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Effect of microstructure evolution on fracture toughness in isothermally aged austenitic stainless steels for cryogenic applications
Cryogenics 40 (2000) 693±700
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Eect of microstructure evolution on fracture toughness in isothermally aged austenitic stainless steels for cryogenic applications Maribel L. Saucedo-Mu~ noz a,b,*, Yutaka Watanabe c, Tetsuo Shoji a, Hideaki Takahashi a a Fracture Research Institute, Tohoku University, Aoba-Ku, Sendai 980, Japan Instituto Politecnico Nacional, ESIQIE, Apartado Postal 75-556, D.F. 07300, Mexico Department of Machine Intelligence and Systems, Graduate School of Engineering, Tohoku University, Aoba-Ku, Sendai 980, Japan b
c
Received 8 September 2000; accepted 15 December 2000
Abstract Two types of austenitic stainless steels JJ1 and JN1 were isothermally aged at temperatures from 600°C to 900°C for 10±1000 min in order to study the microstructural evolution and its eect on fracture toughness at cryogenic temperatures. These steels were developed for applications in the superconducting magnets of a fusion experimental reactor by the Japan Atomic Energy Research Institute. The Charpy V-Notch (CVN) fracture energy at 77 K showed a signi®cant decrease with aging time for both steels. The intergranular precipitation of carbides and nitrides is responsible for the fracture toughness deterioration. The scanning electron microscope (SEM) fractographs showed an intergranular brittle fracture and its fraction also increased with aging time and temperature. The presence of a more abundant intergranular precipitation resulted in a more rapid decrease in fracture toughness with aging time in JN1 steel due to its higher content of C and N, compared to that of JJ1 steel. The volume fraction of precipitates can be uniquely correlated with the reduction in toughness. Ó 2001 Elsevier Science Ltd. All rights reserved. Keywords: Austenitic stainless steels; Fracture toughness; Cryogenic; Microstructure
1. Introduction The 300 series austenitic stainless steels have been used for many years in the construction of components employed in chemical, petroleum, and nuclear power industries due to its corrosion-resistant and mechanical properties [1]. These types of steels also have wide application for cryogenic components, although they were not originally developed for cryogenic use. In fact, a number of cryogenic structures are constructed with 304L or 316L steels [1]. A high level combination of strength and toughness ry > 1200 MPa and JIC > 200 MPa m 1=2 at cryogenic temperatures is required for components for superconducting magnets of fusion reactors because the components need to sustain large electromagnetic forces. This requirement is beyond the capability of conventional steels. The Japan Atomic Energy Research Institute (JAERI) started the development of new cryogenic structural materials in 1982 [2±
*
Corresponding author. Fax: +52-5119-1986. E-mail address: [email protected] (M.L. SaucedoMu~ noz).
4]. The newly developed steels have been used as structural materials for the demo poloidal coils (DPS). Thick plates are used for coil supports and thin plates for the conduit materials and the cryogenic buer tank. Most of the components are exposed to thermal cycles during manufacturing due to the welding processes. Additionally, since superconductors such as Nb3 Al require a reaction heat treatment, the conduit material is also exposed to the heat treatment [3]. As is well known, these thermal processes can promote the intergranular and transgranular precipitation of carbides, nitrides and other phases. The precipitation process in austenitic stainless steel is extremely complex. For example, eighteen dierent precipitate phases were reported in an isothermally aged type 316 steel [5]. Furthermore, the same authors pointed out that the exposure of this type of steel to relatively high temperatures (about 650°C) resulted in a deterioration of the tensile properties, mainly a loss of ductility. Recently, Simmons et al. [7] studied the precipitation behavior in high-nitrogen austenitic steels at elevated temperatures. The authors also reported a drastic decrease of impact properties due to the precipitation of Cr2 N. These precipitates on grain boundaries were responsible for the intergranular
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M.L. Saucedo-Mu~noz et al. / Cryogenics 40 (2000) 693±700
fracture because of the reduction of cohesive strength of grain boundaries. This type of microstructure deterioration also decreases fracture toughness of steels [6±8]. In the present work, two types of austenitic stainless steels, JJ1 and JN1, developed for applications to the superconducting magnets of fusion experimental reactor by JAERI [2], were selected to study the microstructure evolution in isothermal aging and its eect on fracture toughness. Aging temperatures from 600°C to 900°C and short aging times from 10 to 1000 min were selected in order to cover the microstructural evolution, which can be compared to that caused either during the welding process or reaction heat-treatment.
