. 9 to 13) reveals that theinitial condition of the iron influences the rate oftotal work-hardening at the various strain rangesup to about the strain at maximum load. Theamount of carbon initially retained in solid solutionand the size of the ferrite grains apparently affectedthe rate of work-hardening (Ac/A5) in this range ofdeformation. The curves A, B, Cy and D for thequenched and tempered iron (fig. 11) show lower

134

work-hardening rates at most of the testing tempera-tures, especially in the range from about +30° to+ 100° C, than the corresponding curves for thenormalized (fig. 10) or hot-rolled (fig. 12) irons.These lower values for the quenched and temperediron probably are caused mainly by a decrease inthe strain-aging as a result of the smaller amount ofcarbon retained in solid solution. The work-hardening rates were usually higher for the annealed(fig. 9) than for the quenched and tempered iron(fig. 11). As the amount of carbon retained in solidsolution was approximately the same for the ironsin these two conditions, the higher work-hardeningrates of the annealed iron can be attributed to itscoarser ferrite grains (ASTM No. 2 and 4, table 2).The rates for the annealed iron, in general, wereslightly higher than those of either the normalizedor hot-rolled irons, although the amount of carboninitially retained in solid solution of the annealedmaterial was also lower than that of the other twoconditions. This change in the positions of thecurves may be attributed to the predominant effectof the coarser grain size of the annealed iron. Thusthe rise in true stress between yield and maximumload is more rapid for the annealed iron than for thenormalized or hot-rolled irons. The work-hardeningcharacteristics of the normalized (fig. 10) and hot-rolled (fig. 12) irons were approximately alike,except in the temperature range of about +30° to+ 100° C, in which the strain-aging of the normalizediron was appreciably greater. The amount of carbonin solid solution prior to testing was considered tobe about the same for the hot-rolled as for the norm-alized material. The rates of the work-hardeningof the iron reduced 14 and 24 percent in area bycold-drawing, are consider-ably less than those of theiron in the other conditions. The strain-hardeningand strain-aging accompanying the 14- and 24-percent reductions in area of the hot-rolled ironwere greater when the iron was reduced by cold-drawing and subsequently aged in storage than whenthe reduction was obtained in an ordinary tensiontest.

V. Influence of Temperature on theMechanical Properties

The influence of the testing temperature on yieldstress, ultimate stress, true stresses at maximum loadand at fracture, and true strains at maximum loadand at fracture can be readily determined from thetension tests made at the various temperatures andthe true stress-strain curves obtained from thesetests. These data, as summarized in figures 16 to 21,show the relationship of some of these properties. Itis apparent that the ratio of the values of true stressat maximum load to the ultimate stress for thesecurves will depend upon the strain at maximum load.The strains at maximum load varied considerablywith the test temperature and the initial conditionof the iron and usually decreased greatly at thevery low temperatures. Thus some of the pairs ofcurves of ultimate stress, and true stress at maximumload, actually converge at the lowest temperatures.

The dependence of the true stress at fracture on thestrain at fracture is clearly shown in these figuresand will be discussed later.

The combined effects of test temperature andinitial condition of the iron on the above-mentionedmechanical properties are depicted in the compositediagrams of figures 22 to 27. The curves for thelower yield stress (fig. 22) depict closely the effectof temperature on the initial strength (initial re-sistance to flow) of the iron in the various conditions.The effect of strain-aging during the initial yieldingon the yield strength was relatively small or negligi-ble at temperatures in the range 0° to —196° C.The lower yield stress, and hence the initial strengthof the iron, increased continuously with lowering ofthe temperature in this range; this increase in yieldstress with lowering of the temperature from 0° to— 196° C was about 70,000 psi for the iron in eachof the initial conditions. At temperatures in therange of 0° to +100° C, however, the strain-agingof the iron during the initial yielding was usuallyappreciable and is depicted by a leveling off or aslight rise of the curves with increase in temperaturein this range. Thus these curves do not necessarilyindicate an increase in the initial strength of the ironswith increase in temperature from 0° to + 100° C.

