- Email: [email protected]
Creep deformation of Alloys 617 and 276 at 750–950 °C
Materials Science and Engineering A 520 (2009) 184–188
Contents lists available at ScienceDirect
Materials Science and Engineering A journal homepage: www.elsevier.com/locate/msea
Creep deformation of Alloys 617 and 276 at 750–950 ◦ C Ajit K. Roy ∗ , Muhammad H. Hasan, Joydeep Pal Department of Mechanical Engineering, University of Nevada, Las Vegas (UNLV), 4505 Maryland Parkway, Box 454027, Las Vegas, NV 89154-4027, United States
a r t i c l e
i n f o
Article history: Received 4 December 2008 Received in revised form 13 May 2009 Accepted 14 May 2009 Keywords: Alloy 617 Alloy 276 Temperature and stress effect Creep mechanism
a b s t r a c t Alloys 617 and 276 were subjected to time-dependent deformation at elevated temperatures under sustained loading of different magnitudes. The results indicate that Alloy 617 did not exhibit strains exceeding 1 percent (%) in 1000 h at 750, 850 and 950 ◦ C when loaded to 10% of its yield strength (YS) values at these temperatures. However, this alloy was not capable of sustaining higher stresses (0.25YS and 0.35YS) for 1000 h at 850 and 950 ◦ C without excessive deformation. Interestingly, Alloy 617 showed insignificant steady-state creep rate at 750 ◦ C irrespective of the applied stress levels. Alloy 276 almost met the maximum creep deformation criterion when tested at 51 MPa–750 ◦ C. Severe creep deformation of both alloys at 950 ◦ C could be attributed to the dissolution of carbides and intermetallic phases remaining after solution annealing or precipitated during quenching. © 2009 Elsevier B.V. All rights reserved.
1. Introduction Nickel (Ni)-base FCC Alloy 617 has recently been identified [1,2] to be one prime candidate structural material for electricity generation within the purview of the next generation nuclear plant (NGNP) program. The NGNP program is focused on the development of modular high-temperature gas-cooled reactors using helium as a coolant, and a closed-cycle gas turbine for power generation. A maximum operating temperature of 950 ◦ C was previously recommended by the materials advisory board to enhance the efficiency in power generation using nuclear heat. More recently, a reduced reactor core outlet/turbine inlet temperature of 750–800 ◦ C has been suggested to achieve this goal. The selection of Alloy 617 was based on its excellent mechanical properties, oxidation and creep-resistance, and phase stability at elevated temperatures [3]. Alloy 617, which has been proposed to be the primary structural material for use in the intermediate heat exchanger, would possibly undergo time-dependent plastic deformation (creep) at elevated temperatures. Since thermal stresses of different magnitudes will be generated during such operation, it is necessary to determine the susceptibility of this material to creep deformation at applied loads corresponding to certain percentages of its yield strength (YS) values at different temperatures. Conventionally, creep deformation of structural materials is more pronounced at temperatures at or above 50 percent (%) of their melting temperatures (Tm ). Therefore, this investigation was
∗ Corresponding author. Tel.: +1 803 725 3958; fax: +1 803 725 7369. E-mail address: [email protected] (A.K. Roy). 0921-5093/$ – see front matter © 2009 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2009.05.029
performed to evaluate the creep behavior of Alloy 617 at 750, 850 and 950 ◦ C under sustained loading. Efforts have also been made to study the effect of applied load on the extent of deformation at a constant temperature for an identical testing period. Alloy 276, which has been extensively studied [4] at the authors’ laboratories for evaluation of its high-temperature tensile properties, was also included in this investigation to compare its creep behavior to that of Alloy 617. This paper presents the comprehensive test results on both alloys, providing a basic understanding of creep deformation as a function of applied stress at temperatures relevant to the NGNP program. 2. Materials and experimental procedure A vacuum-induction-melting (VIM) practice was used to develop an experimental heat of Alloy 617. The VIM heat was subsequently processed into round bars by forging and hot rolling. The hot rolled bars were then subjected to a thermal treatment that consisted of solution annealing at 1180 ◦ C followed by an oil-quench. On the other hand, Alloy 276 was received in a heat-treated condition in the form of round bars. They were solution annealed at 1163 ◦ C followed by a water quench. The chemical compositions of both alloys are given in Table 1. The test specimens were fabricated from the heat-treated bars of both alloys in such a way that the gage section was parallel to the longitudinal rolling direction. The size requirements prescribed by the ASTM Designation E 139-2000 [5] were followed to machine these specimens. They had an overall length of 101.6mm with a gage length of 37.59-mm, as shown in Fig. 1. Circular grooves were machined at both ends beyond the shoulder region of these specimens to attach dual extensometers for monitoring
A.K. Roy et al. / Materials Science and Engineering A 520 (2009) 184–188
185
Table 1 Chemical compositions of tested materials (wt.%). Material/heat no.