ysis of precipitates was also conducted using the extraction replica technique [5]. 3. Results and discussion 3.1. Toughness reduction due to thermal aging Figs. 1(a)±(c) show the plots of CVN fracture energy at 77 K versus aging time for the JN1 and JJ1 steels aged at 700°C, 800°C and 900°C, respectively. Both steels
2. Materials and experimental procedure Materials used in this work were forged-steel plates of 200 mm thick and their chemical compositions are shown in Table 1. Table 2 shows the mechanical properties of this type of steel at 4 K. The solution treatment of JN1 and JJ1 was carried out at 1075°C and 1050°C, respectively, for 1 h under an argon atmosphere, and then water-quenched. The aging temperatures and times were 600°C, 700°C, 800°C and 900°C and from 10 to 1000 min, respectively. The Charpy V-Notch (CVN) test was conducted at 196°C (77 K), following JIS Z2242 standard [9]. The fracture surface of the CVN-tested specimens and microstructure of the aged specimens were observed with a scanning electron microscope (SEM). The precipitates in the aged samples were extracted electrolitically by dissolution of austenitic matrix in a solution of 10 vol.%HCl±CH3 OH at 4 V. The X-ray diraction pattern of extracted precipitates was measured in a SIEMENS diractometer using KaCu radiation. The TEM microanalysis and observation of precipitates were carried out using a JEOL FX2000 microscope at 200 kV, equipped with energy dispersion X-ray analysis (EDX) facility. The TEM specimens were prepared by electropolishing in a 5 vol.% perchloric acid±20 vol.% glycerin±75 vol.% ethanol electrolyte at 25 V at room temperature. The SEM/EDX microanal-
Fig. 1. CVN fracture energy at 77 K as a function of aging time for JN1 and JJ1 steels aged at (a) 700°C, (b) 800°C and (c) 900°C.
Table 1 Chemical composition of materials (wt%) Material
C
Si
Mn
P
S
Ni
Cr
Al
N
Mo
JN1 JJ1
0.040 0.025
0.97 0.48
3.88 10.13
0.022 0.021
0.001 0.002
15.07 11.79
24.32 12.01
0.023 ±
0.32 0.236
± 4.94
Table 2 Mechanical properties at 4 K of solution treated JN1 and JJ1 steels Material
ry (MPa)
ruts (MPa)
Elongation (%)
JIC kJ m 2
JN1 JJ1
1363 1169
1752 1591
31 39
346 383
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695
showed a monotonotic decrease in the CVN fracture energy with aging time at all the three temperatures. It is also evident that the drop of fracture toughness of JN1 steel is always faster than that of JJ1 steel. The fastest drop of fracture toughness occurred in the JN1 steel samples aged at 900°C. This fact may be attributed to the higher content of C and N in JN1 steel, which can lead to faster kinetics in intergranular precipitation during the aging process, as discussed in a later section. The decrease in CVN fracture energy in aged samples is presented as a function of Larson±Miller parameter (LMP) for JN1 and JJ1 steels in Figs. 2(a) and (b), respectively. LMP is de®ned by the following equation: LMP T C log t;
Fig. 2. CVN fracture energy at 77 K as a function of LMP for (a) JN1 and (b) JJ1 steels.
1
where T is the temperature in K, t is the time in hours and C is a constant, which depends on the material. For these steels, the constant C that best ®ts the experimental data were 8 and 5 for JN and JJ steels, respectively. These plots enable us to predict the fracture toughness for any given aging condition. The typical fracture appearances of CVN test specimens of JN1 are shown in Figs. 3(a)±(i) for the solution
Fig. 3. SEM fractographs of JN1 steel (a) solution treated and aged for (b) 10, (c) 100 and (d) 1000 min at 700°C, for (e) 10, (f) 100 and (g) 1000 min at 800°C and for (h) 10, (i) 100 and (j) 1000 min at 900°C.
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Fig. 4. SEM fractographs of JJ1 steel (a) solution treated and aged for (b) 10, (c) 100 and (d) 1000 min at 700°C, for (e) 10, (f) 100 and (g) 1000 min at 800°C and for (h) 10, (i) 100 and (j) 1000 min at 900°C.
