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Critical cleavage fracture stress characterization of A508 nuclear pressure vessel steels
International Journal of Pressure Vessels and Piping 123-124 (2014) 92e98
Contents lists available at ScienceDirect
International Journal of Pressure Vessels and Piping journal homepage: www.elsevier.com/locate/ijpvp
Critical cleavage fracture stress characterization of A508 nuclear pressure vessel steels Sujun Wu*, Huijin Jin, Yanbin Sun, Luowei Cao School of Materials Science and Engineering, Beihang University, 37 Xueyuan Road, Beijing 100191, China
a r t i c l e i n f o
a b s t r a c t
Article history: Received 9 October 2012 Received in revised form 11 August 2014 Accepted 19 August 2014 Available online 27 August 2014
The critical cleavage fracture stress of SA508 Gr.4N and SA508 Gr.3 low alloy reactor pressure vessel (RPV) steels was studied through the combination of experiments and finite element method (FEM) analysis. The results showed that the value of the local cleavage fracture stress, sF, of SA508 Gr.4N steel was significantly higher than that of SA508 Gr.3 steel. Detailed microstructural analysis was carried out using FEGSEM which revealed much smaller grains, finer and more homogenous carbide particles formed in SA508 Gr.4N steel. Compared with the SA508 Gr.3 steel currently used in the nuclear industry, the SA508 Gr.4N steel possesses higher strength and notch toughness as well as improved cleavage fracture behavior, and is considered a better candidate RPV steel for the next generation nuclear reactors. © 2014 Elsevier Ltd. All rights reserved.
Keywords: Reactor pressure vessel steels A508 steels Cleavage fracture stress
sF
1. Introduction Reactor pressure vessel (RPV) steels are required to endure the combined severe conditions of high pressure, high temperature and intense neutron irradiation. At present, requirements on material with high strength and toughness are rising to increase the power generation capacity and operation life of nuclear power plants [1e3]. Currently, SA508 Gr.3 low alloy steel is the widely used steel for the operating nuclear reactor vessels. However, SA508 Gr.4N low alloy steel with a higher Ni and Cr and lower Mn content compared to SA508 Gr.3 is considered as a candidate material for the new generation of nuclear reactors due to its excellent strength and toughness [4e7]. The low temperature brittle fracture of structural steels is often transgranular cleavage. The critical cleavage fracture stress is a critical factor controlling the cleavage fracture and fracture toughness, which is relatively independent of temperature and strain rate. Thus fracture stress is also one of the fracture criterion parameters connecting macroscopic fracture toughness with the steel microstructure [8e17]. Griffith and Owen [8] used a twodimensional plane strain finite element method (FEM) model to calculate the stress and strain distribution for four point bending (4 PB) specimens. Previous cleavage fracture models [13e16] assumed the peak value syymax ahead of the notch root of the
* Corresponding author. Tel.: þ86 010 8231 6326; fax: þ86 010 8231 7108. E-mail addresses: [email protected], [email protected] (S. Wu). http://dx.doi.org/10.1016/j.ijpvp.2014.08.003 0308-0161/© 2014 Elsevier Ltd. All rights reserved.
4 PB specimens, corresponding to the failure load, as the critical cleavage fracture stress. Some studies also suggested that the cleavage fracture stress sF could be used as an engineering notch toughness parameter of materials and for assessing integrity of structures with notch defects [11,12,18,19]. Previous researches [4e7] on SA508 Gr.4N steel have partly considered the microstructure, tensile properties, Charpy V-notch impact energy and fracture toughness. However, the critical cleavage fracture stress of SA508 Gr.4N steel has not been investigated yet. In this work, the local cleavage fracture stresses of both SA508 Gr.4N and SA508 Gr.3 low alloy steels have been calculated by combining the 4 PB experiments and FEM stress analysis. The micromechanism of cleavage fracture is further discussed based upon the experimental results. 2. Experimental 2.1. Materials and specimens SA508 Gr.4N steel is considered as the candidate for the new generation nuclear reactor pressure vessels. The comparison of its properties with the currently widely used SA508 Gr.3 is therefore of great significance. In this work, both RPV steels were studied which were produced in vacuum induction furnace as 50 kg ingots. The chemical compositions are shown in Table 1. The heat-treatment procedures of the RPV steels are given in Table 2. Metallographic specimens were polished and etched in 5% nital, and microstructures were analyzed by using optical microscopy and field emission
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Table 1 Chemical compositions of the delivery state SA508 steels (wt%). Materials
C
Ni
Cr
Mo
Mn
Si
P
S
SA508 Gr.4N SA508 Gr.3
0.15 0.13
3.26 0.63
1.66 0.19
0.46 0.50
0.34 1.30
0.36 0.21
0.011 0.011
0.008 0.008
Table 2 Heat-treatment conditions of SA508 steels. Materials SA508 Gr.4N SA508 Gr.3
Austenitizing
