Variation of the Fracture Mode in Temper Embrittled 2.25 Cr-1 Mo Steel

JIN YU and C. J. McMAHON, Jr.

Temper embrittled 2.25 Cr-I Mo steel was tested by slow bending of notched specimens at various temperatures, and the fracture mode was examined by SEM fractography. Comparison of the local fracture mode with the load-displacement curves showed that intergranular fracture occurred most prominently in the region where cracking initiated, but that the fracture mode tended to change to cleavage as the cracking propagated and accelerated. When the area fraction of intergranular fracture was plotted as a function of test temperature, a maximum appeared, and the temperature of this maximum tended to increase with specimen hardness. It is argued that the gap between the cleavage fracture stress (or cL) and that of intergranular fracture (o-~ ~) was greatest at some particular temperature, allowing a maximum amount of grain boundary fracture. However, the gap (or cL - o-~) diminished as cracking accelerated, and the fracture mode tended to switch to cleavage. The contrast in behavior between temper embrittled CrMo and NiCr steels is discussed.

I. I N T R O D U C T I O N

THE segregation of group IV B to VI B elements to prior austenite grain boundaries causes temper embrittlement in alloy steels, and the ductile-brittle transition temperature is a widely-accepted tool used to measure the embrittlement susceptibility. An alternative measure is the area fraction of intergranular facets on the fracture surface; both measures are directly related to the amount of segregated impurities on prior austenite grain boundaries. However, the occurrence of grain boundary fracture is sensitive not only to the grain boundary composition, but also to the matrix hardness, ~ grain size, 2 test temperature, 3 etc.

A fractographic survey of partially embrittled specimens may show an inhomogeneous distribution of intergranular facets on the fracture surface, and the fracture mode may vary systematically along the fracture path. For example, in the ICrMoV steel used by Viswanathan and Joshi 3 the local fracture mode changed as a function of distance from the notch root, and the average area fraction of intergranular fracture was highest at a test temperature which depended on microstructure and matrix hardness. The reason why the maximum amount of intergranular fracture occurred at other than the lowest test temperature was not well understood. However, the result suggested that the test temperature and the location of fractographic observations should be selected in a consistent manner if the percent intergranular fracture (pct IG) is to be a useful measure of embrittlement sus- ceptibility. The present report is intended to clarify this matter further.

II. EXPERIMENTAL PROCEDURE

Circumferentially notched cylindrical bars of a phos- phorus-doped 2.25 Cr-1 Mo steel, 6.3 mm diameter, heat treated as shown in Table I and embrittled to achieve near-

JIN YU is Assistant Professor, Department of Materials Science and Engineering, Korea Advanced Institute of Science and Technology, P.O. Box 150 Chongryang, Seoul, Korea. C.J. McMAHON, Jr. is Professor, Department of Materials Science and Engineering, University of Pennsyl- vania, Philadelphia, PA 19104.

Manuscript submitted October 21, 1983.

Table I. Composition and Heat Treatments

Composition: (Wt Pct) C Cr Mo P

0.2* 2.58 1.0" 0.036

Group A Group B

Austenitization 1443 K, 2 h 1223 K, 4 h Grain size ASTM No. 5 ASTM No. 7 Microstmcture 100 pct martensite 100 pct martensite Tempering 923 K, 10 h 923 K, 10 h or

963 K, 13 h Hardness Rc20 Rc20 or Rcl3 Embrittlement step cooling 793 K, 1000 h

Auger PHR (pct) {19.5r 3.5 (Re20) P l z o / F e 7 o 3 - - (Rc 13)

*Nominal value

equilibrium segregation of phosphorus, were tested in slow bending over the temperature range from 77 K to room temperature; the lower temperatures were obtained with a liquid N2 spray. ~,4,5 The end of the cantilever beam specimen was displaced at the rate of 0.085 mm per second, and the loading point was 1.1 cm from the notch root, as reported previously. 4. Scanning electron micrographs at • 300 mag-

*This specimen configuration, originally introduced by Low and co- workers, 8 was used to maximize the number of specimens obtainable from small experimental heats.

nification were taken in various locations over the whole fracture path below the notch tip. The area fraction of inter- granular fracture was measured on SEM micrographs.