C
Mn
Fe
S
Si
Cu
Cr
Alloy 617/HV1160 Alloy 276/Z7437CG
0.06 0.006
0.12 0.42
0.002 5.94
0.009 0.001
0.004 0.008
0.001 22.10 – 15.84
Ni
Al
Ti
Co
Mo
Ta
P
54.80 58.33
0.87 –
0.29 –
12.17 0.10
9.52 15.93
0.001 0.001 – 0.002
V
W
– 0.01
– 3.40
elongation during creep testing. All deformation measured by the extensometers was assumed to occur within the gage section. The loading frames used in creep testing had a lever arm ratio of 20:1. The overall creep rates were determined from elongations measured by the two extensometers placed at the circular grooves. K-type thermocouples were used to monitor temperatures at two locations within the gage section, away from its center. WINCCS (Windows Computer Creep System) version 2 software was used for automatic data acquisition. A split furnace was used to heat the specimens of Alloy 617 in air to 750, 850 and 950 ◦ C under applied stresses of 22, 54 and 78 MPa (750 ◦ C); 24, 59 and 83 MPa (850 ◦ C); and 18, 46 and 64 MPa (950 ◦ C). These stresses represented 10%, 25% and 35% of the YS values of Alloy 617 at these three temperatures. The magnitude of YS at these temperatures was determined from the engineering stress versus engineering strain diagrams generated from tensile testing performed according to the ASTM Designation E 8-04 [6]. The higher YS value of Alloy 617 at 850 ◦ C can be attributed to the occurrence of the yield strength anomaly phenomenon that has recently been reported elsewhere [7]. Alloy 276 was also tested in air at 750, 850 and 950 ◦ C under applied stresses of 51, 49 and 39 MPa, respectively for evaluation of its creep behavior. 3. Results and discussion The optical micrographs of Alloy 617 loaded to 10% of its YS (0.10YS) values at 750, 850 and 950 ◦ C are illustrated in Fig. 2a–c, respectively. Large austenitic grains, annealing twins, and precipitates were observed in all three micrographs. A recent study [8] performed on creep deformation of Alloy 617 suggests that these precipitates could consist of chromium carbide (Cr23 C6 ) and/or intermetallic phase of chromium and molybdenum (Cr–Mo). Austenitic grains, annealing twins and small precipitates were also observed in optical micrographs of Alloy 276, as shown in Fig. 3. The average grain diameters of both alloys were measured by mean linear intercept method [9]. The grain diameters of Alloys 617 and 276 in the as-machined condition were 0.097 and 0.099 mm, respectively. For Alloy 617, the average grain diameter varied between 0.098 and 0.126 mm when tested within a temperature range of 750–950 ◦ C under different applied stress levels. Similarly, grain diameter ranging from 0.101 to 0.127 mm was determined for Alloy 276 within the same temperature range. Considering grain sizes of both alloys in the as-machined condition, and at temperatures of 750, 850 and 950 ◦ C, their standard deviations were ±0.013 and 0.011 mm for Alloys 617 and 276, respectively. Together, the grain size data suggest that there is a tendency for the grain size of these alloys to slightly increase as the temperature/exposure
Fig. 1. Dimensions of creep testing specimen.
Fig. 2. Optical micrographs of Alloy 617, Kalling’s reagent (a) 750 ◦ C, 22 MPa, 1000 h; (b) 850 ◦ C, 24 MPa, 1000 h; (c) 950 ◦ C, 18 MPa, 216 h.
186
A.K. Roy et al. / Materials Science and Engineering A 520 (2009) 184–188
Fig. 3. Optical micrograph of Alloy 276 (51 MPa, 750 ◦ C, 1000 h), HCl + HNO3 + CuCl2 .