treated condition and the various aged conditions. The fracture morphology of JJ1 is shown in Figs. 4(a)±(i) in the same manner. Both of JN1 and JJ1 steels fractured in a ductile manner in the solution treated condition. Intergranular facets were found from all the aged samples, although area fraction of intergranular facets to ductile surface was strongly dependent on aging conditions. The fraction of intergranular brittle fracture increased with aging time and temperature, and seemed consistent with the CVN fracture energy value. Simmons et al. [7] found that the fracture energy decreased with aging time and temperature in a highnitrogen austenitic stainless steel, nominally Fe±18Cr± 5Mn±5Ni±3Mo±0.02C±0.7N. The Cr2 N intergranular precipitation was the main cause for fracture toughness deterioration in this particular steel. 3.2. Microstructural evolution during thermal aging The typical microstructures of JN1 steel aged at 700°C, 800°C and 900°C for 10, 100 and 1000 min are shown in Figs. 5(a)±(i). The samples aged at 700°C for up to 1000 min have precipitates only at grain boundaries, Figs. 5(a)±(c). At 800°C and 900°C, precipitates ®rst formed at grain boundaries, Figs. 5(d), (e) and (g),
and then followed by intragranular precipitation, Figs. 5(f), (h) and (i). Both the precipitations at the grain boundary and in the matrix were more pronounced for longer aging times. The intragranular precipitates can be classi®ed into two types: cellular or discontinuous precipitation and plate-like precipitates, which have a preferred alignment with the austenitic matrix. The morphology of cellular precipitates is similar to that of pearlite in carbon-steels, Fig. 5(i). The formation of this lamellar microstructure initiated at grain boundaries and grew into the austenite matrix, according to the following reaction [7]: c ! c Cr2 N:
2
The volume fraction of the discontinuous precipitation increased with aging time and the maximum value was determined by the point-count grid method [10], to be about 0.04. This value seems to be reasonable, since a volume fraction of 0.1 was reported in an austenitic stainless steel containing 0.42% N, after a long aging [11]. Fig. 6(a) shows the X-ray diraction pattern of extracted precipitates for the JN1 steel aged at 900°C for 300 and 1000 min. It is evident that the precipitates are mostly the Cr2 N and M23 C6 phases for these aging
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Fig. 5. SEM micrographs of JN1 steels aged for (a) 10, (b) 100 and (c) 1000 min at 700°C; for (d) 10, (e) 100 and (f) 1000 min at 800°C and for (g) 10, (h) 100 and (i) 1000 min at 900°C.
conditions. The types of precipitates were found to be the same also for the samples aged at 700°C and 800°C. At 900°C, the relative intensity of XRD peaks for Cr2 N phase increased with respect to those for M23 C6 with aging time. This fact suggests that the Cr2 N precipitation is favored at 900°C and the M23 C6 precipitation is facilitated at temperatures between 700°C and 800°C because the relative intensity of M23 C6 is higher than that of Cr2 N for these aging temperatures. Other works [12,13] have reported that the maximum rates of Cr2 N and M23 C6 precipitation are achieved at about 900°C and 800°C, respectively, in high-nitrogen austenitic stainless steels. The microstructures of JJ1 steel aged at 700°C, 800°C and 900°C for 100, 500 and 1000 min are shown in Figs. 7(a)±(i). The grain boundary precipitation again preceded the intragranular one. The aging treatments for up to 1000 min at any aging temperature indicated no cellular precipitation in this steel. It is also evident that there is a more abundant precipitation in JJ1 steel than that in JN1 steel for the samples aged at 800°C and 900°C for 1000 min. This fact seems to be inconsistent with the expected precipitation behavior in JN1 steel, based on the content of interstitial solutes. That is, JN1 steel has N and C contents higher than those of JJ1 steel
and thus the volume fraction of carbides or nitrides can be higher in the former case [14]. However, as recognized in the XRD results, the precipitation of g Fe2 Mo phase took place in the late stages of aging process for JJ1 steel, promoting a higher volume fraction for this aging condition. It is also important to notice that the intergranular precipitation is more uniform and abundant in JN1 steel, as expected due to its higher interstitial solute content. Fig. 6(b) shows XRD patterns of aged JJ1 steel samples aged at 900°C for 300 and 1000 min. The precipitates consist of Cr2 N; M23 C6 and g Fe2 Mo phases at 800°C and 900°C and are mainly composed of Cr2 N and M23 C6 phases at 700°C. The precipitation of g phase was detected in samples aged at 800°C for 300 min or longer and at 900°C for 100 min or longer. The intensities of XRD peaks for the Cr2 N and g phases are higher than that of the M23 C6 phase in JJ1 samples aged at 800°C and 900°C in contrast with JN1 steel. This suggests that precipitation of Cr2 N and g phases is predominant in JJ1 steel at these temperatures. The formation of g phase has been reported to occur only in austenitic stainless steels with Mo. For example, the 316type steel aged at temperatures between 250°C and 900°C for times longer than 6000 min [13].
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(a)
(b) Fig. 6. XRD patterns of (a) JN1 and (b) JJ1 steel aged at 900°C for 300 and 1000 min.
Fig. 7. SEM micrographs of JJ1 steels aged for (a) 100, (b) 500 and (c) 1000 min at 700°C; for (d) 100, (e) 500 and (f) 1000 min at 800°C for (g) 100, (h) 500 and (i) 1000 min at 900°C.
M.L. Saucedo-Mu~noz et al. / Cryogenics 40 (2000) 693±700
699
(a)
Fig. 9. CVN fracture energy at 77 K versus volume fraction of precipitates for JN1 and JJ1 steels.
cipitation of 0.03. That means that the intergranular precipitation is the main cause of fracture toughness deterioration during aging of both steels. This fact is also con®rmed by the presence of intergranular brittle fracture and its increase with aging time. (b) Fig. 8. TTP diagrams of (a) JN1 and (b) JJ1 steels.