I I I . RESULTS

The 2.25 Cr-1 Mo steel of the present work had been the subject of extensive study. 4 The steel showed a transition temperature shift of only 40 K even in the most embrittled condition, which was much smaller than that of a NiCr steel doped with the same amount of phosphorus. ~ The fracture mode was always found to be a mixture of cleavage and intergranular fracture at low temperatures.

METALLURGICAL TRANSACTIONS A VOLUME 16A, JULY 1985-- 1325

FRACTURE MODE CHANGE AS A FUNCTION OF DISTANCE FROM NOTCH ROOT

X= L X= "~" L X = L

1.00 u

X=0 X= L

L X = L

GRAIN SIZE No. 5, HARDNESS Rc 20, STEP COOLED TEST TEMPERATURE 77 K Fig. 1 --Variation of fracture mode as a function of distance from the notch root. Specimen from Group A tested at 77 K.

Preliminary work was conducted using the step cooled specimen, (Group A, Table I) which had a fracture appear- ance transition temperature (FATT) around 293 K. Fracto- graphs from specimens tested at 77 K and at room temperature are shown in Figures 1 and 2, respectively. At 77 K, the fracture mode was a mixture of intergranular separation and cleavage, and the intergranular fracture was most densely concentrated near the notch root area where fracture started. On the other hand, a mix of ductile rupture and intergranular fracture was obtained at room tern-

perature. Presumably, rupture by microvoid coalescence oc- curred in the latter after plastic bending of the specimen prior to the nucleation of brittle fracture. The two examples differ in the fracture mode, but both exhibit an in- homogeneous fracture surface.

Systematic work on the fracture mode variation was per- formed using isothermally aged specimens of smaller grain size at two hardness levels (Group B, Table I). Auger elec- tron spectroscopy revealed statistically equivalent amounts of grain boundary phosphorus on all the specimens (cf.

1326--VOLUME I6A, JULY 1985 METALLURGICAL TRANSACTIONS A

TEST [EMPERATURE: 295 K

•

Fig. 2--As in Fig. 1; specimens tested at room temperature.

X = L

Table 1). Fractographs similar to Figures 1 and 2 were made along the fracture path, and the area fraction of intergranular fracture (pct ]G) was plotted as a function of distance from the notch root at various temperatures, as shown in Figure 3. At both 77 and 123 K the dominant fracture mode switched from intergranular to cleavage during propagation from the notch tip for both hardness levels. At 173 K the fracture mode fol lowed the sequence: intergranular- cleavage-intergranular; this also occurred at 223 K in the Rc20 specimens. At 273 K, fracture began as rupture and switched to intergranular in the Rc20 specimens, whereas rupture occurred at 273 K in the Rcl3 specimens. When the average value of pct IG was plotted v s test tempera- ture, a maximum was present around 173 K as shown in Figures 4(a) and (b). The occurrence of a maximum in the pct IG at a certain test temperature was also noted by Viswanathan and Joshi 3 in their Charpy tests of 1 CrMoV steels. However, in the latter case the fracture mode

followed the general sequence: rupture-intergranular- cleavage-rupture, which was certainly different from the intergranular-cleavage-intergranular sequence found in the present 2.25 Cr-1 Mo tested in slow bending at 173 K, as shown in Figure 5.

Figure 6 shows the loaddisplacement curves of the speci- mens with the matrix hardness of Rc20 tested at various temperatures. It can be seen that the quasi-brittle fracture was catastrophic at 77 and 123 K. However, at 173 and 223 K, the first running crack was arrested within the sam- ple, leaving an unbroken ligament ahead of the crack tip. Thus, the process of crack initiation and propagation had to be repeated for complete fracture. At 273 K this process had to be repeated a number of times. Essentially the same kinds of load-deflection curves were obtained for the specimens with the hardness Rcl3.

