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Stress-strain relations of HFIR-irradiated austenitic stainless steels
563
Journal of Nuclear Materials 179-181 (1991) 563-567 North-Holland
Stress-strain
relations of HFIR-irradiated
austenitic stainless steels
Shiro Jitsukawa
I, Martin L. Grossbeck 2 and Akimichi Hishinuma 1 ’ Japan Atomic Energy Research Institute, Department of Fuef.r and Materials Research, Tokai-Mum, Ibaraki-ken 319-11, Japan ’ Metals and Ceramics Division, Oak Rigde National Laboratory Oak Ridge, TN 378314376,
USA
Tensile specimens of austenitic staintess steels were irradiated up to 50 dpa at temperature of 573, 673 and 773 K and HFIR. True stress-strain relations derived from the load-displacement curves were analyzed in terms of induced effective strain and strain hardening behavior. The strain hardening exponent of the material was only slightly affected by irradiation at 573 and 673 K. Irradiation at 773 K changes work hardening characteristics. This might have resulted from fine MC precipitation and/or helium produced during irradiation Fracture strain decreases with dose, and the dose dependence was pronounced at 773 K. Heat-to-heat variation was not large between two Ti-modified steels.
procedures
1. In~~uetion
2. Experimental
It is useful to analyze true stress-strain relations of irradiated material to obtain the effect of irradiation on work hardening characteristics and fracture behavior. True stress-strain curves of irradiated alloys were examined by Garr and Pard [l], and de Vries et al. [2]. They examined true stress-strain curves of irradiated 316 strainless steels within the strain range of uniform elongation. True stress-strain relations of irradiated alloys were examined in the present investigation to the point of fracture because of the importance of stress-strain behavior to structural integrity 131. True stress-strain curves were obtained from engineering stress-strain curves of tensile specimens irradiated in the High Flux Isotope Reactor (HFIR). The results indicated that below 673 K, irradiation did not cause a large change in work hardening characteristics in cold-worked alloys. On the other hand, above 773 K, irradiation introduced an apparent increase in the work hardening exponent in such materials, and it was suggested that this change was the result of interaction between dislocations and irradiation produced precipitates and/or He bubbles
Materials tested were US-PCA annealed at 1373 K, aged at 1073 K for 8 h followed and 25% cold-worked; US heat of 20% cold-worked AISI Type 316 stainless steel; and JPCA in solution annealed at 1373 K and cold-worked conditions. In the following, they are designated USPCA-SA, USPCA-B3, US316-CW, JPCA-SA and JPCA-CW. Chemical compositions of these materials are shown elsewhere [7]. Cylindrical tensile specimens 2.0 mm in diameter with an 18.3 mm long reduced section were machined from these materials. Irradiation was performed with a peak thermai neutron flux of approximately 2.5 X 1019 n/m2 s and a fast flux of 1.3 x lOi n/m2 s (E > 0.1 MeV) to fluence levels of 1.7 to 6.0 X 102h n/m2 to give damage levels from 15 to 50 dpa with helium concentrations of 1000 to 4000 appm at temperatures of 573, 673 and 773 K. Tensile tests were conducted at the irradiation temperatures in a vacuum of about 1 X lop3 Pa with strain rates of 4.2 to 5.6 x 10w4/s.
131. In the present study, the effect of irradiation on the true stress-strain relations is examined in terms of heat-to-heat variation, and the effect of irradiation on work hardening characteristics at elevated temperatures is also examined. Some of the data on strengths and elongations. examined in the present experiment have been presented elsewhere [4-61.
Research sponsored by the Japan Atomic Energy Research Institute, Japan and by the Office of Fusion Energy, US Department of Energy, under contract DE-ACOS84OR21400 with the Martin Marietta Energy System, Inc.