The SEM and TEM microanalyses of the precipitates showed that the intergranular M23 C6 precipitates are not only composed of Cr, but also of Fe in JN1 steel and of Fe and Mo in JJ1 steels. Furthermore, Mo is the most enriched element in M23 C6 precipitates of the JJ1 steel. The Cr2 N phase of cellular precipitation was mainly composed of Cr. The intragranular Cr2 N and g phases consisted of Cr and Fe and of Fe and Mo, respectively. All the above results are summarized in the time± temperature±precipitation (TTP) diagrams of JN1 and JJ1 steels, as shown in Figs. 8(a) and (b). In general, it can be noticed that the kinetics of precipitation for JN1 steel are faster than that of JJ1 steel, because of its higher interstitial solute content [14]. The TTP diagrams show that the intergranular precipitation of M23 C6 and M2 N preceded the intragranular precipitation of M2 N and M2 N and g phase in JN1 and JJ1 steels, respectively. 3.3. Correlation between embrittlement and precipitation Fig. 9 shows the plot of CVN fracture energy at 77 K as a function of the total volume fraction of precipitation for JN1 and JJ1 steels. This ®gure shows clearly that fracture energy decreased with increasing precipitate volume fraction. It is also noticed that a volume fraction of precipitation as small as about 0.03 caused a drastical decrease in the fracture energy for JN1 and JJ1 steels. It is important to mention that the intergranular precipitation covers almost all grain boundaries for the case of a sample with a total volume fraction of pre-
4. Conclusions The thermal aging caused a drastical reduction of fracture toughness for both JN1 and JJ1 steels due to the intergranular precipitation of M23 C6 and Cr2 N. The presence of a more uniform and abundant intergranular precipitation resulted in a more rapid decrease in fracture toughness with aging time in JN1 steel due to its higher content of C and N, compared with JJ1 steel. The fractography of aged CVN impact specimens showed mainly intergranular brittle fracture. The volume fraction of precipitates can be uniquely correlated with reduction of toughness. Acknowledgements The authors wish to thank Dr. Hideo Nakajima of JAERI for providing materials used in the present work. References [1] Marshal P. Austenitic stainless steels microstructure and properties. London: Elsevier; 1984. [2] Nakajima H, Yoshida K, Shimamoto S. Development of new cryogenic steels for the superconducting magnets of the fusion experimental reactor. ISIJ Int 1990;30:567±78. [3] Suemune K, Sakamoto T, Ogawa T, Ozaki T, Maehara S. Manufacture and properties of nitrogen-containing Cr±Mn and Cr±Ni austenitic stainless steels for cryogenic use. Adv Cryog Eng 1988;34A:123±9. [4] Nakajima H, Nunoya Y, Nozawa M, Ivano O, Takano K, Ando T, Ohkita S. Development of high strength austenitic stainless
700
[5] [6]
[7] [8]
M.L. Saucedo-Mu~noz et al. / Cryogenics 40 (2000) 693±700 steel for conduit of Nb3 Al conductor. Adv Cryog Eng 1996;42A:323±30. Stoter LP. Thermal ageing eects in AISI type 316 stainless steel. J Mater Sci 1981;16:1039±51. Summers LT, Walsh RP, Miller JR. The cryogenic tensile and fracture toughness properties of austenitic steels given low temperatures short-time Nb3 Sn reaction treatments. Adv Cryog Eng 1996;42A:339±44. Simmons JW, Covino BS, Hawk JA, Dunning JS. Eect of nitride (Cr2 N) precipitation on the mechanical, corrosion and wear properties of austenitic stainless steels. ISIJ Int 1996;36:846±54. Muster WJ, Elster J. Low temperature embrittlement after ageing stainless steels. Cryogenics 1990;30:799±802.
[9] JIS Z-2242 and 2244. Method of impact test for metallic materials. JIS Handbook. Tokyo; 1997. [10] ASM metals handbook, vol. 8. Metals Park, Ohio: ASM; 1977. [11] Kikuchi M, Kajihara M, Choi S. Cellular precipitation involving both substitutional and interstitial solutes: cellular of Cr2 N in Cr± Ni austenitic steels. Mater Sci Eng A 1991;146:131±50. [12] Simmons JW, Atteridge JD, Rawers JS. Sensitization of highnitrogen austenitic stainless steels by dichromium nitride precipitation. Corrosion 1994;50:491±501. [13] Sedrinks AJ. Corrosion of stainless steels. New York: Interscience; 1996. [14] Porter DA, Easterling KE. Phase transformation in metals and alloys. London: Chapman & Hall; 1997.
