- Email: [email protected]
Characterization of He embrittlement of a 9Cr–1Mo steel using local approach of brittle fracture
Available online at www.sciencedirect.com
Engineering Fracture Mechanics 75 (2008) 3570–3580 www.elsevier.com/locate/engfracmech
Characterization of He embrittlement of a 9Cr–1Mo steel using local approach of brittle fracture J. Malaplate a b
a,*
, L. Vincent a, X. Averty b, J. Henry a, B. Marini
a
CEA Saclay, DEN-DANS/DMN/SRMA 91191, Gif-sur-Yvette Cedex, France CEA Saclay, DEN-DANS/DMN/SEMI 91191, Gif-sur-Yvette Cedex, France
Received 22 September 2006; received in revised form 22 December 2006; accepted 12 February 2007 Available online 27 February 2007
Abstract Ferritic-martensitic steels are prime candidate materials for future reactors. We present here the results of a study on the effects of helium implantation on the fracture behavior of 9Cr (T91) martensitic steels. Three-points static bending tests were performed at room temperature on implanted specimens and at 170 C on un-implanted material. All these tests led to brittle fracture. Based on a mechanical analysis of the tests results using Finite Element calculations, we have proposed that the mechanism of brittle fracture is controlled by a double criterion depending on implantation temperature and helium content. Furthermore, by applying the Beremin model, the toughness of helium implanted steel has been evaluated. 2007 Elsevier Ltd. All rights reserved. Keywords: Helium embrittlement; 9Cr martensitic steels; Brittle fracture; Bending tests; Finite element analysis
1. Introduction Ferritic/martensitic steels are prime candidate materials for future fusion reactors and spallation neutron sources due to their excellent thermo-mechanical properties and radiation resistance at temperature above 400 C in a fission spectrum. However, high helium contents will be produced in these materials during operation, which can have a detrimental effect on mechanical properties. Up to now, it has been well established that ferritic/martensitic steels are relatively immune to the so called ‘‘high temperature’’ helium embrittlement phenomenon (see for instance [1]). As regards helium effects at lower temperature on hardening and embrittlement, data are scarce and this topic is still a matter of controversy (see for instance [2,3] for tensile properties and [4–6] for fracture properties). Therefore, the aim of this work is to advance the characterization and understanding of intrinsic effect of helium at low temperature on fracture properties of 9Cr martensitic steel. To this end, three-points static bending tests have been performed at room temperature on 9Cr–1Mo–V–Nb martensitic steel (commercially called
*
Corresponding author. E-mail address: [email protected] (J. Malaplate).
0013-7944/$ - see front matter 2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.engfracmech.2007.02.019
J. Malaplate et al. / Engineering Fracture Mechanics 75 (2008) 3570–3580
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Nomenclature ry req reqmin reqmax r1 rf rfmin rfmax f r rc C r Pr ru þ r u ; ru m m, m+ rw Vp V0 Kj
yield stress equivalent von Mises stress equivalent von Mises stress for the lowest value of the opening displacement equivalent von Mises stress for the highest value of the opening displacement maximum principal stress maximum stress at fracture (‘‘pop-in’’) maximum stress at ‘‘pop-in’’ for the lowest value of the opening displacement maximum stress at ‘‘pop-in’’ for the highest value of the opening displacement average maximum stress at fracture critical fracture stress average critical fracture stress failure probability material constant defined in the Beremin model confidence interval of ru Weibull modulus confidence interval of m Weibull stress plastic volume representative elementary volume stress intensity factor
T91) implanted with helium at different contents and/or temperatures. Bending tests have also been performed on pristine materials at low temperature. The goal was to compare the fracture properties of T91 loaded with helium to those of un-implanted T91 in the brittle domain. Results of tests have been analyzed by Finite Element (FE). A preliminary microstructural characterization has also been carried out on miniature tensile specimens by Transmission Electron Microscopy (TEM). 2. Experimental 2.1. Materials, specimens and helium implantation The studied steel is a mod-9Cr–1Mo steel, i.e. 9Cr–1Mo–V–Nb commercially referred to as T91, received in the normalized and tempered metallurgical condition (solution annealed for 1 h at 1040 C, fast cool, reheated 1 h at 760 C, fast cool). The chemical composition is given in Table 1. Tensile, CT 12.5 and sub-size Charpy specimens were machined from a 15 mm thick plate of this T91. Subsize Charpy specimens had a KLST type geometry, with notch radius of 100 lm for un-implanted material and 60 lm for implanted specimens (see Fig. 1). This smaller notch radius was chosen in order to ensure that crack initiation would happened in the implanted zone of the specimens. Miniature tensile specimens were obtained from an other T91 heat of almost the same composition. More details about treatments, geometries and preparations can be found in [7]. All the implantations (sub-size Charpy and tensile specimens) were carried out at the cyclotron of the Forschungzentrum Ju¨lich. The implantation conditions have already been described in previous articles [2,7]. Due Table 1 Composition of the T91 steel (in weight %) C
Cr
Mo
V
Nb
Ni
Mn
N
P
Si
0.1
8.73
0.99
0.19
0.031
0.023
0.43
0.029
0.021
0.32
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Fig. 1. Dimensions (in mm) of the sub-size KLST Charpy specimens. Note that for the helium implanted specimens the root notch radius value was 60 lm.