3. Analytical
methods
The true stress-strain relation (here after denoted Ts-s) is defined to be the relation between mean tensile stress and mean plastic strain at minimum cross section. However, it was difficult to measure the area at the ~nimum cross section Amin of the irradiated specimen. Therefore, to estimate Amin the following assumptions were made: (i) the volume of the specimen was constant during the tests, (ii) the gage section was uniformly reduced within the plastic strain range to the point of maximum uniform elongation, E: and (iii) necking occurred in the strain range beyond e: and the shape of the neck was approximated by two surfaces of revolution generated by two exponential curves.
0022-3115/91/$03.50 0 1991 - Elsevier Science Publishers B.V. (North-Holland)
564
S. Jitsukawa et al. / Stress-strain relations of HFIR-irradiated aurtenitic SS
(b)
Fig. 1. Schematic illustrations of neck development. Tensile specimen strained: (a) to maximum uniform elongation, r”, and (b) after initiation of necking.
Assumptions (i) and (ii) result in the relation, A, = A&/I, where A, and I, are the original area of the gage section and the gage length, respectively. The variables A,, and I, are instantaneous area of the gage section and instantaneous gage length, respectively. Using this relation, true stress can be obtained from A, and the engineering stress at the corresponding strain in the strain range up to e”,. From assumptions (i) and (iii) a relation between Amin and the elongation after the initiation of necking is derived. Fig. la shows the profile of half of the fractured gage section. The dimensions n, and I, are the radius of the gage section and the gage length at c”,, respectively. Fig. lb shows the profile of
the specimen after the initiation of necking. The radius at the minimum cross section of the gage section, az, was determined by the exponential curve, I.= (exp(x/p) - l), and the position, c, on the x-axis along the tensile direction. The development of the neck was represented by moving the point c in the positive x direction. The gage length. I,, was chosen to satisfy assumption (i) of volume conservation. Thus. the relation between non-uniform elongation ( = /,//, - 1) and A,,,,( = vu:) was approximated. True strain expressed in terms of natural strain at the minimum cross section can be given by ln( A,/A “,,”). The accuracy of this relation is dependent upon the form of the expression for the neck profile, exp( x/p) 1, in fig. 1. Upon examination, only slight differences between the profiles of irradiated and unirradiated specimens were observed. Values of 0.89, 0.87, 0.87 and 0.84 were used for the parameter. p, and applied to the test results at 573, 673, 703 and 773 K, respectively. Values of fracture strain at minimum cross section. er, estimated with assumptions (i) and (iii) were compared with the experimentally obtained values of fracture strain; e; = ln( A,/A, ), where A, is the area of fractured surface and A,, is the original area of the gage section. Agreement between c)t and cr was reasonably good. 4. Results and discussion 4. I. True stress-strain
relations for unirradiated alloys
Fig. 2 shows Ts-s curves for unirradiated JPCA-SA and 56% cold-worked JPCA, and for JPCA-SA irradiated at 673 K to 36 dpa. Also plotted is an estimated stress-strain, U--E curve (dotted) calculated by Ludwig’s power law equation, e = AC”. which is commonly used
TESTED AT
I
Fig. 2. Representative true stress-strain lines represent
the true stress-strain
------ Ludwig’seq. JPCA-SA
curves for unirradiated JPCA-SA and cold worked JPCA, and irradiated JPCA-SA. Dotted
relation calculated by Ludwig’s equation. Plots of true stress against: minimum cross section and (b) c + effective strain ( = c,).