920 C/1.5 h 880 C/1 h
Cooling methods
Tempering
Water quenching Oil quenching
650 C/30 h 640 C/4 h
gun scanning electron microscopy (FEGSEM). Cylindrical specimens with 35 mm gage length and 5 mm diameter were used for the tensile tests. The dimensions and the loading mode of 4 PB notched specimen are shown in Fig. 1. Fracture surfaces of all 4 PB specimens were observed by SEM. 2.2. Mechanical tests Tensile tests and the 4 PB tests of SA508 Gr.4N and SA508 Gr.3 steels were carried out at both 160 C and 196 C with a strain rate of 1 mm/min on a hydraulic servo universal testing machine, and the curves of the load and load-line displacement for 4 PB test specimens were recorded. The fracture load Pf that refers to the load at failure was measured for the 4 PB specimens. The general yield load Pgy of the 4 PB specimens was calculated by the following equation [19]:
Pgy ¼
1:155Cf sy BðW aÞ2 2L
(1)
where B is the specimen thickness, W is the specimen width, a is the notch depth, Cf is the constraint factor taken as 1.22, and L ¼ 12.7 mm is the bending span. 2.3. Finite element method (FEM) calculation A two-dimensional model with four-node biquadratic plane strain reduced integration elements (CPE4R) was employed with the ABAQUS code. Only one-half of the geometry was analyzed due to the symmetry. The total number of elements and nodes were 7259 and 7395, respectively. The element arrangement and plot contours on deformed shape in the vicinity of the notch root are shown in Fig. 2, whose minimum element size is 14.5 mm at the notch tip. The yield strength sy and the true stressestrain curve were obtained from the tensile test results, while the maximum normal stress syy on the plane directly ahead of the notch was calculated based on FEM analysis and 4 PB test results obtained at both 160 C and 196 C.
Fig. 1. The dimensions and the loading mode of 4 PB notched specimens.
Fig. 2. Element arrangement (a) and the plot contours on deformed shape (b) in the vicinity of the notch root.
2.4. Characterization of the critical cleavage fracture stress At a measured Pf/Pgy obtained from 4 PB tests, corresponding stress distribution curves are obtained from FEM results. The local cleavage fracture stress sF was calculated from the normal stress distribution curves ahead of the notch root of 4 PB specimens.
3. Results and discussion 3.1. Microstructure characterization Metallographic observation was carried out for the two steels, and the optical and SEM micrographs are shown in Fig. 3. The analysis of Lee et al. [4] showed that the predominant microstructure of SA508 Gr.4N steel is tempered martensite, while most of the carbides in this steel are Cr-rich carbides, such as M7C3 and M23C6 type, as shown in Fig. 3(a) and (b). However, SA508 Gr.3 steel is tempered bainite microstructure with M3C-type carbides precipitated along lath boundaries and M2C-type carbides dispersed inside laths [20]. The quantitative analysis results of the size distribution of prior austenite grains and carbide particles of SA508 steels are given in Fig. 4 and Table 3. Consider the HallePetch equation [21,22]:
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Fig. 3. Microstructures of SA508 Gr.4N: (a) optical micrograph, (b) SEM micrograph and microstructures of SA508 Gr.3: (c) optical micrograph, (d) SEM micrograph.
Fig. 4. Histograms of the size distributions of prior austenite grains (a) and carbide particles (b) of SA508 Gr.4N, and the size distributions of prior austenite grains (c) and carbide particles (d) of SA508 Gr.3.
sy ¼ si þ ky d1=2
(2)
where sy is the yield strength, si is the stress at infinite grain size, ky is a constant, and d is the grain size. SA508 Gr.4N has smaller average grain size and finer carbide particles than that of SA508 Gr.3, which may contribute to the higher strength of the SA508 Gr.4N.
Table 3 Mean sizes of prior austenite grains and carbide particles of SA508 steels. Materials
Mean prior austenite grain size (mm)
Mean carbide size (mm)
SA508 Gr.4N SA508 Gr.3
16.67 21.65
0.15 0.25
S. Wu et al. / International Journal of Pressure Vessels and Piping 123-124 (2014) 92e98 Table 4 Tensile tests results of SA508 steels. Materials
Temperature ( C)
Yield strength (MPa)
Tensile strength (MPa)
Harden. exp. N
a
SA508 Gr.4N
160 196 160 196
880 1036 738 940
985 1140 835 1020
6.2 5.6 7.8 6.6
4.2 4.52 3.12 3.28
SA508 Gr.3
3.2. Mechanical properties 3.2.1. Tensile test results Tensile data for the SA508 Gr.4N and SA508 Gr.3 steels tested at both 160 C and 196 C are listed in Table 4. It is shown that the
95
yield strength and tensile strength of SA508 Gr.4N steel are higher than that of SA508 Gr.3, resulting from the difference in the microstructures. The strain hardening exponent, N, defined by the RambergeOsgood [23] stress strain law ε/εy ¼ s/sy þ a(s/sy)N (where a is a material constant). The least squares method for fitting the RambergeOsgood curve was used to determine the values of N and a from the true stressestrain data. Taking the logarithm of both sides of RambergeOsgood stress strain law, this leads to the following equation:
ε s s ¼ log a þ N log log εy sy sy
(3)
Linearization of Eq. (3) leads to.