to the thickness of the sub-size Charpy specimens, only a depth of 240 lm (equal to the range of the 34 MeV 3 He particles) was homogeneously implanted with helium, contrary to the miniature tensile specimens, for which totality of the volume was implanted. Three implantation conditions were chosen: 250 C up to a concentration of 0.125 at.% He or of 0.25 at.% He, and 400 C up to a concentration of 0.25 at.% He. For each condition, six specimens were implanted and subsequently shipped to us to be tested. 2.2. Mechanical tests Three-points static bending tests were performed on implanted sub-size Charpy specimens in a glove box in our hot laboratories, because the samples had been activated by implantation. All tests were carried out at room temperature in crosshead displacement control mode at a crosshead speed of 0.1 mm/min. A high frequency data acquisition system also equipped the testing machine in order to allow the detection of possible ‘‘pop-in’’ phenomena inside the implanted zone. The miniature tensile specimens were tested at the Ju¨lich Forschungzentrum. More details about tensile tests can be found in [2]. Tensile, toughness and three-points static bending tests on un-implanted specimens were carried out on a servo-hydraulic machine in a conventional laboratory. These tests were performed at 170 C. The specimens were brought down to 170 C in a chamber mounted around the specimens and acclimatized for about 30 min prior to testing. A more detailed description of mechanical tests can be found in [7]. 2.3. Microstructural analyses The fracture surfaces of some sub-size Charpy specimens were observed by Scanning Electron Microscopy (SEM) on a JEOL 5400. Implantation microstructures were characterized by TEM. Discs of 2 mm diameter were cut from the gauge section of the miniature tensile specimens implanted at 250 C and tested at room temperature, and subsequently electropolished. Observations were carried out on a JEOL 2010-F operated at 200 kV. 3. Results 3.1. Implanted microstructures The preliminary TEM observations on tensile specimens implanted with 0.25 at.% He at 250 C revealed the presence of a high density of small helium bubbles as shown in Fig. 2a. Their density was evaluated to be about 3–4 · 1023 m3. No precise bubble size determination was carried out, however the average detect-
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Fig. 2. TEM micrographs showing the microstructure in mod 9Cr–1Mo steel implanted to 0.25 at.% He at 250 C (a) Bright Field image of the bubble microstructure, underfocus imaging conditions (d f = 2000 nm) and (b) Weak Beam Dark Field image: (g-3g) condition.
able bubble radius appears to be less than 1 nm. Besides, implantation of helium also induced the formation of dislocation loops and small defects clusters (‘‘black dots’’) (Fig. 2a and b). 3.2. Bending tests Thirteen un-implanted sub-size Charpy specimens were submitted to three-points bending tests at 170 C. The fracture surfaces of 6 broken sub-size Charpy specimens were observed by SEM. As expected, the failure mode was predominantly brittle cleavage, even if some intergranular fracture zones were also detected. Among the received implanted specimens, all 6 specimens implanted at 250 C with 0.25 at.% He, five specimens implanted at 400 C with 0.25 at.% He and three at 250 C with 0.125 at.% He were tested in bending. All the implanted specimens tested at room temperature, whatever the implantation conditions, displayed the same behavior. On all the curves, a ‘‘pop-in’’ phenomenon was detected as shown in Fig. 3.