(a) natural
strain ( = C) at
S. Jitsukawa et al. / Stress-strain
to approximate the Ts-s relation of materials. A strength coefficient, A, of 1080 MPa and strain hardening exponent, n, of 0.48 were used for the calculation. Values of A and n were chosen to fit the Ts-s relation of unirradiated JPCA-SA. The agreement between the experimental result and the curve calculated by Ludwig’s equation in fig. 2a was good except for strain ranges below 0.05 and above 1.1. The deviation in the lower strain range may be caused by deformation introduced during machining of the specimens or may reflect a change in flow mechanism [I]. The deviation in the higher strain range above ep (see fig. 2a) might have been caused by the formation of microvoids or internal cracks making the actual area supporting the load smaller than it appears, as indicated by Pugh with respect to copper under high pressure [9]. Despite these small discrepancies Ludwig’s equation can be used to represent the work hardening characteristics of unirradiated JPCA-SA. It is known that the Ts-s relation for cold-worked material can be approximated by Swift’s equation of the form, o=A(~,+r)~,
(1)
where Q is the equivalent strain of cold work prior to the test, and A and n are values for annealed material
P31. Using this relation, good agreement was achieved with 56% cold-worked JPCA above a strain of 0.58 and irradiated JPCA-SA above a strain of 0.71, except in the strain range above cp, as shown in fig. 2b. Good agreement between the Ts-s curve for cold-worked (at ambient temperature) JPCA and the calculated relation in the strain range beyond 0.58 is expected because the temperature dependence of the Ts-s curves is not large below 773 K, and the equivalent strain by cold rolling was 0.58. 4.2. True stress-strain
relation
for irradiated
aIIoys
The Ts-s curve for irradiated JPCA-SA could also be approximated by Swift’s equation using the value of e,, given by co = ( o~,~,/A)“”
- ln(l.Ol),
(2)
where a,,, is a 1% offset proof stress. The value, which gives better agreement with Swift’s equation in the strain range below c,, in fig. 2b is used instead of a 0.2% proof stress as used previously [3] since a 1% offset proof stress. In this analysis, fe is an equivalent strain introduced by irradiation. Irradiation at 573 K to 15 dpa caused an increase in uo.,, to approximately 900 MPa which was followed by a decrease to about 800 MPa during irradiation to 25 dpa. The peak stresses, up and fracture stresses, cr of specimens irradiated to 14 dpa were 1427 MPa and
relations of HFIR-irradiated
austenitic SS
565
1407 MPa, respectively. These values were larger than those of u&radiated specimens. During irradiation to 25 dpa, ei, and or also decreased with dose. Fracture strains were in the strain range of 0.7 to 1.2 (including ce introduced by irradiation) after 14 dpa irradiation, and these values were comparabie to those of unirradiated specimens. Further irradiation induced a slight decrease in ductility. Irradiated USPCA was slightly stronger but less ductile than irradiated JPCA. This analysis introduces the concept of equivalent strain which allows a better assessment of instantaneous plastic behavior of the material. It also permits a more appropriate comparison of irradiated with unirradiated cold-worked and annealed alloys. However, caution must be used in interpreting such large values of strain since co is often a large portion of the total strain, and observed post-irradiation elongation is rather small. The irradiation response of tensile properties at 673 K was qualitatively similar to that at 573 K, however irradiation hardening and loss in ductility were increased compared with those at 573 K. The dose dependence of the tensile properties at 773 K was different. The proof stresses and the fracture strains of the specimens irradiated to 26 dpa were around 700 MPa and 0.48, respectively. Further irradiation up to a dose of 50 dpa caused a continuous decrease in er to 0.1 or less, while a,,,, did not exhibit a large change with dose. Irradiated USPCA were also slightly stronger but less ductile compared with irradiated JPCA. Logarithmic plots of typical results were made to evaluate the effect of irradiation on work hardening characteristics as shown in fig. 3. Curves for unirradiated specimens were also plotted