From comparisons of the load-displacement curve with the local fracture mode at 173 and 223 K, it was deduced that

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0 . 0 4 % P ASTM. GRAIN SIZE No. 7 AGED IO00 h AT 7 9 3 K

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( b ) HARDNESS Rc 13

TEST TEMP K

�9 - 77

�9 123

�9 173

x 2 2 3

o 2 7 3

0 L/2 L

Fig. 3 - - P c t IG at three locations at several test temperatures. Specimen from Group B; hardness level: (a) Rc20 and (b) Rcl3.

x

the intergranular fracture mode occurred preferentially at the crack initiation and re-initiation sites, but that cleavage oc- curred preferentially during rapid crack propagation. The tendency for intergranular fracture to occur maximally at 173 K was revealed more strongly when fractographs were taken only from the notch tip area where fracture initiated, as shown in Figure 7. This tendency appears less strong when the pct IG is averaged over the specimen, as in Figure 4.

IV. DISCUSSION

In the present temper embrittled CrMo steel and in the CrMoV steel studied previously by Viswanathan and Joshi 3 two effects were observed which are absent in steels which contain a significant amount of nickel (i.e., several percent). First, the amount of intergranular fracture is nonuniform over the fracture surface, being greatest in regions of frac- ture initiation, and, second, it passes through a maximum at a particular test temperature. We believe these effects are related and that they can be understood in terms of the factors which raise the applied stress necessary for plastic flow in a ferritic steel: a temperature reduction, an increased strain rate (or crack growth rate, for crack-tip plasticity), and the usual carbon/carbide-related hardening.

First, let us examine the general differences between the present CrMo steel and a NiCr steel, using the analysis of the latter of Kameda and McMahon, 6 who measured the

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r r

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I00 150 200 250

4 20

T E S T TEMPERATURE ( K ) Fig. 4- -Average pct IG from Fig. 3 as a function of the test temperature.

local fracture stress in notched bars as a function of tem- perature and the degree of temper embrittlement. As shown earlier by Knott and Cottrell, 7 the fracture stress in a notched specimen nearly coincides with the stress for initial yielding (at the notch tip) at low temperatures, but as the test tem- perature is increased and the yield stress drops, the fracture stress increases until it exceeds the general yield stress. This is shown schematically for both the unembrittled and em- brittled conditions in a NiCr steel 6 in Figure 8(a).

The present CrMo steel differs from the NiCr steel in two important ways: first, because it lacks nickel, its toughness, or resistance to cleavage fracture, is lower. Hence, the curve for the unembrittled condition is moved to a higher tem- perature range in Figure 8(b). Second, because it contains molybdenum, which partially scavenges embrittling ele- ments like phosphorus, 4 it suffers less temper embrittlement than the NiCr steel. Hence, the curve for the embrittled condition is moved to a lower temperature range than in the NiCr steel (Figure 8(b)). This narrows the difference be- tween the unembrittled and embrittled conditions in the CrMo steel and alters the type of fracture-mode transition produced by temper embrittlement.

As indicated schematically in Figure 8(a), the fracture stress curves for cleavage and for intergranular fracture in a

1328--VOLUME 16A, JULY 1985 METALLURGICAL TRANSACTIONS A

I00 p

X z O

X = L

Fig. 5 - -Frac ture mode in specimen from Group B with hardness level Rc20, tested at 173 K.

strongly embrittled NiCr steel are widely separated, and temper embrittlement produces a transition from rupture to intergranular fracture at most test temperatures. However, because of the considerations described above, the transition is from cleavage to intergranular fractures over a substantial range of temperature in the CrMo steel, and we envision the following: at temperature T~ in the CrMo steel the two brittle fracture modes compete almost evenly, and at temperature T3, where general yield precedes fracture, the occurrence of plastic rupture limits the amount of intergranular fracture. However, the intermediate temperature T2 is too low for rupture to occur, and the difference between the cleavage stress and the intergranular fracture stress is large enough so

that the latter is strongly favored. In this way the presence of an optimum temperature for intergranular fracture in the CrMo steel can be rationalized.