1000
Load (N)
800 600 400 200 0 0
10
20
30
40
50
Opening displacement (µm) Fig. 3. Applied load as a function of opening displacement for T91 implanted at 250 C with 0.25 at.% He tested in three-points bending at room temperature. For all specimens, a ‘‘pop-in’’ (indicated by the arrow) occurred during testing. Although, bending curves corresponding to the other implantation conditions are not shown for clarity purposes, they also displayed the same phenomenon. Finite Element calculated curve for T91 implanted at 400 C with 0.25 at.% He is plotted as well (thick curve).
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This behavior corresponds to the initiation and propagation of a brittle crack in the implanted zone, followed by crack arrest in the un-implanted area of the sample. As can be seen in the SEM micrograph of Fig. 4, the fracture surface consists of two different zones. The first one, just below the notch root, has a fully brittle fracture appearance, mixture of cleavage (Fig. 4b) and intergranular (Fig. 4c) fracture modes. The thickness of this zone is equal to 240 lm, which corresponds exactly to the a particles implantation depth. The second zone presents a fully ductile appearance, as expected for T91 submitted to static bending tests at room temperature. 3.3. Mechanical analysis of the tests Three-dimensional FE calculations of the bending tests were performed to determine the stress and strain fields and the stress at the onset of failure. One quarter of the specimen was meshed as can be seen in Fig. 5, which represents 10,120 elements and 11,511 nodes. The constitutive behaviors used as input data in the FE calculations have been modeled based on the results of tensile tests performed at room temperature or 170 C on cylindrical tensile specimens for the unimplanted material, and at room temperature on miniature tensile specimens for the different implantation conditions. All the curves have been plotted in Fig. 6. For the specimens implanted at 250 C, the fracture in tensile tests happened at yield stress, so the plastic behavior of these materials have been extrapolated by
Fig. 4. SEM micrographs showing the fracture surface of a T91 specimen implanted with 0.25 at.% He after three-points bending test at room temperature. (a) general view of the broken specimen; (b) and (c) details of the fracture surface in the brittle helium implanted zone.
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Fig. 5. Mesh of the bending test. (a) general view of the quarter of the test sample and (b) detail of the notch area, dimension of the mesh in the crack propagation direction in the implanted zone: 25 lm.
Un-implanted, RT (1) Un-implanted, -170˚C (2) 1250 appm / 250˚C, RT - Considere (3) 1250 appm / 250˚C, RT - Rigid-plastic (4) 2500 appm / 250˚C, RT - Considere (5) 2500 appm / 250˚C, RT - Rigid-plastic (6) 2500 appm / 400˚C, RT (7)
True stress (MPa)
2500 2000
(5) (3) (2) (1) (7) (6) (4)
1500 1000 500 0 0
20
40
60 True strain (%)
80
100
Fig. 6. Curves used as input data for the FE calculations.
two models, a rigid-plastic behavior and a ‘‘Considere’’ behavior (necking happens at the yield stress dr/de = ry). In the case of un-implanted T91, only one elastic–plastic constitutive behavior (obtained from the tensile tests at 170 C) was used as input data, whereas two constitutive behaviors were used for the implanted specimens, one for calculations in the un-implanted zone (obtained from the tensile tests at room temperature on un-implanted specimen) and the other for calculations in the helium implanted area (obtained from the tensile tests at room temperature on implanted T91). The applied load as a function of opening displacement for the bending tests on implanted and un-implanted material was calculated and good agreement with experimental data was found, at least up to load at which the ‘‘pop-in’’ occurred for the implanted material as can be seen in Fig. 3. For the materials implanted at 250 C, the two models (rigid-plastic and ‘‘Considere’’) give almost the same results.