time increases, ranging from ASTM grain sizes within 4–3 for asmachined alloy to ASTM grain size of 3 after 950 ◦ C exposure. The results of creep testing involving Alloy 617 are shown in Fig. 4 in the form of % creep versus time as a function of temperature at applied stresses equivalent to 0.10YS (22, 24 and 18 MPa), 0.25YS (54, 59 and 46 MPa) and 0.35YS (78, 83 and 64 MPa). It is interesting to note that the magnitude of instantaneous elastic plus plastic strain resulting from the initial applied stress was relatively higher at elevated temperatures. The modulus of elasticity is known [10] to decrease with increasing temperature, which could possibly account for the enhanced anelastic strain at higher temperatures. Further, the primary creep curve was relatively shorter at higher initial applied stresses, irrespective of the testing temperature. At 18 MPa–950 ◦ C (Fig. 4(a)), this alloy exhibited a very short steadystate region, followed by an extended third stage. On the contrary, substantially longer secondary creep regions were observed in this alloy at 750 and 850 ◦ C under initial applied stresses of 22 and 24 MPa, respectively (Fig. 4(a)). With increasing applied stress levels (0.25YS and 0.35YS) and temperature (Fig. 4(b, c)), the steady-state region became shorter and finally disappeared at 950 ◦ C, showing only a steeper tertiary creep curve. Assuming that a structural material must not undergo creep deformation exceeding 1% strain following 1000 h of loading at different stress levels, it could be stated that Alloy 617 may not be capable of withstanding an operating temperature of 950 ◦ C at applied stresses above 10% of its YS value. Data shown in Fig. 4(b) indicate that Alloy 617 may be able to sustain an operating temperature of 850 ◦ C for periods of up to 850–900 h when loaded to an applied stress of 59 MPa. Further, this alloy reached a tertiary stage almost immediately when loaded to a higher stress level of 83 MPa at a similar temperature, as shown in Fig. 4(c). Thus, the inference is strong that Alloy 617 may not be suitable for application under sustained loading for 1000 h both at 0.25YS and 0.35YS stress levels at 850 ◦ C or higher. It is, however, interesting to note that this alloy was capable to sustain all three levels of stress at 750 ◦ C by virtue of its prolonged and stable steady-state creep rates even beyond 1000 h of testing (Fig. 4). Thus, based on the overall data, it may be concluded that Alloy 617 could be safely used as a heat exchanger material within an operating temperature of 750 ◦ C at applied stresses corresponding to its 0.10YS, 0.25YS and 0.35YS values. Perhaps, this alloy may also be considered for a similar application at 850 ◦ C at applied stresses not exceeding its 0.25YS value. The results of creep testing involving Alloy 276 under sustained loads corresponding to its 0.25YS values (51, 49 and 39 MPa) at three temperatures are illustrated in Fig. 5. These data indicate that this
Fig. 4. Creep curves of Alloy 617 vs. temperature and applied stress. (a) Applied stress = 0.10YS. (b) Applied stress = 0.25YS. (c) Applied stress = 0.35YS.
A.K. Roy et al. / Materials Science and Engineering A 520 (2009) 184–188
Fig. 5. Creep curves of Alloy 276 vs. temperature and applied stress.
alloy did not experience any significant steady-state creep at 950 ◦ C. Rather, it suffered from extensive deformation within the initial 200 h of testing, showing a very steep slope in the tertiary stage. With respect to its creep deformation at 850 ◦ C, this alloy exhibited strain exceeding 1% following 1000 h of loading at 49 MPa. However, at 750 ◦ C, this alloy was very close to meeting the maximum strain criterion of 1% under an initial applied stress of 51 MPa for a loading period of 1000 h. A comparison of steady-state creep rates of Alloys 617 and 276, subjected to applied stresses corresponding to their individual 0.25YS values at 750, 850 and 950 ◦ C, is illustrated in Fig. 6, clearly showing a superior creep-resistance of Alloy 617. Also, a variation of steady-state creep rate with applied stress for Alloy 617 at three tested temperatures is shown in Fig. 7. Literature data [8] on Alloy 617 have also been superimposed in this figure for comparison of its creep rates at similar temperatures. It is obvious that the creep rates obtained from this study were significantly higher compared to those cited in the literature, possibly due to the differences in the tested environmental conditions (air versus He + O2 ). The resultant data also indicate that, with the exception of 850 and 950 ◦ C, the normal temperature-compensated power law did not apply to this
Fig. 6. log10 (creep rate) vs. 1/T and applied stress.
187
Fig. 7. log10 (creep rate) vs. log10 (applied stress).
alloy, possibly due to the changes in microstructure and anomalous yield strength behavior at elevated temperatures. Alloys 617 and 276 are both solid-solution strengthened alloys. It has been reported [11,12] that carbides of M23 C6 and M6 C types, and Cr–Mo intermetallic phase can be precipitated at their grain boundaries and as intragranular particles during solution annealing at elevated temperatures. While both types of carbides can co-exist at the grain boundaries, M23 C6 is Cr-rich and M6 C has higher Mo content. However, a preferential precipitation of Cr-rich M23 C6 type carbide (Cr23 C6 ) appears to be more reasonable than Mo-rich M6 C carbide (Mo6 C) due to a greater diffusion coefficient of Cr than Mo in Ni-base alloys. The formation of precipitates in the matrix can also lead to the development of subgrains. The formation of subgrains and grain boundary precipitation can inhibit dislocation motion [7,8,13], thus preventing accelerated deformation rates of these alloys at 750 and 850 ◦ C under relatively lower applied stress levels (0.10YS and 0.25YS), leading to the occurrence of prolonged steady-state region in the creep curves. Fig. 8 illustrates dislocation pile-ups at the grain boundaries, and subgrain formation within the matrix of Alloy 617 tested under an applied stress of 59 MPa at 850 ◦ C. At 950 ◦ C, the unstable intragranular carbides and grain boundary carbides may undergo dissolution, subsequently causing migration of carbides and grain boundaries that could lead to the initiation of voids at the carbide phases existing on the grain boundaries. Finally, recrystallization occurs as creep deformation progresses with eventual breaking up of grains [14].