The combined effects of the ferrite grain size andthe solubility of carbon on the initial strength of theiron are indicated in the relative positions of theyield stress curves (fig. 22). The hot-rolled iron(ASTM grain No. 4 to 5) and the normalized iron,(ASTM grain No. 3 to 4) should have about thesame amount of carbon retained in solid solution.At each test temperature throughout the range of— 196° to +100° C, however, the yield stress of thehot-rolled material (curve B) is higher than thecorresponding yield stress of the normalized material(curve F). Similarly, curve Ey for the quenchedand tempered iron (ASTM grain No. 4 to 5), ishigher throughout this temperature range thancurve A, for the annealed iron (ASTM grain No. 2to 3), even though both conditions are consideredto have about the same amount of carbon retainedin solid solution. These results show a generaltrend of an increase in yield stress as the ferritegrain size decreases. It is noteworthy, however,that the relatively high yield strengths (curves Cand D of fig. 22) and also the relatively high ultimatestresses and true stresses at maximum load (curvesC and D of figs. 23 and 24) of the cold-drawn ironsare due to the strain-hardening and strain-agingduring the prior cold-drawing operations and theaging during storage. These high-strength valuesare not due to a grain size factor, as the averageferrite grain size was the same for the hot-rolled,quenched and tempered, and cold-drawn conditions.

The ultimate stress (fig. 23) and true stress atmaximum load (fig. 24) both increased as the testtemperature was decreased from about +30° to— 196° C; the true stress at maximum load of theannealed specimens did not change appreciably asthe temperature was decreased from —154° to— 196° C. Both stresses either changed little or

135

increased slightly as the temperature was raised fromabout +30° to +60° C, and both increased consider-ably as the temperature was raised from +60° to+ 100° C, because of the increase with temperatureof the strain-aging accompanying the deformation.The strain-aging varies with the initial condition ofthe iron and especially with the amount of carboninitially retained in solid solution. Thus, the ultimatestresses and true stresses at maximum load for thehot-rolled iron (curves B) are relatively higher attemperatures above —120° C than those for thequenched and tempered iron (curves E), even thoughthe average grain size was the same. These valuesfor the normalized iron (curves F) are relativelyhigher at most of the test temperatures above— 120° C, than those for the annealed iron (curvesA). This would be expected, as the normalized ironinitially had a smaller ferrite grain size and retainedmore carbon in solid solution. However, it is believedthe effect of the ferrite grain size on the strain-agingwas relatively small. The increase in ultimate stressand true stress at maximum load with increase intemperature in the range +60° to +100° C for thecold-drawn iron (curves C and D) was relativelysmall, as appreciable strain-aging had occurredduring cold-drawing and subsequent storage. At thevery low temperatures, —154° to —196° C, therelative positions of the curves of true stress atmaximum load (fig. 24) depend greatly upon thestrain at maximum load; the greater the strain thegreater is the true stress.

The curves representing the influence of testtemperature on true strain at maximum load areshown in figure 25. The strains at maximum loadfor the annealed (curve A), hot-rolled (curve B),normalized (curve F), and quenched and tempered(curve E) irons show certain common features. Thestrains remained approximately constant at testtemperatures ranging from +60° to —30° C and, ingeneral, decreased as the temperature was loweredfrom —30° to —196° C. The curve B for the hot-rolled iron however, exhibits a moderate increase instrain as the temperature is lowered from about— 125° to —160° C; a smaller increase is also notedin the curve F for the normalized iron. At a temper-ature of —196° C, the strains at maximum loadapparently were influenced greatly by the ferritegrain size; the larger the grain size the smaller was thestrain. The strain for the annealed iron was con-siderably smaller than that for the quenched andtempered iron. The iron in both conditions wasconsidered to have about the same amount of carboninitially retained in solid solution. Similarly, thestrain at —196° C for the normalized iron wasconsiderably smaller than that of the hot-rolled iron.