with dotted lines in fig. 3. Straight lines in the figure are Ts-s curves calculated by Ludwig’s power law equation with A and n for unirradiated JPCA-SA. Ts-s curves of unirradiated and irradiated specimens, which were irradiated at 573 and 673 K, agreed well with the straight lines from Ludwig’s equation in the strain range between ce and co + ep. Results for the specimens tested at 573 K and 673 K, demonstrated that work hardening characteristics of irradiated specimens in the strain range below fp could be well represented by eqs. (1) and (2). Results for the specimens tested at 773 K were different. As shown in fig. 3b, the Ts-s curve for the unirradiated cold-worked specimen (labeled No. 4-l) agreed well with Ludwig’s power law relation, while the Ts-s curves of irradiated specimens, using co from Swift’s equation, did not. The stress-strain curves of irradiated specimens lay above the straight line of Ludwig’s equation. The deviation of the stress-strain curves of irradiated specimens have been reported by Hishinuma and Jitsukawa [3], who suggested that fine MC precipitates, formed during irradiation, caused the deviation by increasing the flow stress. This is not inconsistent with the present test results for US316-CW (AA-42) because US316 contains approximately 0.1 wt%
566
S. Jitsukawa
et al. / Stress-strain
relations of HFIR-irradiated
(b)
(a) -
5
<
UXA43,
austenitrc SS
Pldpa(EL-21)
US316-CW, 26dpa
US?CA-E%3.14+a(EL-15) -----JPCA-SA,Odpa
I
773K
---JPCA-SA,Odpa =1060MPa,
n=0.4
500 I
I IIIII 0.5
1.0 c co
f
G
Fig. 3. Logarithmic plots of true stress-strain curves for USPCA-B3 and US316-CW specimens irradiated at (a) 573 K and 673 K. and (b) 773 K. True stresses of an EBR-II irradiated AISI Type 316 stainless steel at yield and ultimate tensile strength [I] are also plotted.
niobium [7], which is known to be a strong MC type carbide forming element. However, the larger deviation observed for US316-CW compared to that for USPCAB3 does not seem to be completely consistent with the contents of MC forming elements in these alloys; USPCA contains 0.25 wtX titanium. This implies that other mechanisms, which impede the dislocation motion were also operating. Transmuted helium produces cavities, and possibly complexes, that may impede the dislocation motion. To examine the role of helium on Ts-s curves, a stress-strain curve for a fast breeder reactor (EBR-II)-irradiated specimen, which contained less than one percent of the helium production rate in the HFIRirradiated specimen [l], has been replotted to examine the role of helium on Ts-s curves. The plots for the yield and the ultimate strengths, which were labeled Y and U, respectively, were shown in fig. 3b. These points were on or very close to the straight line predicted by eq. (1). Irradiation by the EBR-II at 773 K did not induce a large change in the work hardening characteristics. As a result, it may be concluded that transmuted helium affected the flow properties. However, data are sparse, and further work is anticipated to clarify the effect of helium in this temperature range. 5. Conclusions (1) An equation of the form B = A(c, + 6)” was used successfully to predict work hardening characteristics of
both unirradiated and irradiated stainless steels below 673 K. Cold-work and irradiation each introduced an effective strain, co, which accounted for the effects of prior cold-work and irradiation on ductility. (2) The effect of irradiation on the strain hardening exponent, n, was small for temperatures below 673 K. (3) At a temperature of 773 K, irradiation in the HFIR changed work hardening characteristics. It is suggested that this was caused by irradiation-produced fine MC particles and/or transmutation-produced helium. (4) USPCA, irrespective of treatment prior to irradiation, was slightly stronger but less ductile than JPCA after irradiation.
References
PI K.R. Garr and A.G. Pard, ASTM-STP
611 (1976) 72.
PI M.I. de Vries, J.D. Elen, G.J. Tjoa and A. Mastenboek,
in: Irradiation Embrittlement and Creep in Fuel Cladding and Core Components (BNES, London 1972) p. 47. ]31 A. Hishinuma and S. Jitsukawa, J. Nucl. Mater. 169 (1989) 241. A. Hishinuma and M.L. Gross141 M.P. Tanaka, S. Hamada, beck, J. Nucl. Mater. 155-157 (1988) 957. in these Proceedings (ICFRM-4) J. Nucl. (51 M.L. Grossbeck, Mater. 179-181 (1991).