Fig. 8. TEM micrograph of Alloy 617 (59 MPa, 850 ◦ C, 1000 h).
188
A.K. Roy et al. / Materials Science and Engineering A 520 (2009) 184–188
4. Summary and conclusions Alloys 617 and 276 have been tested under sustained loading for evaluation of their creep deformation behavior at 750, 850 and 950 ◦ C. Alloy 617 has been loaded at stress levels corresponding to 10%, 25% and 35% of its YS values at these temperatures. However, creep deformation of Alloy 276 has been investigated only at applied stresses equivalent to 25% of its YS values at similar temperatures. The significant results and conclusions drawn from this study are given below: • The primary stage of creep was very short for both alloys irrespective of the testing temperature and the applied stress. • Both alloys exhibited prolonged and slower steady-state creep rate at 850 ◦ C when loaded to 25% of their YS values for a comparable duration. While Alloy 617 was very close to meeting the maximum strain criterion of 1% in 1000 h of loading, Alloy 276 suffered from excessive deformation. • Severe creep deformation, characterized by the formation of an instantaneous tertiary region, was observed with Alloy 617 when testing was performed at 850 and 950 ◦ C under applied stresses corresponding to 35% of its YS values (83 and 64 MPa) at these temperatures. However, this alloy met the deformation acceptance criterion at 950 ◦ C when loaded to its 0.10YS value. • Alloy 617 was capable of sustaining all three levels of applied stress (22, 54 and 78 MPa) for durations exceeding 1000 h at 750 ◦ C. Considering a maximum allowable strain of 1% following 1000 h of sustained loading, this alloy may be most suitable for use as a heat exchanger material at operating temperatures up to 750 ◦ C under applied stresses equivalent to its 0.10YS, 0.25YS and 0.35YS values. • Alloy 276 suffered from severe creep deformation at 950 ◦ C at an applied stress of 39 MPa (0.25YS), showing a steep tertiary curve.
However, this alloy was very close to meeting the maximum strain criterion at 750 ◦ C under an applied stress of 51 MPa (0.25YS). • Dissolution of carbides and intermetallic phases formed and precipitated during thermal treatment that can promote hardening of the tested materials, could possibly account for void formation leading to recrystallization and subsequent break-up of grains at 950 ◦ C. Acknowledgment This work was funded by the United States Department of Energy under grant number DE-FC07-04ID14566. References [1] W. Ren, R. Swindeman, ASME Pressure Vessels and Piping Division Conference, Texas, 2007. [2] B.S. Rao, H.P. Meurer, H. Schuster, Mater. Sci. Eng. A 104 (1988) 37–51. [3] W.L. Mankins, J.C. Hosier, T.H. Bassford, Metall. Mater. Trans. B 5 (1974) 2579–2590. [4] A.K. Roy, J. Pal, C. Mukhopadhyay, Mater. Sci. Eng. A 474 (2008) 363–370. [5] ASTM Designation E 139-2000, Creep Rupture and Stress-Rupture Tests of Metallic Materials, American Society for Testing and Materials (ASTM) International. [6] ASTM Designation E 8-2004, Standard Test Methods for Tensile Testing of Metallic Materials, American Society for Testing and Materials (ASTM) International. [7] A.K. Roy, V. Marthandam, Mater. Sci. Eng. A, (2009), doi:10.1016/j.mesa.2009. 03.090. [8] P.S. Shankar, K. Natesan, J. Nucl. Mater. 366 (2007) 26–36. [9] ASTM E 112-1996, Standard Test Methods for Determining Average Grain Size, American Society for Testing and Materials (ASTM) International. [10] A.K. Roy, J. Pal, M.H. Hasan, J. Eng. Mater. Technol., in press. [11] M. Raghavan, B.J. Berkowitz, J.C. Scanlon, Metall. Mater. Trans. A 13 (1982) 979–984. [12] H.M. Tawancy, J. Mater. Sci. 31 (1996) 3929–3936. [13] V. Marthandam, Tensile deformation, toughness and crack propagation studies of Alloy 617, Ph.D. Dissertation, Department of Mechanical Engineering, University of Nevada, Las Vegas, April 2008. [14] S. Kihara, J.B. Newkirk, A. Ohtomo, Y. Saiga, Metall. Mater. Trans. A 11A (1980) 1019–1031.