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Eect of microstructure evolution on fracture toughness in isothermally aged austenitic stainless steels for cryogenic applications Maribel L. Saucedo-Mu~ noz a,b,*, Yutaka Watanabe c, Tetsuo Shoji a, Hideaki Takahashi a a Fracture Research Institute, Tohoku University, Aoba-Ku, Sendai 980, Japan Instituto Politecnico Nacional, ESIQIE, Apartado Postal 75-556, D.F. 07300, Mexico Department of Machine Intelligence and Systems, Graduate School of Engineering, Tohoku University, Aoba-Ku, Sendai 980, Japan b
c
Received 8 September 2000; accepted 15 December 2000
Abstract Two types of austenitic stainless steels JJ1 and JN1 were isothermally aged at temperatures from 600°C to 900°C for 10±1000 min in order to study the microstructural evolution and its eect on fracture toughness at cryogenic temperatures. These steels were developed for applications in the superconducting magnets of a fusion experimental reactor by the Japan Atomic Energy Research Institute. The Charpy V-Notch (CVN) fracture energy at 77 K showed a signi®cant decrease with aging time for both steels. The intergranular precipitation of carbides and nitrides is responsible for the fracture toughness deterioration. The scanning electron microscope (SEM) fractographs showed an intergranular brittle fracture and its fraction also increased with aging time and temperature. The presence of a more abundant intergranular precipitation resulted in a more rapid decrease in fracture toughness with aging time in JN1 steel due to its higher content of C and N, compared to that of JJ1 steel. The volume fraction of precipitates can be uniquely correlated with the reduction in toughness. Ó 2001 Elsevier Science Ltd. All rights reserved. Keywords: Austenitic stainless steels; Fracture toughness; Cryogenic; Microstructure
1. Introduction The 300 series austenitic stainless steels have been used for many years in the construction of components employed in chemical, petroleum, and nuclear power industries due to its corrosion-resistant and mechanical properties [1]. These types of steels also have wide application for cryogenic components, although they were not originally developed for cryogenic use. In fact, a number of cryogenic structures are constructed with 304L or 316L steels [1]. A high level combination of strength and toughness ry > 1200 MPa and JIC > 200 MPa m 1=2 at cryogenic temperatures is required for components for superconducting magnets of fusion reactors because the components need to sustain large electromagnetic forces. This requirement is beyond the capability of conventional steels. The Japan Atomic Energy Research Institute (JAERI) started the development of new cryogenic structural materials in 1982 [2±
*
Corresponding author. Fax: +52-5119-1986. E-mail address: [email protected] (M.L. SaucedoMu~ noz).
4]. The newly developed steels have been used as structural materials for the demo poloidal coils (DPS). Thick plates are used for coil supports and thin plates for the conduit materials and the cryogenic buer tank. Most of the components are exposed to thermal cycles during manufacturing due to the welding processes. Additionally, since superconductors such as Nb3 Al require a reaction heat treatment, the conduit material is also exposed to the heat treatment [3]. As is well known, these thermal processes can promote the intergranular and transgranular precipitation of carbides, nitrides and other phases. The precipitation process in austenitic stainless steel is extremely complex. For example, eighteen dierent precipitate phases were reported in an isothermally aged type 316 steel [5]. Furthermore, the same authors pointed out that the exposure of this type of steel to relatively high temperatures (about 650°C) resulted in a deterioration of the tensile properties, mainly a loss of ductility. Recently, Simmons et al. [7] studied the precipitation behavior in high-nitrogen austenitic steels at elevated temperatures. The authors also reported a drastic decrease of impact properties due to the precipitation of Cr2 N. These precipitates on grain boundaries were responsible for the intergranular
0011-2275/01/$ - see front matter Ó 2001 Elsevier Science Ltd. All rights reserved. PII: S 0 0 1 1 - 2 2 7 5 ( 0 1 ) 0 0 0 0 4 - 2
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M.L. Saucedo-Mu~noz et al. / Cryogenics 40 (2000) 693±700
fracture because of the reduction of cohesive strength of grain boundaries. This type of microstructure deterioration also decreases fracture toughness of steels [6±8]. In the present work, two types of austenitic stainless steels, JJ1 and JN1, developed for applications to the superconducting magnets of fusion experimental reactor by JAERI [2], were selected to study the microstructure evolution in isothermal aging and its eect on fracture toughness. Aging temperatures from 600°C to 900°C and short aging times from 10 to 1000 min were selected in order to cover the microstructural evolution, which can be compared to that caused either during the welding process or reaction heat-treatment.