Y ¼ log a þ NX
(4)
where Y ¼ logðε=εy s=sy Þ, and X ¼ logðs=sy Þ, sy and εy represent the yielding stress and the corresponding strain. While fitting the line, Eq. (4), only those points should be taken into account where the value of the plastic strain is higher than the established limit εy. The values of N and a are given in Table 4. It can be seen that the strain hardening exponent of SA508 Gr.4N is slightly lower than that of SA508 Gr.3, which implies that the SA508 Gr.4N has relatively higher strain hardening ability. Fig. 5. The value of Pf/Pgy at fracture for each 4 PB specimen of SA508 Gr.4N and SA508 Gr.3 steels.
Fig. 6. The typical cleavage fracture behavior for (a) SA508 Gr.4N and (b) SA508 Gr.3 steels.
Fig. 7. Distribution of syy/sy ahead of notch root for 4 PB specimens tested at (a) 160 C and (b) 196 C of SA508 Gr.4N steel calculated by FEM.
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3.2.2. 4 PB test results and fracture surface observation Fracture load Pf of SA508 Gr.4N and SA508 Gr.3 steels was measured from 4 PB tests. The value of Pf/Pgy for each 4 PB specimen is given in Fig. 5. It is difficult to avoid notch defects and notch-like geometries in the engineering structures and components [17]. Thus, it is important to evaluate the notch toughness of SA508 Gr.4N steels. In Wang et al.'s study [11], the notch toughness was characterized by Pf/Pgy for fracture. 4 PB test results show that the notch toughness is much higher at 160 C than that at 196 C for both SA508 Gr.4N and SA508 Gr.3 steels. Fractographic observation was carried out on the fracture surfaces of SA508 Gr.4N and SA508 Gr.3 steels. The typical cleavage fracture behavior for both steels was revealed as shown in Fig. 6. It was found that all specimens of both SA508 Gr.4N and SA508 Gr.3 steels tested at both 160 C and 196 C fractured in cleavage manner with only a few of them showing the principal initiation site ahead of the notch root. 3.3. FEM calculation results FEM calculation results of stress distribution syy/sy ahead of the notch tip of 4 PB specimens of SA508 Gr.4N and SA508 Gr.3 steels tested at both 160 C and 196 C are shown in Figs. 7 and 8, respectively, where P is the applied load, X is the distance from the notch root. The results demonstrate that the normal stress syy increases with increasing Pf/Pgy, and the peak value of syy moves away from the notch root, resulting in the extension of area subject to high syy implying higher energy was absorbed before failure for both SA508 Gr.4N and SA508 Gr.3 steels. The loadedisplacement curves obtained by FE analyses compared with the actual test loadedisplacement curves at 196 C for both steels are shown in
Fig. 8. Distribution of syy/sy ahead of notch root for 4 PB specimens tested at (a) 160 C and (b) 196 C of SA508 Gr.3 steel calculated by FEM.
Fig. 9. It can be seen that the loadedisplacement curves based on FE analyses are well consistent with the experimental results for both steels. 3.4. Local cleavage fracture stress (sF) The local cleavage fracture stress, sF, can be obtained from the normal stress distribution curve ahead of notch root and the experimental results of the 4 PB specimens. The relationship between the maximum value of syy/sy (Figs. 7 and 8) and the corresponding applied load ratio P/Pgy of SA508 Gr.4N and SA508 Gr.3 steels is shown in Fig. 10. The maximum syy value (syymax) corresponding to the load ratio at fracture of the specimen (Pf/Pgy) is considered as the local cleavage fracture stress sF. Fig. 11 presents the values of local cleavage fracture stress, sF, for both the SA508 Gr.4N and SA508 Gr.3 steels. It can be seen from Fig. 11 that the sF values of the SA508 Gr.4N steel are significantly higher than those of SA508 Gr.3 steel tested at both 160 C and 196 C. The sF values of SA508 Gr.4N steel are in a range of 2005e2150 MPa at 160 C and 2040e2228 MPa at 196 C, respectively. In general, SA508 Gr.4N steel is considered to have a much better cleavage fracture behavior than SA508 Gr.3 steel. The syymax related to Pf/Pgy represents the intensity of the stress field ahead of a notch (Figs. 7, 8 and 10). Therefore, sF is closely related to the macroscopic notch toughness Pf/Pgy. sF is proposed to
Fig. 9. The comparison of loadedisplacement curves between FE analyses and experimental data at 196 C for (a) SA508 Gr.4N and (b) SA508 Gr.3 steels.