S. Jitsukawa
et al. / Stress-sirain
[6] M.L. Grossbeck, T. Sawai, S. Jitsukawa and L.T. Gibson, ADIP Quarterly Progress Report, DOE/ER-0313/6 (ORNL, Oak Ridge, 1989) p. 259. [7] A.F. Rowcliffe, M.L. Grossbeck and S. Jitsukawa, ADIP
relations of HFIR-irradiated
austenitic SS
Quarterly Progress Report, DOE/ER-0045/12 Oak Ridge, 1984) p. 38. [S] P.B. Mellor, J. Mech. Phys. Solids 5 (1956) 5. [9] H.L.D. Pugh, ASTM-STP 374 (1964) 68.
567 (ORNL,
Journal of Nuclear Materials 179-181 (1991) 563-567 North-Holland
Stress-strain
relations of HFIR-irradiated
austenitic stainless steels
Shiro Jitsukawa
I, Martin L. Grossbeck 2 and Akimichi Hishinuma 1 ’ Japan Atomic Energy Research Institute, Department of Fuef.r and Materials Research, Tokai-Mum, Ibaraki-ken 319-11, Japan ’ Metals and Ceramics Division, Oak Rigde National Laboratory Oak Ridge, TN 378314376,
USA
Tensile specimens of austenitic staintess steels were irradiated up to 50 dpa at temperature of 573, 673 and 773 K and HFIR. True stress-strain relations derived from the load-displacement curves were analyzed in terms of induced effective strain and strain hardening behavior. The strain hardening exponent of the material was only slightly affected by irradiation at 573 and 673 K. Irradiation at 773 K changes work hardening characteristics. This might have resulted from fine MC precipitation and/or helium produced during irradiation Fracture strain decreases with dose, and the dose dependence was pronounced at 773 K. Heat-to-heat variation was not large between two Ti-modified steels.
procedures
1. In~~uetion
2. Experimental
It is useful to analyze true stress-strain relations of irradiated material to obtain the effect of irradiation on work hardening characteristics and fracture behavior. True stress-strain curves of irradiated alloys were examined by Garr and Pard [l], and de Vries et al. [2]. They examined true stress-strain curves of irradiated 316 strainless steels within the strain range of uniform elongation. True stress-strain relations of irradiated alloys were examined in the present investigation to the point of fracture because of the importance of stress-strain behavior to structural integrity 131. True stress-strain curves were obtained from engineering stress-strain curves of tensile specimens irradiated in the High Flux Isotope Reactor (HFIR). The results indicated that below 673 K, irradiation did not cause a large change in work hardening characteristics in cold-worked alloys. On the other hand, above 773 K, irradiation introduced an apparent increase in the work hardening exponent in such materials, and it was suggested that this change was the result of interaction between dislocations and irradiation produced precipitates and/or He bubbles
Materials tested were US-PCA annealed at 1373 K, aged at 1073 K for 8 h followed and 25% cold-worked; US heat of 20% cold-worked AISI Type 316 stainless steel; and JPCA in solution annealed at 1373 K and cold-worked conditions. In the following, they are designated USPCA-SA, USPCA-B3, US316-CW, JPCA-SA and JPCA-CW. Chemical compositions of these materials are shown elsewhere [7]. Cylindrical tensile specimens 2.0 mm in diameter with an 18.3 mm long reduced section were machined from these materials. Irradiation was performed with a peak thermai neutron flux of approximately 2.5 X 1019 n/m2 s and a fast flux of 1.3 x lOi n/m2 s (E > 0.1 MeV) to fluence levels of 1.7 to 6.0 X 102h n/m2 to give damage levels from 15 to 50 dpa with helium concentrations of 1000 to 4000 appm at temperatures of 573, 673 and 773 K. Tensile tests were conducted at the irradiation temperatures in a vacuum of about 1 X lop3 Pa with strain rates of 4.2 to 5.6 x 10w4/s.