ysis of precipitates was also conducted using the extraction replica technique [5]. 3. Results and discussion 3.1. Toughness reduction due to thermal aging Figs. 1(a)±(c) show the plots of CVN fracture energy at 77 K versus aging time for the JN1 and JJ1 steels aged at 700°C, 800°C and 900°C, respectively. Both steels
2. Materials and experimental procedure Materials used in this work were forged-steel plates of 200 mm thick and their chemical compositions are shown in Table 1. Table 2 shows the mechanical properties of this type of steel at 4 K. The solution treatment of JN1 and JJ1 was carried out at 1075°C and 1050°C, respectively, for 1 h under an argon atmosphere, and then water-quenched. The aging temperatures and times were 600°C, 700°C, 800°C and 900°C and from 10 to 1000 min, respectively. The Charpy V-Notch (CVN) test was conducted at 196°C (77 K), following JIS Z2242 standard [9]. The fracture surface of the CVN-tested specimens and microstructure of the aged specimens were observed with a scanning electron microscope (SEM). The precipitates in the aged samples were extracted electrolitically by dissolution of austenitic matrix in a solution of 10 vol.%HCl±CH3 OH at 4 V. The X-ray diraction pattern of extracted precipitates was measured in a SIEMENS diractometer using KaCu radiation. The TEM microanalysis and observation of precipitates were carried out using a JEOL FX2000 microscope at 200 kV, equipped with energy dispersion X-ray analysis (EDX) facility. The TEM specimens were prepared by electropolishing in a 5 vol.% perchloric acid±20 vol.% glycerin±75 vol.% ethanol electrolyte at 25 V at room temperature. The SEM/EDX microanal-
Fig. 1. CVN fracture energy at 77 K as a function of aging time for JN1 and JJ1 steels aged at (a) 700°C, (b) 800°C and (c) 900°C.
Table 1 Chemical composition of materials (wt%) Material
C
Si
Mn
P
S
Ni
Cr
Al
N
Mo
JN1 JJ1
0.040 0.025
0.97 0.48
3.88 10.13
0.022 0.021
0.001 0.002
15.07 11.79
24.32 12.01
0.023 ±
0.32 0.236
± 4.94
Table 2 Mechanical properties at 4 K of solution treated JN1 and JJ1 steels Material
ry (MPa)
ruts (MPa)
Elongation (%)
JIC kJ m 2
JN1 JJ1
1363 1169
1752 1591
31 39
346 383
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695
showed a monotonotic decrease in the CVN fracture energy with aging time at all the three temperatures. It is also evident that the drop of fracture toughness of JN1 steel is always faster than that of JJ1 steel. The fastest drop of fracture toughness occurred in the JN1 steel samples aged at 900°C. This fact may be attributed to the higher content of C and N in JN1 steel, which can lead to faster kinetics in intergranular precipitation during the aging process, as discussed in a later section. The decrease in CVN fracture energy in aged samples is presented as a function of Larson±Miller parameter (LMP) for JN1 and JJ1 steels in Figs. 2(a) and (b), respectively. LMP is de®ned by the following equation: LMP T C log t;
Fig. 2. CVN fracture energy at 77 K as a function of LMP for (a) JN1 and (b) JJ1 steels.
1
where T is the temperature in K, t is the time in hours and C is a constant, which depends on the material. For these steels, the constant C that best ®ts the experimental data were 8 and 5 for JN and JJ steels, respectively. These plots enable us to predict the fracture toughness for any given aging condition. The typical fracture appearances of CVN test specimens of JN1 are shown in Figs. 3(a)±(i) for the solution
Fig. 3. SEM fractographs of JN1 steel (a) solution treated and aged for (b) 10, (c) 100 and (d) 1000 min at 700°C, for (e) 10, (f) 100 and (g) 1000 min at 800°C and for (h) 10, (i) 100 and (j) 1000 min at 900°C.
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Fig. 4. SEM fractographs of JJ1 steel (a) solution treated and aged for (b) 10, (c) 100 and (d) 1000 min at 700°C, for (e) 10, (f) 100 and (g) 1000 min at 800°C and for (h) 10, (i) 100 and (j) 1000 min at 900°C.