Contents lists available at ScienceDirect
Materials Science and Engineering A journal homepage: www.elsevier.com/locate/msea
Creep deformation of Alloys 617 and 276 at 750–950 ◦ C Ajit K. Roy ∗ , Muhammad H. Hasan, Joydeep Pal Department of Mechanical Engineering, University of Nevada, Las Vegas (UNLV), 4505 Maryland Parkway, Box 454027, Las Vegas, NV 89154-4027, United States
a r t i c l e
i n f o
Article history: Received 4 December 2008 Received in revised form 13 May 2009 Accepted 14 May 2009 Keywords: Alloy 617 Alloy 276 Temperature and stress effect Creep mechanism
a b s t r a c t Alloys 617 and 276 were subjected to time-dependent deformation at elevated temperatures under sustained loading of different magnitudes. The results indicate that Alloy 617 did not exhibit strains exceeding 1 percent (%) in 1000 h at 750, 850 and 950 ◦ C when loaded to 10% of its yield strength (YS) values at these temperatures. However, this alloy was not capable of sustaining higher stresses (0.25YS and 0.35YS) for 1000 h at 850 and 950 ◦ C without excessive deformation. Interestingly, Alloy 617 showed insignificant steady-state creep rate at 750 ◦ C irrespective of the applied stress levels. Alloy 276 almost met the maximum creep deformation criterion when tested at 51 MPa–750 ◦ C. Severe creep deformation of both alloys at 950 ◦ C could be attributed to the dissolution of carbides and intermetallic phases remaining after solution annealing or precipitated during quenching. © 2009 Elsevier B.V. All rights reserved.
1. Introduction Nickel (Ni)-base FCC Alloy 617 has recently been identified [1,2] to be one prime candidate structural material for electricity generation within the purview of the next generation nuclear plant (NGNP) program. The NGNP program is focused on the development of modular high-temperature gas-cooled reactors using helium as a coolant, and a closed-cycle gas turbine for power generation. A maximum operating temperature of 950 ◦ C was previously recommended by the materials advisory board to enhance the efficiency in power generation using nuclear heat. More recently, a reduced reactor core outlet/turbine inlet temperature of 750–800 ◦ C has been suggested to achieve this goal. The selection of Alloy 617 was based on its excellent mechanical properties, oxidation and creep-resistance, and phase stability at elevated temperatures [3]. Alloy 617, which has been proposed to be the primary structural material for use in the intermediate heat exchanger, would possibly undergo time-dependent plastic deformation (creep) at elevated temperatures. Since thermal stresses of different magnitudes will be generated during such operation, it is necessary to determine the susceptibility of this material to creep deformation at applied loads corresponding to certain percentages of its yield strength (YS) values at different temperatures. Conventionally, creep deformation of structural materials is more pronounced at temperatures at or above 50 percent (%) of their melting temperatures (Tm ). Therefore, this investigation was
∗ Corresponding author. Tel.: +1 803 725 3958; fax: +1 803 725 7369. E-mail address: [email protected] (A.K. Roy). 0921-5093/$ – see front matter © 2009 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2009.05.029
performed to evaluate the creep behavior of Alloy 617 at 750, 850 and 950 ◦ C under sustained loading. Efforts have also been made to study the effect of applied load on the extent of deformation at a constant temperature for an identical testing period. Alloy 276, which has been extensively studied [4] at the authors’ laboratories for evaluation of its high-temperature tensile properties, was also included in this investigation to compare its creep behavior to that of Alloy 617. This paper presents the comprehensive test results on both alloys, providing a basic understanding of creep deformation as a function of applied stress at temperatures relevant to the NGNP program. 2. Materials and experimental procedure A vacuum-induction-melting (VIM) practice was used to develop an experimental heat of Alloy 617. The VIM heat was subsequently processed into round bars by forging and hot rolling. The hot rolled bars were then subjected to a thermal treatment that consisted of solution annealing at 1180 ◦ C followed by an oil-quench. On the other hand, Alloy 276 was received in a heat-treated condition in the form of round bars. They were solution annealed at 1163 ◦ C followed by a water quench. The chemical compositions of both alloys are given in Table 1. The test specimens were fabricated from the heat-treated bars of both alloys in such a way that the gage section was parallel to the longitudinal rolling direction. The size requirements prescribed by the ASTM Designation E 139-2000 [5] were followed to machine these specimens. They had an overall length of 101.6mm with a gage length of 37.59-mm, as shown in Fig. 1. Circular grooves were machined at both ends beyond the shoulder region of these specimens to attach dual extensometers for monitoring
A.K. Roy et al. / Materials Science and Engineering A 520 (2009) 184–188
185
Table 1 Chemical compositions of tested materials (wt.%). Material/heat no.