The strains at maximum load for the cold-drawnirons (curves C and D), as a result of the strain-hardening and strain-aging prior to test are verymuch smaller than those for the iron in the otherconditions. It should be pointed out, however, thatthe temperatures at which the strain begins to de-crease rapidly with decrease in temperature are about

100 deg. C lower for the cold-drawn irons than forthe other irons.

The curves representing the influence of test tem-perature on the true strain at fracture are shown infigure 26; the true strain at fracture is a measure ofthe ductility of a metal. Curves A, E, C, and D forthe annealed, quenched and tempered, and colddrawn irons are similar in shape. As the tempera-ture is decreased from +100° to —196° C, the ductil-ity first gradually rises to a maximum and then de-creases continuously to very low values at —196° C.There is no abrupt change in the slope of these curves(A, E) C, and D) and no sharp transition tempera-ture.4 The ductility at fracture of the hot-rolledand normalized irons varied considerably with thetest temperature, and two maxima are. exhibited ineach curve (curves B and F). The ductility wasgreatly reduced as the temperature was lowered from— 30° to —78° for curve F, and to —120° C forcurve B. With further decrease in temperature to— 154° C, however, the ductility increased slightlyand then decreased greatly as the temperature waslowered to —-196° C. As mentioned earlier, thecurves representing the ductility at maximum loadof the iron in these two conditions (fig. 25) exhibitthe same trend, although to a smaller extent. Thischaracteristic is probably associated in some way withthe greater amount of carbon initially retained insolid solution in these irons than in the other irons.

The temperature at which the maximum ductilityis retained depends upon the initial condition of theiron (fig. 26). This temperature is highest for thenormalized iron, curve F, and lowest for the cold-drawn iron, curves C and D. The temperature ofthe maximum ductility of the cold-drawn irons isabout 50° C below that of the hot-rolled iron, curveB. The reduction in the amount of carbon retainedin solid solution by the strain-aging of the iron duringthe cold-drawing operations and storage resulted ina shift of the temperature of maximum ductility toa lower value.

The fracture stress curves (fig. 27) indicate ageneral trend of an increase in fracture stress withdecrease in temperature between +30° and about— 160° C. However, many deviations from thistrend are shown in the curves. All of the stressvalues are decreased as the temperature is loweredfrom about —160° to —196° C because of the largedecrease in ductility. The increase in the fracturestress values accompanying a rise of testing tem-perature from +60° to +100° C is due to the pre-dominant influence of the strain-aging. The curvesB and F for the hot-rolled and normalized ironsexhibit a large decrease in fracture stress values asthe temperature is lowered from about —30° toabout —80° C, a direct result of the rapid decreasein ductility of the specimens as the temperature islowered through this range (fig. 26).

The influence of prior cold-drawing of the hot-rolled iron on the fracture stress, as determined in

4 A transition temperature is generally considered as a narrow temperaturerange in which the fracture changes from a shear to a cleavage type accompaniedby a large decrease in ductility.

136

subsequent tension tests, is also illustrated in figure27 by the shape of the curves B, O, and D. Thestrain-hardening and strain-aging accompanying thecold-drawing of the iron and aging during subsequentstorage did not completely remove all of the charac-teristics specific to the hot-rolled iron. Curve C forthe iron reduced 14 percent in area by cold-drawingexhibits a smaller dip in the region near —78° Cthan that in curve B for the hot-rolled iron. Cold-drawing of the iron to 24-percent reduction in areawas sufficient to eliminate the dip in the fracturestress curve D in the region near — 78° C.

The reduction of the amount of carbon retained insolid solution by a heat treatment of the initiallyhot-rolled iron, such as represented by the quenchedand tempered, or annealed condition, also influencesthe fracture stresses as indicated by a nearly con-tinuous rise of the curves E and A of figure 27 as thetemperature is lowered from about +30° to about-160° C.

VI. Summary

True stress-strain curves for ingot iron in variousconditions and tested in tension at temperaturesranging from —196° to +100° C, are presented.Derived graphs also are presented showing theeffects of the testing temperature and initial condi-tion of the ingot iron on the mechanical propertiesand work-hardening characteristics.