treated condition and the various aged conditions. The fracture morphology of JJ1 is shown in Figs. 4(a)±(i) in the same manner. Both of JN1 and JJ1 steels fractured in a ductile manner in the solution treated condition. Intergranular facets were found from all the aged samples, although area fraction of intergranular facets to ductile surface was strongly dependent on aging conditions. The fraction of intergranular brittle fracture increased with aging time and temperature, and seemed consistent with the CVN fracture energy value. Simmons et al. [7] found that the fracture energy decreased with aging time and temperature in a highnitrogen austenitic stainless steel, nominally Fe±18Cr± 5Mn±5Ni±3Mo±0.02C±0.7N. The Cr2 N intergranular precipitation was the main cause for fracture toughness deterioration in this particular steel. 3.2. Microstructural evolution during thermal aging The typical microstructures of JN1 steel aged at 700°C, 800°C and 900°C for 10, 100 and 1000 min are shown in Figs. 5(a)±(i). The samples aged at 700°C for up to 1000 min have precipitates only at grain boundaries, Figs. 5(a)±(c). At 800°C and 900°C, precipitates ®rst formed at grain boundaries, Figs. 5(d), (e) and (g),
and then followed by intragranular precipitation, Figs. 5(f), (h) and (i). Both the precipitations at the grain boundary and in the matrix were more pronounced for longer aging times. The intragranular precipitates can be classi®ed into two types: cellular or discontinuous precipitation and plate-like precipitates, which have a preferred alignment with the austenitic matrix. The morphology of cellular precipitates is similar to that of pearlite in carbon-steels, Fig. 5(i). The formation of this lamellar microstructure initiated at grain boundaries and grew into the austenite matrix, according to the following reaction [7]: c ! c Cr2 N:
2
The volume fraction of the discontinuous precipitation increased with aging time and the maximum value was determined by the point-count grid method [10], to be about 0.04. This value seems to be reasonable, since a volume fraction of 0.1 was reported in an austenitic stainless steel containing 0.42% N, after a long aging [11]. Fig. 6(a) shows the X-ray diraction pattern of extracted precipitates for the JN1 steel aged at 900°C for 300 and 1000 min. It is evident that the precipitates are mostly the Cr2 N and M23 C6 phases for these aging
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Fig. 5. SEM micrographs of JN1 steels aged for (a) 10, (b) 100 and (c) 1000 min at 700°C; for (d) 10, (e) 100 and (f) 1000 min at 800°C and for (g) 10, (h) 100 and (i) 1000 min at 900°C.
conditions. The types of precipitates were found to be the same also for the samples aged at 700°C and 800°C. At 900°C, the relative intensity of XRD peaks for Cr2 N phase increased with respect to those for M23 C6 with aging time. This fact suggests that the Cr2 N precipitation is favored at 900°C and the M23 C6 precipitation is facilitated at temperatures between 700°C and 800°C because the relative intensity of M23 C6 is higher than that of Cr2 N for these aging temperatures. Other works [12,13] have reported that the maximum rates of Cr2 N and M23 C6 precipitation are achieved at about 900°C and 800°C, respectively, in high-nitrogen austenitic stainless steels. The microstructures of JJ1 steel aged at 700°C, 800°C and 900°C for 100, 500 and 1000 min are shown in Figs. 7(a)±(i). The grain boundary precipitation again preceded the intragranular one. The aging treatments for up to 1000 min at any aging temperature indicated no cellular precipitation in this steel. It is also evident that there is a more abundant precipitation in JJ1 steel than that in JN1 steel for the samples aged at 800°C and 900°C for 1000 min. This fact seems to be inconsistent with the expected precipitation behavior in JN1 steel, based on the content of interstitial solutes. That is, JN1 steel has N and C contents higher than those of JJ1 steel
and thus the volume fraction of carbides or nitrides can be higher in the former case [14]. However, as recognized in the XRD results, the precipitation of g Fe2 Mo phase took place in the late stages of aging process for JJ1 steel, promoting a higher volume fraction for this aging condition. It is also important to notice that the intergranular precipitation is more uniform and abundant in JN1 steel, as expected due to its higher interstitial solute content. Fig. 6(b) shows XRD patterns of aged JJ1 steel samples aged at 900°C for 300 and 1000 min. The precipitates consist of Cr2 N; M23 C6 and g Fe2 Mo phases at 800°C and 900°C and are mainly composed of Cr2 N and M23 C6 phases at 700°C. The precipitation of g phase was detected in samples aged at 800°C for 300 min or longer and at 900°C for 100 min or longer. The intensities of XRD peaks for the Cr2 N and g phases are higher than that of the M23 C6 phase in JJ1 samples aged at 800°C and 900°C in contrast with JN1 steel. This suggests that precipitation of Cr2 N and g phases is predominant in JJ1 steel at these temperatures. The formation of g phase has been reported to occur only in austenitic stainless steels with Mo. For example, the 316type steel aged at temperatures between 250°C and 900°C for times longer than 6000 min [13].
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(a)
(b) Fig. 6. XRD patterns of (a) JN1 and (b) JJ1 steel aged at 900°C for 300 and 1000 min.
Fig. 7. SEM micrographs of JJ1 steels aged for (a) 100, (b) 500 and (c) 1000 min at 700°C; for (d) 100, (e) 500 and (f) 1000 min at 800°C for (g) 100, (h) 500 and (i) 1000 min at 900°C.