131. In the present study, the effect of irradiation on the true stress-strain relations is examined in terms of heat-to-heat variation, and the effect of irradiation on work hardening characteristics at elevated temperatures is also examined. Some of the data on strengths and elongations. examined in the present experiment have been presented elsewhere [4-61.
Research sponsored by the Japan Atomic Energy Research Institute, Japan and by the Office of Fusion Energy, US Department of Energy, under contract DE-ACOS84OR21400 with the Martin Marietta Energy System, Inc.
3. Analytical
methods
The true stress-strain relation (here after denoted Ts-s) is defined to be the relation between mean tensile stress and mean plastic strain at minimum cross section. However, it was difficult to measure the area at the ~nimum cross section Amin of the irradiated specimen. Therefore, to estimate Amin the following assumptions were made: (i) the volume of the specimen was constant during the tests, (ii) the gage section was uniformly reduced within the plastic strain range to the point of maximum uniform elongation, E: and (iii) necking occurred in the strain range beyond e: and the shape of the neck was approximated by two surfaces of revolution generated by two exponential curves.
0022-3115/91/$03.50 0 1991 - Elsevier Science Publishers B.V. (North-Holland)
564
S. Jitsukawa et al. / Stress-strain relations of HFIR-irradiated aurtenitic SS
(b)
Fig. 1. Schematic illustrations of neck development. Tensile specimen strained: (a) to maximum uniform elongation, r”, and (b) after initiation of necking.
Assumptions (i) and (ii) result in the relation, A, = A&/I, where A, and I, are the original area of the gage section and the gage length, respectively. The variables A,, and I, are instantaneous area of the gage section and instantaneous gage length, respectively. Using this relation, true stress can be obtained from A, and the engineering stress at the corresponding strain in the strain range up to e”,. From assumptions (i) and (iii) a relation between Amin and the elongation after the initiation of necking is derived. Fig. la shows the profile of half of the fractured gage section. The dimensions n, and I, are the radius of the gage section and the gage length at c”,, respectively. Fig. lb shows the profile of
the specimen after the initiation of necking. The radius at the minimum cross section of the gage section, az, was determined by the exponential curve, I.= (exp(x/p) - l), and the position, c, on the x-axis along the tensile direction. The development of the neck was represented by moving the point c in the positive x direction. The gage length. I,, was chosen to satisfy assumption (i) of volume conservation. Thus. the relation between non-uniform elongation ( = /,//, - 1) and A,,,,( = vu:) was approximated. True strain expressed in terms of natural strain at the minimum cross section can be given by ln( A,/A “,,”). The accuracy of this relation is dependent upon the form of the expression for the neck profile, exp( x/p) 1, in fig. 1. Upon examination, only slight differences between the profiles of irradiated and unirradiated specimens were observed. Values of 0.89, 0.87, 0.87 and 0.84 were used for the parameter. p, and applied to the test results at 573, 673, 703 and 773 K, respectively. Values of fracture strain at minimum cross section. er, estimated with assumptions (i) and (iii) were compared with the experimentally obtained values of fracture strain; e; = ln( A,/A, ), where A, is the area of fractured surface and A,, is the original area of the gage section. Agreement between c)t and cr was reasonably good. 4. Results and discussion 4. I. True stress-strain
relations for unirradiated alloys
Fig. 2 shows Ts-s curves for unirradiated JPCA-SA and 56% cold-worked JPCA, and for JPCA-SA irradiated at 673 K to 36 dpa. Also plotted is an estimated stress-strain, U--E curve (dotted) calculated by Ludwig’s power law equation, e = AC”. which is commonly used
TESTED AT
I
Fig. 2. Representative true stress-strain lines represent
the true stress-strain
------ Ludwig’seq. JPCA-SA
curves for unirradiated JPCA-SA and cold worked JPCA, and irradiated JPCA-SA. Dotted
relation calculated by Ludwig’s equation. Plots of true stress against: minimum cross section and (b) c + effective strain ( = c,).