The true stress-strain curves for tension tests attemperatures ranging from —196° to +100° C withannealed, normalized, quenched and tempered, hot-rolled and cold-drawn ingot iron are not linear be-tween maximum load and fracture; the divergencefrom a linear relationship generally increases as thetesting temperature is lowered.

The logarithmic graphs of the true stress-straindata for ingot iron tested in tension at temperaturesranging from —196° to +100° C are not linear, andthus the true stress-strain relationship cannot berepresented accurately by the equation, a=k8m.

The plastic deformation of the ingot iron is notuniform during extension. Some local contractionusually occurs at the initial yielding of the iron andduring deformation between yield and maximumload. The deformation from maximum load to frac-ture is frequently nonuniform, as indicated by acyclic trend in this portion of the true stress-straincurves.

The "rim effect" obtained during the fracture ofingot iron is considerable except at the very lowtesting temperatures. Accurate values of the frac-ture stress, therefore, cannot be determined fromdiameter measurements of the ingot iron specimensafter fracture, except possibly for those tested atvery low temperatures.

Deformation by twinning (Neumann bands) occursduring the deformation of specimens of ingot irontested in tension at temperatures below —120° C,and the amount of twinning depends upon the testingtemperature and the prior strain history of thespecimen; the twinning increases with decrease oftemperature below —120° C and decreases with

prior cold-working of the iron at room temperature.The work-hardening of ingot iron during test in

tension at temperatures below —120° C is mainlystrain-hardening, and the rate of total work-hard-ening increases with decrease in the testing tempera-ture. As the temperature is increased above —120°C, the influence of strain-aging in increasing the totalwork-hardening rate appears to be appreciable.Above room temperature strain-aging is the predom-inating factor, and the total work-hardening rateincreases greatly with increase in temperature.

The effect of the ferrite grain size is manifested inthe yield stress values; the yield stress decreaseswith increase in the grain size. The work-hardeningrates at very small strains also are influenced by theferrite grain size; the rate increases with increase inthe grain size. The difference in the grain size ofthe iron in the various initial conditions, however,was not sufficient to greatly affect the other mechan-ical properties.

The variation of the strains at maximum load andat fracture with temperature, as determined in ten-sion tests, depends upon the initial condition of theiron; it varies with the amount of carbon retainedin solid solution and the prior strain history. Thetransition temperature ranges, as determined byvariations in the strains at fracture, in general, arerelatively broad.

The fracture stress curves (true stress at initialfracture) show a trend of an increase in the fracturestress values with decrease in testing temperaturebelow room temperature. However, several factors,such as strain-hardening, strain-aging, and strain atfracture greatly influence the fracture stresses andresult in many deviations from this trend.

The authors are indebted to C. R. Johnson, R. L.Bloss, Lavaria B. Weinrich, and Fannie A. Wilkinsonfor assistance in this investigation.

VII. References[1] P. Ludwik, Elemente der technologischen Mechanik

(Julius Springer, Berlin, 1909).[2] C. W. MacGregor, The true stress-strain tension test—

its role in modern materials testing, J. Franklin Inst.238, 111, 159 (1944).

[3] E. Voce, True stress-strain curves, Metal Treatment 15,53 (summer 1948).

[4] G. W. Geil and N. L. Carwile, A reduction-of-area gagefor use at low temperatures, J. Research NBS 43,527 (1949) RP2044.

[5] C. W. MacGregor, Relation between stress and reductionin area for tensile tests of metals, Trans. Am. Inst.Min. Met. Engrs. 124, 208 (1937).

[6] M. Gensamer, E. B. Pearsall, and G. V. Smith, Themechanical properties of the isothermal decompositionproducts of austenite, Trans. Am. Soc. Metals 28,380 (1940).

[7] C. E. Lacy and M. Gensamer, The tensile properties ofalloyed ferrites, Trans. Am. Soc. Metals 32, 88 (1944).

[8] J. H. Hollomon, The effect of heat treatment and carboncontent on the work hardening characteristics ofseveral steels, Trans. Am. Soc. Metals 32, 123 (1944).