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97
particle, and the cleavage fracture stress of such steels decreased with carbide particle radius increasing. Therefore, sF also can be calculated by the following equation:
1=2 sF ¼ pEgp = 1 y2 C0
(5)
where E is Young's modulus, gp is effective surface energy, y is Poisson's ratio, C0 is related to the size of relevant microstructural feature. Therefore, the scatter of sF demonstrates that it depends on the size distribution of the weakest constituent (large grains and large second-phase particles) [12e17,25]. sF indicates the resistance to cleavage fracture which includes the critical event of crack nucleation and propagation. In the crack nucleation process, the carbide particle size distribution plays a dominant role, while the grain size distribution plays the major role in the propagation process. Thus, sF can be used to analyze the micromechanism of cleavage fracture, and the value of sF is related to the microstructure of steels. According to the quantitative analysis (Fig. 4) in this work, the values of sF of SA508 Gr.4N steel are much higher than that of SA508 Gr.3 which can be attributed to the finer and more homogenous carbide particles and the smaller grains in SA508 Gr.4N steel. As reported by Wu and Knott [13], the fracture toughness KIC values of steels up to 80 MN/m3/2 can be predicted from the critical local cleavage fracture stress, sF, using the well established RKR model [9]. The curves between the predicted fracture toughness KIC and the test temperature can be obtained from which the ductileebrittle transition temperature can be determined. Therefore, sF can also be used as a fracture criterion parameter connecting macroscopic cleavage fracture and toughness with microstructures in A508 nuclear pressure vessel steels. 4. Conclusions Fig. 10. The variation of maximum value of syy/sy with applied load ratio Pf/Pgy for (a) SA508 Gr.4N and (b) SA508 Gr.3 steels.
be an engineering notch toughness parameter for materials and of value for assessing the integrity of structures with notch defects by the criterion syymax sF [17]. Previous studies [16,24] have shown that the dominant factor for cleavage fracture of steels is the ferrite grain size, and the local fracture stress sF is related to the large grains rather than the mean value of the grain size distribution. It can be predicted from the coarse grain distributions that the value of sF is lower and the cleavage fracture is easier to occur when compared to fine grain patches. Curry and Knott [25] suggested that cleavage fracture in spheroidized steels was due to the propagation of penny-shaped crack nuclei produced in a second phase carbide
(1) Both the yield and tensile strengths of SA508 Gr.4N steel are higher than that of SA508 Gr.3 steel, while the RambergeOsgood strain hardening exponent, N, of SA508 Gr.4N steel is lower than that of SA508 Gr.3 at both 160 C and 196 C, respectively. (2) The notch toughness of the SA508 Gr.4N steel characterized by Pf/Pgy and the area under the loadedisplacement curve is higher than that of SA508 Gr.3 steel. (3) The local cleavage fracture stress, sF, of the SA508 Gr.4N steel is significantly higher than that of SA508 Gr.3 steel, which can be attributed to smaller grains, finer and more homogenous carbide particles in SA508 Gr.4N steel. Acknowledgment The authors would like to acknowledge the financial support from the National Science and Technology Key Project: Life Management Technology of Nuclear Power Plant (2011ZX06004-002) and the Fundamental Research Funds for the Central Universities with Grant No. YWF-10-01-B20.The authors would also like to thank Prof. PEJ Flewitt of the University of Bristol for helpful discussions. References
Fig. 11. The cleavage fracture stress sF of SA508 Gr.4N and SA508 Gr.3 steels.
[1] Kim SH, Im YR, Lee SH, Lee HC, Oh YJ, Hong JH. Effects of alloying elements on mechanical and fracture properties of base metals and simulated heataffected zones of SA508 steels. Metall Trans A 2001;32A:903e12. [2] Im YR, Oh YJ, Lee BJ, Hong JH, Lee HC. Effects of carbide precipitation on the strength and Charpy impact properties of low carbon Mn-Ni-Mo bainitic steels. J Nucl Mater 2001;297:138e48.
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[3] Chhibber R, Singh H, Arora N, Dutta BK. Micromechanical modeling of reactor pressure vessel steel. Mater Des 2012;36:258e74. [4] Lee BS, Kim MC, Hong JH. Characterization of high strength and high toughness Ni-Mo-Cr alloy steels for nuclear application. Int J Pres Vessels Pip 2010;87:74e80. [5] Lee KH, Kim MC, Lee BS, Wee DM. Analysis of the master curve approach on the fracture toughness properties of SA508 Gr.4N NieMoeCr low alloy steels for reactor pressure vessels. J Nucl Mater 2010;403:68e74. [6] Lee KH, Park SG, Kim MC, Lee BS, Wee DM. Characterization of transition behavior in SA508 Gr.4N NieCreMo low alloy steels with microstructural alteration by Ni and Cr contents. Mater Sci Eng A 2011;529:156e63. [7] Lee KH, Park SG, Kim MC, Lee BS. Cleavage fracture toughness of tempered martensitic Ni-Cr-Mo low alloy steel with different martensite fraction. Mater Sci Eng A 2012;534:75e82. [8] Griffith JR, Owen DRJ. An elastic-plastic stress analysis for a notched bar in plane strain bending. J Mech Phys Solids 1971;19:419e31. [9] Ritchie RO, Knott JF, Rice JR. On the relationship between critical tensile stress and fracture toughness in mild steel. J Mech Phys Solids 1973;21:395e410. [10] Ritchie RO, Server WL, Wullaert RA. Critical fracture stress and fracture strain models for the prediction of lower and upper shelf toughness in nuclear pressure vessel steels. Metall Trans A 1979;10:1557e70. [11] Wang GZ, Liu YG, Chen JH. Investigation of cleavage fracture initiation in notched specimens of a C-Mn steel with carbides and inclusions. Mater Sci Eng A 2004;369:181e91. [12] Wang GZ, Wang YL, Chen JH. Effects of loading rate on the local cleavage fracture stress sf in notched specimens. Eng Fract Mech 2005;72:675e89. [13] Wu SJ, Knott JF. Effects of degradation on the mechanical properties and fracture toughness of a steel pressure-vessel weld metal. Int J Pres Vessels Pip 2003;80:807e15.