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Table 2 Results of the 3D FE calculations; N is the number of specimens tested [He] (appm) 0 0 1250 2300 2500
Timpl (C)
ry
rfmin
rfmax
r f
reqmin/ry
reqmax/ry
N
250 250 400
501 (RT) 815 (170 C) 1003 1090 737
1870 1690 1660 1490
2180 2020 1760 1790
2025 1855 1710 1640
1.33 1.02 1.01 1.15
1.78 1.04 1.01 1.27
13 3 6 5
2500
RT
σ (MPa)
2000
σf
-170˚C
1500
Timpl = 250˚C σy at RT
1000 500
Timpl = 400˚C 0 0
500
1000
1500
2000
2500
3000
Helium content (appm) f as a function of helium content. Fig. 7. Evolution of yield stress ry and average maximum principal stress at fracture r
The main results obtained from these 3D FE calculations have been gathered in Table 2. For the maximum principal stress r1, the maximum stress values, calculated at the onset of the fracture, have been reported in this table and referred to as rf. In each case a minimum (rfmin) and a maximum (rfmax) value corresponding, respectively to the lowest and highest values of the opening displacement at failure (or at the onset of the ‘‘pop-in’’) are indicated together with the average value ð rf Þ. As expected, the maximum stress values were located in the implanted zone on the plan at the notch root referred as A in Fig. 5. Calculated values indicated that failure happened at lower r1 values for implanted specimens tested at room temperature than for unimplanted samples tested at 170 C. The ratio of the equivalent von Mises stress (req) to the yield stress (ry), which gives a qualitative indication of the local amount of plastic deformation, is also indicated (as in the case of the maximum principal stress at fracture, reqmin and reqmax were calculated from the lowest and the highest values of the opening displacement at the onset of the ‘‘pop-in’’). Obtained values for this ratio indicated that there was very little plastic deformation in specimens implanted at 250 C, contrary to un-implanted specimens and samples implanted at 400 C. In Fig. 7 are reported the evolution of measured yield stress and calculated maximum principal stress at rupture as a function of helium content. The two ry curves show the hardening associated with helium implantation, whereas the other curve illustrates the embrittlement effect. 4. Discussion The experiments performed in this study clearly show a drastic embrittlement of T91 steel by the implanted helium. Indeed, this helium implantation leads to a decrease of the value of the brittle fracture stress, since for the same geometry, the brittle fracture occurs for un-implanted specimens for values of the maximum principal stress r1, calculated from three-points static bending tests at 170 C, in the range 1870–2180 MPa, whereas for implanted specimens tested at room temperature, the values of r1 are almost all below
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1850 MPa. Obviously, the implanted helium lowers the grain boundary cohesion, since there is a significant amount of intergranular fracture zones (40–50% of the brittle fracture area) on the fracture surface of implanted specimens, whereas fracture mode on un-implanted T91 is fully cleavage. These results, particularly the notable amount of intergranular fracture areas, seem to indicate that a ‘‘Non-Hardening Embrittlement’’ (NHE) effect is added to the Hardening Embrittlement effect linked to the implantation, i.e. the increase of the yield stress as can be seen in Fig. 7 [4]. Yamato and co-workers [8] have analyzed the published data sets on irradiation-induced embrittlement. On the one hand, results suggest that up to concentrations several hundred appm (400–600 appm), embrittlement is dominated by a hardening mechanism at irradiation temperature below about 400 C. On the other hand, for higher concentrations, the scarce data [9,10] are consistent with the hypothesis that accumulation of helium sufficiently weakens grain boundaries to the point where a NHE effect emerges as signaled by intergranular fracture. In addition, the FE calculations indicate that for the specimens implanted with helium at 250 C, ‘‘pop-in’’ crack propagation can be triggered only if some plastic deformation is introduced in the material (ratio von Mises stress to yield stress just above 1). This behavior can be summarized by the following criterion: failure happens if (req P ry), which could mean that flaws or microcracks have to be created by plasticity before failure. This suggests that a limiting step in the fracture mechanism is the nucleation of a crack. On the contrary, for un-implanted specimens and specimens implanted at 400 C, the apparition of plasticity is not the limiting step. In this case, the criterion of failure is (r1 P rc), where rc is the critical fracture stress. The critical step in the fracture process is here the propagation of the microcracks. Hence it is proposed that the brittle failure of the specimens is controlled by a double criterion depending on helium content and implantation temperature. This double criterion has already been proposed for other materials [11,12]. In the cases of the crack propagation controlled mechanism, for which the critical fracture stress can be measured, the values of rc are equal to the rf values calculated by FE. It concerns the un-implanted T91 and the material implanted at 400 C with 0.25 at.% He. The experimental associated failure probabilities, Pr, were determined according to the following equation: Pr ¼