[9] C. Zener and J. H. Hollomon, Plastic flow and ruptureof metals, Trans. Am. Soc. Metals 33, 163 (1944).

[10] John H. Hollomon, Tensile deformation, Trans. Am.Inst. Min. Met. Engrs. 162, 268 (1945).

137

[11] T. A. Read, H. Markus, and J. M. McCaughey, Plasticflow and rupture of steel at high hardness levels—partof book Fracturing of metals, Am. Soc. Metals (Cleve-land, Ohio, 1948).

[12] L. Siegle, Effect of ferritic grain size on the true stress-strain tensile properties and notch impact strength ofingot iron at low temperatures. Thesis presented tothe Graduate School, Univ. Pennsylvania (1948).

[13] E. J. Ripling, Tensile properties of a heat treated lowalloy steel at subzero temperatures, Am. Soc. MetalsPreprint No. 1 (1949).

[14] E. J. Ripling and G. Sachs, The effect of strain-tempera-ture history on the flow and fracture characteristics ofan annealed steel, J. Metals 1, Sec. 3, 205 (1949).

[15] J. E. Dorn, A. Goldberg, and T. E. Tietz, The effect ofthermal-mechanical history on the strain hardening ofmetals, Trans. Am. Inst. Min. Met. Engrs. 180, 205(1949).

[16] D. J. McAdam, Jr., G. W. Geil, and W. D. Jenkins,Influence of plastic extension and compression on thefracture stress of metals, Trans. Am. Soc. TestingMaterials 47, 554 (1947).

[17] D. J. McAdam, Jr., G. W. Geil, and R. W. Mebs, Influ-ence of plastic deformation, combined stresses and lowtemperatures on the breaking stress of ferritic steels,Trans. Am. Inst. Min. Met. Engrs. 112, 323 (1947).

[18] D. J. McAdam, Jr., G. W. Geil, D. H. Woodard, andW. D. Jenkins, Influence of size and the stress systemon the flow stress and fracture stress of metals, Trans.Am. Inst. Min. Met. Engrs. 180, 363 (1949).

[19] D. J. McAdam, Jr., G. W. Geil, and Frances Jane Crom-well, Flow, fracture and ductility of metals, Trans.Am. Inst. Min. Met. Engrs. 175, 306 (1948).

[20] D. J. McAdam, Jr., G. W. Geil, D. H. Woodard, andW. D. Jenkins, Influence of strain aging on the, fracturestress of low carbon steel, Trans. Am. Inst. Min.Met. Engrs. 176, 436 (1948).

[21] E. Voce, The relationship between stress and strain forhomogeneous deformation, J. Inst. Metals 74, 537(1948).

[22] M. Gensamer, E. Saibel, J. T. Ransom, and R. E.Lowrie, The fracture of metals (Am. Weld. Soc, NewYork, N. Y., 1947).

[23] L. Siegle and R. M. Brick, mechanical properties ofmetals at low temperatures: A survey, Trans. Am.Soc. Metals 40, 813 (1948).

[24] J. H. Hollomon, The problem of fracture, Welding J. 25,Res. Supplement 534s (1946).

[25] H. W. Gillett and F. T. McGuire, Behavior of ferriticsteels at low temperatures, Nat. Research CouncilWar Met. Comm. Reports No. W-78 and W-118(June 1944).

[26] A. E. White and C. A. Siebert, Report on literaturesurvey on low temperature properties of metals (J. W.Edwards, Ann Arbor, Mich., 1947).

[27] E. Orowan, Fracture and strength of solids, Report toMinistry of Supply, United Kingdom (1949).

[28] M. Gensamer, Strength and ductility, Trans. Am. Soc.Metals 36, 30 (1946)

[29] M. Gensamer, The yield point in metals, Trans. Am.Inst. Min. Met. Engrs. 138, 104 (1938).