[14] Ritchie RO, Thompson AW. On macroscopic and microscopic analyses for crack initiation and crack growth toughness in ductile alloys. Metall Trans A 1985;16A:233e48. [15] Lin T, Evans AG, Ritchie RO. A statistical model of brittle fracture by transgranular cleavage. J Mech Phys Solids 1986;34:477e97. [16] Wu SJ, Davis CL. Effect of duplex ferrite grain size distribution on local fracture stresses of Nb-microalloyed steels. Mater Sci Eng A 2004;387e389:456e60. [17] Wang GZ, Wang YL, Xuan FZ, Tu ST, Wang ZD. Cleavage fracture behavior of a C-Mn vessel steel at various loading rates in notched specimens. Int J Pres Vessels Pip 2008;85:720e7. [18] Samant AV, Lewandowski JJ. Effects of test temperature, grain size, and alloy additions on the cleavage fracture stress of polycrystalline niobium. Metall Trans A 1997;28A:389e99. [19] Alexander DJ, Lewandowski JJ, Sisak WJ, Thompson AW. Yielding and work hardening effects in notched bend bars. J Mech Phys Solids 1986;34:433. [20] Lee S, Kim S, Hwang B, Lee BS, Lee CG. Effect of carbide distribution on the fracture toughness in the transition temperature region of an SA 508 steel. Acta Mater 2002;50:4755e62. [21] Hall EO. The deformation and ageing of mild steel. 3. discussion of results. Proc Phys Soc 1951;64B:747. [22] Petch NJ. The cleavage strength of polycrystals. J Iron Steel Inst 1953;174: 25e8. [23] Ramberg W, Osgood WR. Description of stress-strain curves by three parameters. Technical Report, Technical Note No. 902. NACA; 1943. p. 4. [24] Shehata MT, Boyd JD. Advances in physical metallurgy and applications of steel. In: Proc. Conf., University of Liverpool, Liverpool, Sept. 1981. Warrendale, PA: TMS; 1982. p. 229e36. [25] Curry DA, Knott JF. Effect of microstructure on cleavage fracture stress in steel. Metal Sci 1978:511e4.
Contents lists available at ScienceDirect
International Journal of Pressure Vessels and Piping journal homepage: www.elsevier.com/locate/ijpvp
Critical cleavage fracture stress characterization of A508 nuclear pressure vessel steels Sujun Wu*, Huijin Jin, Yanbin Sun, Luowei Cao School of Materials Science and Engineering, Beihang University, 37 Xueyuan Road, Beijing 100191, China
a r t i c l e i n f o
a b s t r a c t
Article history: Received 9 October 2012 Received in revised form 11 August 2014 Accepted 19 August 2014 Available online 27 August 2014
The critical cleavage fracture stress of SA508 Gr.4N and SA508 Gr.3 low alloy reactor pressure vessel (RPV) steels was studied through the combination of experiments and finite element method (FEM) analysis. The results showed that the value of the local cleavage fracture stress, sF, of SA508 Gr.4N steel was significantly higher than that of SA508 Gr.3 steel. Detailed microstructural analysis was carried out using FEGSEM which revealed much smaller grains, finer and more homogenous carbide particles formed in SA508 Gr.4N steel. Compared with the SA508 Gr.3 steel currently used in the nuclear industry, the SA508 Gr.4N steel possesses higher strength and notch toughness as well as improved cleavage fracture behavior, and is considered a better candidate RPV steel for the next generation nuclear reactors. © 2014 Elsevier Ltd. All rights reserved.
Keywords: Reactor pressure vessel steels A508 steels Cleavage fracture stress
sF
1. Introduction Reactor pressure vessel (RPV) steels are required to endure the combined severe conditions of high pressure, high temperature and intense neutron irradiation. At present, requirements on material with high strength and toughness are rising to increase the power generation capacity and operation life of nuclear power plants [1e3]. Currently, SA508 Gr.3 low alloy steel is the widely used steel for the operating nuclear reactor vessels. However, SA508 Gr.4N low alloy steel with a higher Ni and Cr and lower Mn content compared to SA508 Gr.3 is considered as a candidate material for the new generation of nuclear reactors due to its excellent strength and toughness [4e7]. The low temperature brittle fracture of structural steels is often transgranular cleavage. The critical cleavage fracture stress is a critical factor controlling the cleavage fracture and fracture toughness, which is relatively independent of temperature and strain rate. Thus fracture stress is also one of the fracture criterion parameters connecting macroscopic fracture toughness with the steel microstructure [8e17]. Griffith and Owen [8] used a twodimensional plane strain finite element method (FEM) model to calculate the stress and strain distribution for four point bending (4 PB) specimens. Previous cleavage fracture models [13e16] assumed the peak value syymax ahead of the notch root of the
* Corresponding author. Tel.: þ86 010 8231 6326; fax: þ86 010 8231 7108. E-mail addresses: [email protected], [email protected] (S. Wu). http://dx.doi.org/10.1016/j.ijpvp.2014.08.003 0308-0161/© 2014 Elsevier Ltd. All rights reserved.