(a) natural
strain ( = C) at
S. Jitsukawa et al. / Stress-strain
to approximate the Ts-s relation of materials. A strength coefficient, A, of 1080 MPa and strain hardening exponent, n, of 0.48 were used for the calculation. Values of A and n were chosen to fit the Ts-s relation of unirradiated JPCA-SA. The agreement between the experimental result and the curve calculated by Ludwig’s equation in fig. 2a was good except for strain ranges below 0.05 and above 1.1. The deviation in the lower strain range may be caused by deformation introduced during machining of the specimens or may reflect a change in flow mechanism [I]. The deviation in the higher strain range above ep (see fig. 2a) might have been caused by the formation of microvoids or internal cracks making the actual area supporting the load smaller than it appears, as indicated by Pugh with respect to copper under high pressure [9]. Despite these small discrepancies Ludwig’s equation can be used to represent the work hardening characteristics of unirradiated JPCA-SA. It is known that the Ts-s relation for cold-worked material can be approximated by Swift’s equation of the form, o=A(~,+r)~,
(1)
where Q is the equivalent strain of cold work prior to the test, and A and n are values for annealed material
P31. Using this relation, good agreement was achieved with 56% cold-worked JPCA above a strain of 0.58 and irradiated JPCA-SA above a strain of 0.71, except in the strain range above cp, as shown in fig. 2b. Good agreement between the Ts-s curve for cold-worked (at ambient temperature) JPCA and the calculated relation in the strain range beyond 0.58 is expected because the temperature dependence of the Ts-s curves is not large below 773 K, and the equivalent strain by cold rolling was 0.58. 4.2. True stress-strain
relation
for irradiated
aIIoys
The Ts-s curve for irradiated JPCA-SA could also be approximated by Swift’s equation using the value of e,, given by co = ( o~,~,/A)“”
- ln(l.Ol),
(2)
where a,,, is a 1% offset proof stress. The value, which gives better agreement with Swift’s equation in the strain range below c,, in fig. 2b is used instead of a 0.2% proof stress as used previously [3] since a 1% offset proof stress. In this analysis, fe is an equivalent strain introduced by irradiation. Irradiation at 573 K to 15 dpa caused an increase in uo.,, to approximately 900 MPa which was followed by a decrease to about 800 MPa during irradiation to 25 dpa. The peak stresses, up and fracture stresses, cr of specimens irradiated to 14 dpa were 1427 MPa and
relations of HFIR-irradiated
austenitic SS
565
1407 MPa, respectively. These values were larger than those of u&radiated specimens. During irradiation to 25 dpa, ei, and or also decreased with dose. Fracture strains were in the strain range of 0.7 to 1.2 (including ce introduced by irradiation) after 14 dpa irradiation, and these values were comparabie to those of unirradiated specimens. Further irradiation induced a slight decrease in ductility. Irradiated USPCA was slightly stronger but less ductile than irradiated JPCA. This analysis introduces the concept of equivalent strain which allows a better assessment of instantaneous plastic behavior of the material. It also permits a more appropriate comparison of irradiated with unirradiated cold-worked and annealed alloys. However, caution must be used in interpreting such large values of strain since co is often a large portion of the total strain, and observed post-irradiation elongation is rather small. The irradiation response of tensile properties at 673 K was qualitatively similar to that at 573 K, however irradiation hardening and loss in ductility were increased compared with those at 573 K. The dose dependence of the tensile properties at 773 K was different. The proof stresses and the fracture strains of the specimens irradiated to 26 dpa were around 700 MPa and 0.48, respectively. Further irradiation up to a dose of 50 dpa caused a continuous decrease in er to 0.1 or less, while a,,,, did not exhibit a large change with dose. Irradiated USPCA were also slightly stronger but less ductile compared with