C
Mn
Fe
S
Si
Cu
Cr
Alloy 617/HV1160 Alloy 276/Z7437CG
0.06 0.006
0.12 0.42
0.002 5.94
0.009 0.001
0.004 0.008
0.001 22.10 – 15.84
Ni
Al
Ti
Co
Mo
Ta
P
54.80 58.33
0.87 –
0.29 –
12.17 0.10
9.52 15.93
0.001 0.001 – 0.002
V
W
– 0.01
– 3.40
elongation during creep testing. All deformation measured by the extensometers was assumed to occur within the gage section. The loading frames used in creep testing had a lever arm ratio of 20:1. The overall creep rates were determined from elongations measured by the two extensometers placed at the circular grooves. K-type thermocouples were used to monitor temperatures at two locations within the gage section, away from its center. WINCCS (Windows Computer Creep System) version 2 software was used for automatic data acquisition. A split furnace was used to heat the specimens of Alloy 617 in air to 750, 850 and 950 ◦ C under applied stresses of 22, 54 and 78 MPa (750 ◦ C); 24, 59 and 83 MPa (850 ◦ C); and 18, 46 and 64 MPa (950 ◦ C). These stresses represented 10%, 25% and 35% of the YS values of Alloy 617 at these three temperatures. The magnitude of YS at these temperatures was determined from the engineering stress versus engineering strain diagrams generated from tensile testing performed according to the ASTM Designation E 8-04 [6]. The higher YS value of Alloy 617 at 850 ◦ C can be attributed to the occurrence of the yield strength anomaly phenomenon that has recently been reported elsewhere [7]. Alloy 276 was also tested in air at 750, 850 and 950 ◦ C under applied stresses of 51, 49 and 39 MPa, respectively for evaluation of its creep behavior. 3. Results and discussion The optical micrographs of Alloy 617 loaded to 10% of its YS (0.10YS) values at 750, 850 and 950 ◦ C are illustrated in Fig. 2a–c, respectively. Large austenitic grains, annealing twins, and precipitates were observed in all three micrographs. A recent study [8] performed on creep deformation of Alloy 617 suggests that these precipitates could consist of chromium carbide (Cr23 C6 ) and/or intermetallic phase of chromium and molybdenum (Cr–Mo). Austenitic grains, annealing twins and small precipitates were also observed in optical micrographs of Alloy 276, as shown in Fig. 3. The average grain diameters of both alloys were measured by mean linear intercept method [9]. The grain diameters of Alloys 617 and 276 in the as-machined condition were 0.097 and 0.099 mm, respectively. For Alloy 617, the average grain diameter varied between 0.098 and 0.126 mm when tested within a temperature range of 750–950 ◦ C under different applied stress levels. Similarly, grain diameter ranging from 0.101 to 0.127 mm was determined for Alloy 276 within the same temperature range. Considering grain sizes of both alloys in the as-machined condition, and at temperatures of 750, 850 and 950 ◦ C, their standard deviations were ±0.013 and 0.011 mm for Alloys 617 and 276, respectively. Together, the grain size data suggest that there is a tendency for the grain size of these alloys to slightly increase as the temperature/exposure
Fig. 1. Dimensions of creep testing specimen.
Fig. 2. Optical micrographs of Alloy 617, Kalling’s reagent (a) 750 ◦ C, 22 MPa, 1000 h; (b) 850 ◦ C, 24 MPa, 1000 h; (c) 950 ◦ C, 18 MPa, 216 h.
186
A.K. Roy et al. / Materials Science and Engineering A 520 (2009) 184–188
Fig. 3. Optical micrograph of Alloy 276 (51 MPa, 750 ◦ C, 1000 h), HCl + HNO3 + CuCl2 .