4 PB specimens, corresponding to the failure load, as the critical cleavage fracture stress. Some studies also suggested that the cleavage fracture stress sF could be used as an engineering notch toughness parameter of materials and for assessing integrity of structures with notch defects [11,12,18,19]. Previous researches [4e7] on SA508 Gr.4N steel have partly considered the microstructure, tensile properties, Charpy V-notch impact energy and fracture toughness. However, the critical cleavage fracture stress of SA508 Gr.4N steel has not been investigated yet. In this work, the local cleavage fracture stresses of both SA508 Gr.4N and SA508 Gr.3 low alloy steels have been calculated by combining the 4 PB experiments and FEM stress analysis. The micromechanism of cleavage fracture is further discussed based upon the experimental results. 2. Experimental 2.1. Materials and specimens SA508 Gr.4N steel is considered as the candidate for the new generation nuclear reactor pressure vessels. The comparison of its properties with the currently widely used SA508 Gr.3 is therefore of great significance. In this work, both RPV steels were studied which were produced in vacuum induction furnace as 50 kg ingots. The chemical compositions are shown in Table 1. The heat-treatment procedures of the RPV steels are given in Table 2. Metallographic specimens were polished and etched in 5% nital, and microstructures were analyzed by using optical microscopy and field emission
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Table 1 Chemical compositions of the delivery state SA508 steels (wt%). Materials
C
Ni
Cr
Mo
Mn
Si
P
S
SA508 Gr.4N SA508 Gr.3
0.15 0.13
3.26 0.63
1.66 0.19
0.46 0.50
0.34 1.30
0.36 0.21
0.011 0.011
0.008 0.008
Table 2 Heat-treatment conditions of SA508 steels. Materials SA508 Gr.4N SA508 Gr.3
Austenitizing
920 C/1.5 h 880 C/1 h
Cooling methods
Tempering
Water quenching Oil quenching
650 C/30 h 640 C/4 h
gun scanning electron microscopy (FEGSEM). Cylindrical specimens with 35 mm gage length and 5 mm diameter were used for the tensile tests. The dimensions and the loading mode of 4 PB notched specimen are shown in Fig. 1. Fracture surfaces of all 4 PB specimens were observed by SEM. 2.2. Mechanical tests Tensile tests and the 4 PB tests of SA508 Gr.4N and SA508 Gr.3 steels were carried out at both 160 C and 196 C with a strain rate of 1 mm/min on a hydraulic servo universal testing machine, and the curves of the load and load-line displacement for 4 PB test specimens were recorded. The fracture load Pf that refers to the load at failure was measured for the 4 PB specimens. The general yield load Pgy of the 4 PB specimens was calculated by the following equation [19]:
Pgy ¼
1:155Cf sy BðW aÞ2 2L
(1)
where B is the specimen thickness, W is the specimen width, a is the notch depth, Cf is the constraint factor taken as 1.22, and L ¼ 12.7 mm is the bending span. 2.3. Finite element method (FEM) calculation A two-dimensional model with four-node biquadratic plane strain reduced integration elements (CPE4R) was employed with the ABAQUS code. Only one-half of the geometry was analyzed due to the symmetry. The total number of elements and nodes were 7259 and 7395, respectively. The element arrangement and plot contours on deformed shape in the vicinity of the notch root are shown in Fig. 2, whose minimum element size is 14.5 mm at the notch tip. The yield strength sy and the true stressestrain curve were obtained from the tensile test results, while the maximum normal stress syy on the plane directly ahead of the notch was calculated based on FEM analysis and 4 PB test results obtained at both 160 C and 196 C.
Fig. 1. The dimensions and the loading mode of 4 PB notched specimens.
Fig. 2. Element arrangement (a) and the plot contours on deformed shape (b) in the vicinity of the notch root.
2.4. Characterization of the critical cleavage fracture stress At a measured Pf/Pgy obtained from 4 PB tests, corresponding stress distribution curves are obtained from FEM results. The local cleavage fracture stress sF was calculated from the normal stress distribution curves ahead of the notch root of 4 PB specimens.
3. Results and discussion 3.1. Microstructure characterization Metallographic observation was carried out for the two steels, and the optical and SEM micrographs are shown in Fig. 3. The analysis of Lee et al. [4] showed that the predominant microstructure of SA508 Gr.4N steel is tempered martensite, while most of the carbides in this steel are Cr-rich carbides, such as M7C3 and M23C6 type, as shown in Fig. 3(a) and (b). However, SA508 Gr.3 steel is tempered bainite microstructure with M3C-type carbides precipitated along lath boundaries and M2C-type carbides dispersed inside laths [20]. The quantitative analysis results of the size distribution of prior austenite grains and carbide particles of SA508 steels are given in Fig. 4 and Table 3. Consider the HallePetch equation [21,22]:
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Fig. 3. Microstructures of SA508 Gr.4N: (a) optical micrograph, (b) SEM micrograph and microstructures of SA508 Gr.3: (c) optical micrograph, (d) SEM micrograph.