irradiated JPCA. Logarithmic plots of typical results were made to evaluate the effect of irradiation on work hardening characteristics as shown in fig. 3. Curves for unirradiated specimens were also plotted with dotted lines in fig. 3. Straight lines in the figure are Ts-s curves calculated by Ludwig’s power law equation with A and n for unirradiated JPCA-SA. Ts-s curves of unirradiated and irradiated specimens, which were irradiated at 573 and 673 K, agreed well with the straight lines from Ludwig’s equation in the strain range between ce and co + ep. Results for the specimens tested at 573 K and 673 K, demonstrated that work hardening characteristics of irradiated specimens in the strain range below fp could be well represented by eqs. (1) and (2). Results for the specimens tested at 773 K were different. As shown in fig. 3b, the Ts-s curve for the unirradiated cold-worked specimen (labeled No. 4-l) agreed well with Ludwig’s power law relation, while the Ts-s curves of irradiated specimens, using co from Swift’s equation, did not. The stress-strain curves of irradiated specimens lay above the straight line of Ludwig’s equation. The deviation of the stress-strain curves of irradiated specimens have been reported by Hishinuma and Jitsukawa [3], who suggested that fine MC precipitates, formed during irradiation, caused the deviation by increasing the flow stress. This is not inconsistent with the present test results for US316-CW (AA-42) because US316 contains approximately 0.1 wt%
566
S. Jitsukawa
et al. / Stress-strain
relations of HFIR-irradiated
(b)
(a) -
5
<
UXA43,
austenitrc SS
Pldpa(EL-21)
US316-CW, 26dpa
US?CA-E%3.14+a(EL-15) -----JPCA-SA,Odpa
I
773K
---JPCA-SA,Odpa =1060MPa,
n=0.4
500 I
I IIIII 0.5
1.0 c co
f
G
Fig. 3. Logarithmic plots of true stress-strain curves for USPCA-B3 and US316-CW specimens irradiated at (a) 573 K and 673 K. and (b) 773 K. True stresses of an EBR-II irradiated AISI Type 316 stainless steel at yield and ultimate tensile strength [I] are also plotted.
niobium [7], which is known to be a strong MC type carbide forming element. However, the larger deviation observed for US316-CW compared to that for USPCAB3 does not seem to be completely consistent with the contents of MC forming elements in these alloys; USPCA contains 0.25 wtX titanium. This implies that other mechanisms, which impede the dislocation motion were also operating. Transmuted helium produces cavities, and possibly complexes, that may impede the dislocation motion. To examine the role of helium on Ts-s curves, a stress-strain curve for a fast breeder reactor (EBR-II)-irradiated specimen, which contained less than one percent of the helium production rate in the HFIRirradiated specimen [l], has been replotted to examine the role of helium on Ts-s curves. The plots for the yield and the ultimate strengths, which were labeled Y and U, respectively, were shown in fig. 3b. These points were on or very close to the straight line predicted by eq. (1). Irradiation by the EBR-II at 773 K did not induce a large change in the work hardening characteristics. As a result, it may be concluded that transmuted helium affected the flow properties. However, data are sparse, and further work is anticipated to clarify the effect of helium in this temperature range. 5. Conclusions (1) An equation of the form B = A(c, + 6)” was used successfully to predict work hardening characteristics of
both unirradiated and irradiated stainless steels below 673 K. Cold-work and irradiation each introduced an effective strain, co, which accounted for the effects of prior cold-work and irradiation on ductility. (2) The effect of irradiation on the strain hardening exponent, n, was small for temperatures below 673 K. (3) At a temperature of 773 K, irradiation in the HFIR changed work hardening characteristics. It is suggested that this was caused by irradiation-produced fine MC particles and/or transmutation-produced helium. (4) USPCA, irrespective of treatment prior to irradiation, was slightly stronger but less ductile than JPCA after irradiation.
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