The tendency for a switch from intergranular to cleavage fracture, once rapid crack propagation starts in the CrMo steel, can also be understood in terms of Figure 8. Since yielding in a ferritic steel is a thermally-activated process, an increase in strain rate produces the same effect as a decrease in temperature. Once crack propagation begins, the relevant volume is that around the crack tip, and the relevant "strain rate" is then proportional to the crack velocity. Hence, acceleration of fracture brings the two brittle fracture modes into competition, just as would occur at a lower test

METALLURGICAL TRANSACTIONS A VOLUME 16A, JULY 1985-- 1329

o; C3 g J

0 . 0 4 P A S T M GRAIN SIZE N=7 HARDNESS Re 2 0 AGED t O 0 0 h AT 7 9 3 K

2

0 ' ~

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1 2 3 K

173K

2 2 3 K

2 7 3 K

0 I 0 .2 0.3 0.4

D I S P L A C E M E N T A T O . 5 i n F R O M N O T C H ( i nch )

Fig. 6 - - Load-displacement curves of specimens from Group B with hard- ness level Rc20 at several test temperatures.

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/ / I

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i I I 11~ I

I00 150 200 250 TEST TEMPERATURE ( K )

Fig. 7 - - Variation of the pct IG at the notch root with the test temperature for specimen from Group B with hardness level (a) Rc20 and (b) Rcl3.

1330--VOLUME 16A, JULY 1985

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TEST TEMPERATURE Fig. 8--Schematic plots of fracture stresses v s test temperature for (a) NiCr and (b) CrMo steels in the unembrittled and embrittled conditions.

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L~

Io 300

2O0

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2 0 30 4 0 I I I ! I

V i s w n n a t h a n ~ dosh i

I CrMoV - - 0 . 0 2 5 P

GRAIN SIZE

PRESENT WORK A S T M No.7

2 1 Cr - I M o - O . 0 4 P

I I I I 2 0 0 250 300 350

HARDNESS ( DPH ) Fig. 9--Variation of the test temperature to give maximum pct IG with the specimen hardness.

temperature. In this case the cleavage mode presumably dominates for topological reasons, since it can produce a fracture surface of less total area than is required for inter- granular fracture.

The effect of matrix hardness can also be rationalized in terms of Figure 8. Increasing the hardness simply raises the curves for initial and general yielding. By the arguments given above for the CrMo steel, one would then expect the temperature at which the maximum amount of intergranular fracture occurs to increase with hardness. This tendency is seen to some degree in Figures 4 and 7 and also in Figure 9, in which the present results are compared with those of

METALLURGICAL TRANSACTIONS A

Viswanathan and Joshi. 3 However, in the latter case allow- ance must also be made for the greater strain rate in their impact tests.

V. SUMMARY

There are important differences between NiCr and CrMo steel as far as temper embrittlement is concerned. An under- standing of these differences can enable one to understand the nonuniform distribution of intergranular fracture in a temper embrittled CrMo steel, as well as the occurrence of an optimum temperature for intergranular fracture in this type of steel. These factors must be kept in mind when using the fracture appearance in a CrMo or CrMoV steel to assess the extent of temper embrittlement. They also show why temper embrittlement in such a steel can be of significance even when intergranular fracture comprises only a minor portion of the overall fracture appearance.

A C K N O W L E D G M E N T S

Support for this work was received from the Pennsylvania Sc ience and E n g i n e e r i n g F o u n d a t i o n under Gran t No. 242 and from the National Science Foundation MRL program at the University of Pennsylvania under Grant No. DMR76-80994.

REFERENCES

1. R. A. Mulford, C. J. McMahon, Jr., D. P. Pope, and H. C. Feng: Metall. Trans. A, 1976, vol. 7A, p. 1183.

2. J. M. Capus: J. Iron Steel Inst., 1962, vol. 200, p. 922. 3. R. Viswanathan and A. Joshi: Metall. Trans. A, 1975, vol. 6A, p. 2289. 4. J. Yu and C.J . McMahon, Jr.: Metall. Trans. A, 1980, vol. I IA,

p. 277. 5. B.J . Schulz and C.J . McMahon, Jr.: in Alloy Effects in Temper

Embrittlement, ASTM STP 499, 1972, p, 104. 6. J. Kameda and C.J. McMahon, Jr.: Metall. Trans. A, 1980, vol. I1A,

p. 91. 7. J.F. Knott and A. H. Cottre|l: lnl. Iron and Steel Inst., 1963, vol. 201,

p. 249. 8. J. R. Low, Jr., D.R. Stein, A. M. Turkalo, and R. R Laforce: Trans.

TMS-AIME, 1968, vol. 242, p. 14.

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