Fig. 4. Histograms of the size distributions of prior austenite grains (a) and carbide particles (b) of SA508 Gr.4N, and the size distributions of prior austenite grains (c) and carbide particles (d) of SA508 Gr.3.
sy ¼ si þ ky d1=2
(2)
where sy is the yield strength, si is the stress at infinite grain size, ky is a constant, and d is the grain size. SA508 Gr.4N has smaller average grain size and finer carbide particles than that of SA508 Gr.3, which may contribute to the higher strength of the SA508 Gr.4N.
Table 3 Mean sizes of prior austenite grains and carbide particles of SA508 steels. Materials
Mean prior austenite grain size (mm)
Mean carbide size (mm)
SA508 Gr.4N SA508 Gr.3
16.67 21.65
0.15 0.25
S. Wu et al. / International Journal of Pressure Vessels and Piping 123-124 (2014) 92e98 Table 4 Tensile tests results of SA508 steels. Materials
Temperature ( C)
Yield strength (MPa)
Tensile strength (MPa)
Harden. exp. N
a
SA508 Gr.4N
160 196 160 196
880 1036 738 940
985 1140 835 1020
6.2 5.6 7.8 6.6
4.2 4.52 3.12 3.28
SA508 Gr.3
3.2. Mechanical properties 3.2.1. Tensile test results Tensile data for the SA508 Gr.4N and SA508 Gr.3 steels tested at both 160 C and 196 C are listed in Table 4. It is shown that the
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yield strength and tensile strength of SA508 Gr.4N steel are higher than that of SA508 Gr.3, resulting from the difference in the microstructures. The strain hardening exponent, N, defined by the RambergeOsgood [23] stress strain law ε/εy ¼ s/sy þ a(s/sy)N (where a is a material constant). The least squares method for fitting the RambergeOsgood curve was used to determine the values of N and a from the true stressestrain data. Taking the logarithm of both sides of RambergeOsgood stress strain law, this leads to the following equation:
ε s s ¼ log a þ N log log εy sy sy
(3)
Linearization of Eq. (3) leads to.
Y ¼ log a þ NX
(4)
where Y ¼ logðε=εy s=sy Þ, and X ¼ logðs=sy Þ, sy and εy represent the yielding stress and the corresponding strain. While fitting the line, Eq. (4), only those points should be taken into account where the value of the plastic strain is higher than the established limit εy. The values of N and a are given in Table 4. It can be seen that the strain hardening exponent of SA508 Gr.4N is slightly lower than that of SA508 Gr.3, which implies that the SA508 Gr.4N has relatively higher strain hardening ability. Fig. 5. The value of Pf/Pgy at fracture for each 4 PB specimen of SA508 Gr.4N and SA508 Gr.3 steels.
Fig. 6. The typical cleavage fracture behavior for (a) SA508 Gr.4N and (b) SA508 Gr.3 steels.
Fig. 7. Distribution of syy/sy ahead of notch root for 4 PB specimens tested at (a) 160 C and (b) 196 C of SA508 Gr.4N steel calculated by FEM.
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3.2.2. 4 PB test results and fracture surface observation Fracture load Pf of SA508 Gr.4N and SA508 Gr.3 steels was measured from 4 PB tests. The value of Pf/Pgy for each 4 PB specimen is given in Fig. 5. It is difficult to avoid notch defects and notch-like geometries in the engineering structures and components [17]. Thus, it is important to evaluate the notch toughness of SA508 Gr.4N steels. In Wang et al.'s study [11], the notch toughness was characterized by Pf/Pgy for fracture. 4 PB test results show that the notch toughness is much higher at 160 C than that at 196 C for both SA508 Gr.4N and SA508 Gr.3 steels. Fractographic observation was carried out on the fracture surfaces of SA508 Gr.4N and SA508 Gr.3 steels. The typical cleavage fracture behavior for both steels was revealed as shown in Fig. 6. It was found that all specimens of both SA508 Gr.4N and SA508 Gr.3 steels tested at both 160 C and 196 C fractured in cleavage manner with only a few of them showing the principal initiation site ahead of the notch root. 3.3. FEM calculation results FEM calculation results of stress distribution syy/sy ahead of the notch tip of 4 PB specimens of SA508 Gr.4N and SA508 Gr.3 steels tested at both 160 C and 196 C are shown in Figs. 7 and 8, respectively, where P is the applied load, X is the distance from the notch root. The results demonstrate that the normal stress syy increases with increasing Pf/Pgy, and the peak value of syy moves away from the notch root, resulting in the extension of area subject to high syy implying higher energy was absorbed before failure for both SA508 Gr.4N and SA508 Gr.3 steels. The loadedisplacement curves obtained by FE analyses compared with the actual test loadedisplacement curves at 196 C for both steels are shown in
Fig. 8. Distribution of syy/sy ahead of notch root for 4 PB specimens tested at (a) 160 C and (b) 196 C of SA508 Gr.3 steel calculated by FEM.