time increases, ranging from ASTM grain sizes within 4–3 for asmachined alloy to ASTM grain size of 3 after 950 ◦ C exposure. The results of creep testing involving Alloy 617 are shown in Fig. 4 in the form of % creep versus time as a function of temperature at applied stresses equivalent to 0.10YS (22, 24 and 18 MPa), 0.25YS (54, 59 and 46 MPa) and 0.35YS (78, 83 and 64 MPa). It is interesting to note that the magnitude of instantaneous elastic plus plastic strain resulting from the initial applied stress was relatively higher at elevated temperatures. The modulus of elasticity is known [10] to decrease with increasing temperature, which could possibly account for the enhanced anelastic strain at higher temperatures. Further, the primary creep curve was relatively shorter at higher initial applied stresses, irrespective of the testing temperature. At 18 MPa–950 ◦ C (Fig. 4(a)), this alloy exhibited a very short steadystate region, followed by an extended third stage. On the contrary, substantially longer secondary creep regions were observed in this alloy at 750 and 850 ◦ C under initial applied stresses of 22 and 24 MPa, respectively (Fig. 4(a)). With increasing applied stress levels (0.25YS and 0.35YS) and temperature (Fig. 4(b, c)), the steady-state region became shorter and finally disappeared at 950 ◦ C, showing only a steeper tertiary creep curve. Assuming that a structural material must not undergo creep deformation exceeding 1% strain following 1000 h of loading at different stress levels, it could be stated that Alloy 617 may not be capable of withstanding an operating temperature of 950 ◦ C at applied stresses above 10% of its YS value. Data shown in Fig. 4(b) indicate that Alloy 617 may be able to sustain an operating temperature of 850 ◦ C for periods of up to 850–900 h when loaded to an applied stress of 59 MPa. Further, this alloy reached a tertiary stage almost immediately when loaded to a higher stress level of 83 MPa at a similar temperature, as shown in Fig. 4(c). Thus, the inference is strong that Alloy 617 may not be suitable for application under sustained loading for 1000 h both at 0.25YS and 0.35YS stress levels at 850 ◦ C or higher. It is, however, interesting to note that this alloy was capable to sustain all three levels of stress at 750 ◦ C by virtue of its prolonged and stable steady-state creep rates even beyond 1000 h of testing (Fig. 4). Thus, based on the overall data, it may be concluded that Alloy 617 could be safely used as a heat exchanger material within an operating temperature of 750 ◦ C at applied stresses corresponding to its 0.10YS, 0.25YS and 0.35YS values. Perhaps, this alloy may also be considered for a similar application at 850 ◦ C at applied stresses not exceeding its 0.25YS value. The results of creep testing involving Alloy 276 under sustained loads corresponding to its 0.25YS values (51, 49 and 39 MPa) at three temperatures are illustrated in Fig. 5. These data indicate that this
Fig. 4. Creep curves of Alloy 617 vs. temperature and applied stress. (a) Applied stress = 0.10YS. (b) Applied stress = 0.25YS. (c) Applied stress = 0.35YS.
A.K. Roy et al. / Materials Science and Engineering A 520 (2009) 184–188
Fig. 5. Creep curves of Alloy 276 vs. temperature and applied stress.
alloy did not experience any significant steady-state creep at 950 ◦ C. Rather, it suffered from extensive deformation within the initial 200 h of testing, showing a very steep slope in the tertiary stage. With respect to its creep deformation at 850 ◦ C, this alloy exhibited strain exceeding 1% following 1000 h of loading at 49 MPa. However, at 750 ◦ C, this alloy was very close to meeting the maximum strain criterion of 1% under an initial applied stress of 51 MPa for a loading period of 1000 h. A comparison of steady-state creep rates of Alloys 617 and 276, subjected to applied stresses corresponding to their individual 0.25YS values at 750, 850 and 950 ◦ C, is illustrated in Fig. 6, clearly showing a superior creep-resistance of Alloy 617. Also, a variation of steady-state creep rate with applied stress for Alloy 617 at three tested temperatures is shown in Fig. 7. Literature data [8] on Alloy 617 have also been superimposed in this figure for comparison of its creep rates at similar temperatures. It is obvious that the creep rates obtained from this study were significantly higher compared to those cited in the literature, possibly due to the differences in the tested environmental conditions (air versus He + O2 ). The resultant data also indicate that, with the exception of 850 and 950 ◦ C, the normal temperature-compensated power law did not apply to this
Fig. 6. log10 (creep rate) vs. 1/T and applied stress.
187
Fig. 7. log10 (creep rate) vs. log10 (applied stress).
alloy, possibly due to the changes in microstructure and anomalous yield strength behavior at elevated temperatures. Alloys 617 and 276 are both solid-solution strengthened alloys. It has been reported [11,12] that carbides of M23 C6 and M6 C types, and Cr–Mo intermetallic phase can be precipitated at their grain boundaries and as intragranular particles during solution annealing at elevated temperatures. While both types of carbides can co-exist at the grain boundaries, M23 C6 is Cr-rich and M6 C has higher Mo content. However, a preferential precipitation of Cr-rich M23 C6 type carbide (Cr23 C6 ) appears to be more reasonable than Mo-rich M6 C carbide (Mo6 C) due to a greater diffusion coefficient of Cr than Mo in Ni-base alloys. The formation of precipitates in the matrix can also lead to the development of subgrains. The formation of subgrains and grain boundary precipitation can inhibit dislocation motion [7,8,13], thus preventing accelerated deformation rates of these alloys at 750 and 850 ◦ C under relatively lower applied stress levels (0.10YS and 0.25YS), leading to the occurrence of prolonged steady-state region in the creep curves. Fig. 8 illustrates dislocation pile-ups at the grain boundaries, and subgrain formation within the matrix of Alloy 617 tested under an applied stress of 59 MPa at 850 ◦ C. At 950 ◦ C, the unstable intragranular carbides and grain boundary carbides may undergo dissolution, subsequently causing migration of carbides and grain boundaries that could lead to the initiation of voids at the carbide phases existing on the grain boundaries. Finally, recrystallization occurs as creep deformation progresses with eventual breaking up of grains [14].