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(a)
Fig. 9. CVN fracture energy at 77 K versus volume fraction of precipitates for JN1 and JJ1 steels.
cipitation of 0.03. That means that the intergranular precipitation is the main cause of fracture toughness deterioration during aging of both steels. This fact is also con®rmed by the presence of intergranular brittle fracture and its increase with aging time. (b) Fig. 8. TTP diagrams of (a) JN1 and (b) JJ1 steels.
The SEM and TEM microanalyses of the precipitates showed that the intergranular M23 C6 precipitates are not only composed of Cr, but also of Fe in JN1 steel and of Fe and Mo in JJ1 steels. Furthermore, Mo is the most enriched element in M23 C6 precipitates of the JJ1 steel. The Cr2 N phase of cellular precipitation was mainly composed of Cr. The intragranular Cr2 N and g phases consisted of Cr and Fe and of Fe and Mo, respectively. All the above results are summarized in the time± temperature±precipitation (TTP) diagrams of JN1 and JJ1 steels, as shown in Figs. 8(a) and (b). In general, it can be noticed that the kinetics of precipitation for JN1 steel are faster than that of JJ1 steel, because of its higher interstitial solute content [14]. The TTP diagrams show that the intergranular precipitation of M23 C6 and M2 N preceded the intragranular precipitation of M2 N and M2 N and g phase in JN1 and JJ1 steels, respectively. 3.3. Correlation between embrittlement and precipitation Fig. 9 shows the plot of CVN fracture energy at 77 K as a function of the total volume fraction of precipitation for JN1 and JJ1 steels. This ®gure shows clearly that fracture energy decreased with increasing precipitate volume fraction. It is also noticed that a volume fraction of precipitation as small as about 0.03 caused a drastical decrease in the fracture energy for JN1 and JJ1 steels. It is important to mention that the intergranular precipitation covers almost all grain boundaries for the case of a sample with a total volume fraction of pre-
4. Conclusions The thermal aging caused a drastical reduction of fracture toughness for both JN1 and JJ1 steels due to the intergranular precipitation of M23 C6 and Cr2 N. The presence of a more uniform and abundant intergranular precipitation resulted in a more rapid decrease in fracture toughness with aging time in JN1 steel due to its higher content of C and N, compared with JJ1 steel. The fractography of aged CVN impact specimens showed mainly intergranular brittle fracture. The volume fraction of precipitates can be uniquely correlated with reduction of toughness. Acknowledgements The authors wish to thank Dr. Hideo Nakajima of JAERI for providing materials used in the present work. References [1] Marshal P. Austenitic stainless steels microstructure and properties. London: Elsevier; 1984. [2] Nakajima H, Yoshida K, Shimamoto S. Development of new cryogenic steels for the superconducting magnets of the fusion experimental reactor. ISIJ Int 1990;30:567±78. [3] Suemune K, Sakamoto T, Ogawa T, Ozaki T, Maehara S. Manufacture and properties of nitrogen-containing Cr±Mn and Cr±Ni austenitic stainless steels for cryogenic use. Adv Cryog Eng 1988;34A:123±9. [4] Nakajima H, Nunoya Y, Nozawa M, Ivano O, Takano K, Ando T, Ohkita S. Development of high strength austenitic stainless
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M.L. Saucedo-Mu~noz et al. / Cryogenics 40 (2000) 693±700 steel for conduit of Nb3 Al conductor. Adv Cryog Eng 1996;42A:323±30. Stoter LP. Thermal ageing eects in AISI type 316 stainless steel. J Mater Sci 1981;16:1039±51. Summers LT, Walsh RP, Miller JR. The cryogenic tensile and fracture toughness properties of austenitic steels given low temperatures short-time Nb3 Sn reaction treatments. Adv Cryog Eng 1996;42A:339±44. Simmons JW, Covino BS, Hawk JA, Dunning JS. Eect of nitride (Cr2 N) precipitation on the mechanical, corrosion and wear properties of austenitic stainless steels. ISIJ Int 1996;36:846±54. Muster WJ, Elster J. Low temperature embrittlement after ageing stainless steels. Cryogenics 1990;30:799±802.
[9] JIS Z-2242 and 2244. Method of impact test for metallic materials. JIS Handbook. Tokyo; 1997. [10] ASM metals handbook, vol. 8. Metals Park, Ohio: ASM; 1977. [11] Kikuchi M, Kajihara M, Choi S. Cellular precipitation involving both substitutional and interstitial solutes: cellular of Cr2 N in Cr± Ni austenitic steels. Mater Sci Eng A 1991;146:131±50. [12] Simmons JW, Atteridge JD, Rawers JS. Sensitization of highnitrogen austenitic stainless steels by dichromium nitride precipitation. Corrosion 1994;50:491±501. [13] Sedrinks AJ. Corrosion of stainless steels. New York: Interscience; 1996. [14] Porter DA, Easterling KE. Phase transformation in metals and alloys. London: Chapman & Hall; 1997.