[30] J. R. Low, Jr. and T. A. Prater, Final report on plasticflow of aluminum aircraft sheet under combined

.loads—II OSRD Report No. 4052 (Aug. 22, 1944).[31] J. R. Low, Jr. and F. Garofalo, Precision determination

of stress-strain curves in the plastic range, Proc. Soc.Exper. Stress Anal. 4, [2], 16 (1947).

138

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139

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FIGURE 3. True stress-strain curves obtained in tension tests FIGURE 5. True stress-strain curves obtained in tension testsat various temperatures with normalized ingot iron. at various temperatures with hot-rolled ingot iron.

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150

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FIGURE 6. True stress-strain curves obtained in tension tests FIGURE 7. True stress-strain curves obtained in tension testsat various temperatures with ingot iron as cold-drawn, 14- at various temperatures with ingot iron as cold-drawn, 24-percent reduction of area. percent reduction of area.

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Effect of testing temperature on the rate of work-hardening of cold-drawn ingot iron.

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from fracture, others all % in. from fracture; etched with, Jrpercent nilal, X,/()().A, -120° C; B, -138° 0; C, -164° 0; i>, -188° C; Band A', -196° 0.

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FiauRB ir>. Micro8tructure of the fracture region, after test in tension at 196° C with ingot iron in various conditions;etched with 4 percent nilal, /[loo.

A, annealed; B, quenched and tempered; C, aormallzed; /', hot-rolled; /'.', cold-drawn, L4-percent redaction of area; F, cold-drawn, 24-percent reductionof area.

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INGOT IRON, ANNEALEDV - - U P P E R YIELD STRESS/ \ - -LOWER YIELD STRESSO - - U L T I M A T E STRESSO - - T R U E STRESS AT MAXIMUM LOAD• - - T R U E STRESS AT FRACTUREA - - T R U E STRAIN AT MAXIMUM LOADD - - T R U E STRAIN AT FRACTURE

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FIGURE 16. Effect of testing temperature on various mechanicalproperties of the annealed ingot iron.

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FIGURE 18. Effect of testing temperature on various mechanicalproperties of the quenched and tempered ingot iron.

FIGURE 19. Effect of testing temperature on various mechanicalproperties of the hot-rolled ingot iron.

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GOT IRON% REDUCTION OF AREASTRESSSTRESS

RESSAT MAXIMUM LOADAT FRACTUREAT MAXIMUM LOADAT FRACTURE

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-100 0TEMPERATURE, °C

FIGURE 20. Effect of testing temperature on various mechanicalproperties of the ingot iron as cold-drawn, 14-percent reductionof area.

-100 0TEMPERATURE, °C

FIGURE 21. Effect of testing temperature on various mechanicalproperties of the ingot iron as cold-drawn, 2^-percent reductionof area.

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140

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INGOT IRONYIELD STRESS

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FIGURE 22. Effect of testing temperature and the initial condi-tion of the ingot iron oh the lower yield stress.

FIGURE 23. Effect of testing temperature and the initial condi-tion of the ingot iron on the ultimate stress.

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-200 -100 0TEMPERATURE, °C

100

-200 -100 0TEMPERATURE. °C

100

FIGURE 24. Effect of testing temperature and the initial condi-tion of the ingot iron on the true stress at maximum load.

FIGURE 25. Effect of testing temperature and the initial condi-tion of the ingot iron on the strain at maximum load.

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— ANNEALED— HOT ROLLED- - C O L D DRAWN. I4%REDUCTION OF AREA

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FIGURE 26. Effect of testing temperature and the initial condi-tion of the ingot iron on the strain at initial fracture.

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INGOT IRONTRUE STRESS AT FRACTURE

—ANNEALED—HOT ROLLED—COLD DRAWN, 14% REDUCTION OF AREA—COLD DRAWN, 24%REDUCTION OF AREA— QUENCHED AND TEMPERED— NORMALIZED

-200 -100 0TEMPERATURE. °C

100

FIGURE 27. Effect of testing temperature and the initial condi-tion of the ingot iron on the true stress at initial fracture.

WASHINGTON, April 11, 1950.

147