Fig. 8. TEM micrograph of Alloy 617 (59 MPa, 850 ◦ C, 1000 h).
188
A.K. Roy et al. / Materials Science and Engineering A 520 (2009) 184–188
4. Summary and conclusions Alloys 617 and 276 have been tested under sustained loading for evaluation of their creep deformation behavior at 750, 850 and 950 ◦ C. Alloy 617 has been loaded at stress levels corresponding to 10%, 25% and 35% of its YS values at these temperatures. However, creep deformation of Alloy 276 has been investigated only at applied stresses equivalent to 25% of its YS values at similar temperatures. The significant results and conclusions drawn from this study are given below: • The primary stage of creep was very short for both alloys irrespective of the testing temperature and the applied stress. • Both alloys exhibited prolonged and slower steady-state creep rate at 850 ◦ C when loaded to 25% of their YS values for a comparable duration. While Alloy 617 was very close to meeting the maximum strain criterion of 1% in 1000 h of loading, Alloy 276 suffered from excessive deformation. • Severe creep deformation, characterized by the formation of an instantaneous tertiary region, was observed with Alloy 617 when testing was performed at 850 and 950 ◦ C under applied stresses corresponding to 35% of its YS values (83 and 64 MPa) at these temperatures. However, this alloy met the deformation acceptance criterion at 950 ◦ C when loaded to its 0.10YS value. • Alloy 617 was capable of sustaining all three levels of applied stress (22, 54 and 78 MPa) for durations exceeding 1000 h at 750 ◦ C. Considering a maximum allowable strain of 1% following 1000 h of sustained loading, this alloy may be most suitable for use as a heat exchanger material at operating temperatures up to 750 ◦ C under applied stresses equivalent to its 0.10YS, 0.25YS and 0.35YS values. • Alloy 276 suffered from severe creep deformation at 950 ◦ C at an applied stress of 39 MPa (0.25YS), showing a steep tertiary curve.
However, this alloy was very close to meeting the maximum strain criterion at 750 ◦ C under an applied stress of 51 MPa (0.25YS). • Dissolution of carbides and intermetallic phases formed and precipitated during thermal treatment that can promote hardening of the tested materials, could possibly account for void formation leading to recrystallization and subsequent break-up of grains at 950 ◦ C. Acknowledgment This work was funded by the United States Department of Energy under grant number DE-FC07-04ID14566. References [1] W. Ren, R. Swindeman, ASME Pressure Vessels and Piping Division Conference, Texas, 2007. [2] B.S. Rao, H.P. Meurer, H. Schuster, Mater. Sci. Eng. A 104 (1988) 37–51. [3] W.L. Mankins, J.C. Hosier, T.H. Bassford, Metall. Mater. Trans. B 5 (1974) 2579–2590. [4] A.K. Roy, J. Pal, C. Mukhopadhyay, Mater. Sci. Eng. A 474 (2008) 363–370. [5] ASTM Designation E 139-2000, Creep Rupture and Stress-Rupture Tests of Metallic Materials, American Society for Testing and Materials (ASTM) International. [6] ASTM Designation E 8-2004, Standard Test Methods for Tensile Testing of Metallic Materials, American Society for Testing and Materials (ASTM) International. [7] A.K. Roy, V. Marthandam, Mater. Sci. Eng. A, (2009), doi:10.1016/j.mesa.2009. 03.090. [8] P.S. Shankar, K. Natesan, J. Nucl. Mater. 366 (2007) 26–36. [9] ASTM E 112-1996, Standard Test Methods for Determining Average Grain Size, American Society for Testing and Materials (ASTM) International. [10] A.K. Roy, J. Pal, M.H. Hasan, J. Eng. Mater. Technol., in press. [11] M. Raghavan, B.J. Berkowitz, J.C. Scanlon, Metall. Mater. Trans. A 13 (1982) 979–984. [12] H.M. Tawancy, J. Mater. Sci. 31 (1996) 3929–3936. [13] V. Marthandam, Tensile deformation, toughness and crack propagation studies of Alloy 617, Ph.D. Dissertation, Department of Mechanical Engineering, University of Nevada, Las Vegas, April 2008. [14] S. Kihara, J.B. Newkirk, A. Ohtomo, Y. Saiga, Metall. Mater. Trans. A 11A (1980) 1019–1031.