Fig. 9. It can be seen that the loadedisplacement curves based on FE analyses are well consistent with the experimental results for both steels. 3.4. Local cleavage fracture stress (sF) The local cleavage fracture stress, sF, can be obtained from the normal stress distribution curve ahead of notch root and the experimental results of the 4 PB specimens. The relationship between the maximum value of syy/sy (Figs. 7 and 8) and the corresponding applied load ratio P/Pgy of SA508 Gr.4N and SA508 Gr.3 steels is shown in Fig. 10. The maximum syy value (syymax) corresponding to the load ratio at fracture of the specimen (Pf/Pgy) is considered as the local cleavage fracture stress sF. Fig. 11 presents the values of local cleavage fracture stress, sF, for both the SA508 Gr.4N and SA508 Gr.3 steels. It can be seen from Fig. 11 that the sF values of the SA508 Gr.4N steel are significantly higher than those of SA508 Gr.3 steel tested at both 160 C and 196 C. The sF values of SA508 Gr.4N steel are in a range of 2005e2150 MPa at 160 C and 2040e2228 MPa at 196 C, respectively. In general, SA508 Gr.4N steel is considered to have a much better cleavage fracture behavior than SA508 Gr.3 steel. The syymax related to Pf/Pgy represents the intensity of the stress field ahead of a notch (Figs. 7, 8 and 10). Therefore, sF is closely related to the macroscopic notch toughness Pf/Pgy. sF is proposed to
Fig. 9. The comparison of loadedisplacement curves between FE analyses and experimental data at 196 C for (a) SA508 Gr.4N and (b) SA508 Gr.3 steels.
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particle, and the cleavage fracture stress of such steels decreased with carbide particle radius increasing. Therefore, sF also can be calculated by the following equation:
1=2 sF ¼ pEgp = 1 y2 C0
(5)
where E is Young's modulus, gp is effective surface energy, y is Poisson's ratio, C0 is related to the size of relevant microstructural feature. Therefore, the scatter of sF demonstrates that it depends on the size distribution of the weakest constituent (large grains and large second-phase particles) [12e17,25]. sF indicates the resistance to cleavage fracture which includes the critical event of crack nucleation and propagation. In the crack nucleation process, the carbide particle size distribution plays a dominant role, while the grain size distribution plays the major role in the propagation process. Thus, sF can be used to analyze the micromechanism of cleavage fracture, and the value of sF is related to the microstructure of steels. According to the quantitative analysis (Fig. 4) in this work, the values of sF of SA508 Gr.4N steel are much higher than that of SA508 Gr.3 which can be attributed to the finer and more homogenous carbide particles and the smaller grains in SA508 Gr.4N steel. As reported by Wu and Knott [13], the fracture toughness KIC values of steels up to 80 MN/m3/2 can be predicted from the critical local cleavage fracture stress, sF, using the well established RKR model [9]. The curves between the predicted fracture toughness KIC and the test temperature can be obtained from which the ductileebrittle transition temperature can be determined. Therefore, sF can also be used as a fracture criterion parameter connecting macroscopic cleavage fracture and toughness with microstructures in A508 nuclear pressure vessel steels. 4. Conclusions Fig. 10. The variation of maximum value of syy/sy with applied load ratio Pf/Pgy for (a) SA508 Gr.4N and (b) SA508 Gr.3 steels.
be an engineering notch toughness parameter for materials and of value for assessing the integrity of structures with notch defects by the criterion syymax sF [17]. Previous studies [16,24] have shown that the dominant factor for cleavage fracture of steels is the ferrite grain size, and the local fracture stress sF is related to the large grains rather than the mean value of the grain size distribution. It can be predicted from the coarse grain distributions that the value of sF is lower and the cleavage fracture is easier to occur when compared to fine grain patches. Curry and Knott [25] suggested that cleavage fracture in spheroidized steels was due to the propagation of penny-shaped crack nuclei produced in a second phase carbide
(1) Both the yield and tensile strengths of SA508 Gr.4N steel are higher than that of SA508 Gr.3 steel, while the RambergeOsgood strain hardening exponent, N, of SA508 Gr.4N steel is lower than that of SA508 Gr.3 at both 160 C and 196 C, respectively. (2) The notch toughness of the SA508 Gr.4N steel characterized by Pf/Pgy and the area under the loadedisplacement curve is higher than that of SA508 Gr.3 steel. (3) The local cleavage fracture stress, sF, of the SA508 Gr.4N steel is significantly higher than that of SA508 Gr.3 steel, which can be attributed to smaller grains, finer and more homogenous carbide particles in SA508 Gr.4N steel. Acknowledgment The authors would like to acknowledge the financial support from the National Science and Technology Key Project: Life Management Technology of Nuclear Power Plant (2011ZX06004-002) and the Fundamental Research Funds for the Central Universities with Grant No. YWF-10-01-B20.The authors would also like to thank Prof. PEJ Flewitt of the University of Bristol for helpful discussions. References
Fig. 11. The cleavage fracture stress sF of SA508 Gr.4N and SA508 Gr.3 steels.
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