ði 0:5Þ N
ð1Þ
where N stands for the number of experiments and i for the rank of a given result (the experiments were ranked from 1 to N based on increasing opening displacement). The Beremin model of brittle fracture was then applied to these two materials. In this model, the failure probability is described by a Weibull distribution, which can be written as follows: m rw P r ¼ 1 exp ð2Þ ru where m and ru are material constants characterizing the fracture behavior. The Weibull stress, rw, is defined as: Z 1=m m dV rw ¼ rc ð3Þ V0 VP with Vp the volume plastically deformed and V0 a representative elementary volume, chosen as 25 lm3. The parameters m and ru of this model were adjusted to fit to the experimental results. Simulated curves and experimental results of the un-implanted material and implanted T91 at 400 C are plotted in Fig. 8. The obtained values of the fitting parameters are gathered in Table 3, as well as 95% confidence intervals values, calculated according to [13]. The m values of the two cases are almost equal, which means that the scattering effect is the same. On the contrary, a significant effect of helium implantation is noticeable on ru values. From the m and ru material parameters values, the Weibull stress rw, and then the failure probability, can be computed by FE simulations for a given mechanical test and specimen geometry. This was done, based on the un-implanted material parameters, in the case of toughness test on CT 12.5 specimens. From a CT 12.5 specimen mesh, the stress and strain fields have been determined and then the stress intensity factors have been calculated and the failure probabilities have been computed by applying formulae (2) and (3) with
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Failure probability Pr
1.0 0.8 0.6 0.4
2500 appm / 400˚C - Experiment 2500 appm / 400˚C - Simulation
0.2
Un-implanted - Experiment Un-implanted - Simulation
0.0 0.00
0.05
0.10
0.15
0.20
0.25
0.30
Opening displacement (mm) Fig. 8. Experimental (dots) and calculated (lines) failure probabilities as a function of opening displacement for T91 tested in three-points bending (empty dots for un-implanted specimens and full dots for specimens implanted at 400 C).
Table 3 Values of the fracture parameters m and ru for T91 un-implanted and 0.25 at.% He implanted at 400 C, as defined in the Beremin model, þ þ identified using static bending tests (respectively at 170 C and room temperature). Confidence intervals ðr u ; ru ; m ; m Þ corresponding to a confidence level of 95% are indicated as well
Un-implanted Timpl = 400 C
ru (MPa)
r u
rþ u
m
m
m+
3997 2433
3787 2186
4221 2716
13.2 14.8
7.2 4.2
18.8 24.5
the m and ru values obtained from the charpy tests. Experimental and calculated curves at 170 C are plotted in Fig. 9. There is a small shift between simulated and experimental results. However, only 4 tests were carried out. Likewise, the failure probability for a toughness test on CT 12.5 specimen with m and ru parameters values determined for the material implanted at 400 C was computed. The obtained curves are also plotted in Fig. 9. 1.0
Implanted
Failure probability Pr
0.9
Un-implanted
0.8 0.7 0.6 Experiment, -170˚C Pr inf, -170˚C Pr mean, -170˚C Pr sup, -170˚C Pr inf, impl, RT Pr mean, impl, RT Pr sup, impl, RT
0.5 0.4 0.3 0.2 0.1 0.0 0
10
20
30
40
50
60
70
Kj (MPa.m 0.5) Fig. 9. Experimental and calculated failure probabilities as a function of stress intensity factor for T91 CT specimens tested at 170 C. Failure probabilities calculated for T91 CT implanted with 0.25 at.% He at 400 C and tested at room temperature are also indicated. Pr þ inf and Pr sup correspond to curves calculated using ðm; r u Þ resp. ðm; ru Þ values reported in Table 3. Pr mean was calculated using the optimum value (m, ru).
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From the calculated failure probabilities, the mean toughness for the un-implanted material tested at 170 C and for the material implanted at 400 C and tested at room temperature are deduced. These values are 30 MPa Æ m0.5 and 16 MPa Æ m0.5, respectively. 5. Summary/Conclusion The effects of helium implantation on the fracture properties of a mod-9Cr–1Mo martensitic steel were studied by three-points static bending tests on implanted and un-implanted specimens, performed at room temperature and 170 C, respectively. Results were analyzed by FE coupled with SEM observations of fracture surfaces. The main results are summarized below: • A ‘‘pop-in’’ phenomenon was systematically noted during the bending tests at room temperature on implanted samples. The SEM observations of their fracture surfaces revealed that the failure mode in the implanted zone was fully brittle with both cleavage and intergranular fracture (and fully ductile in the un-implanted zone). By contrast, the fracture mode for pristine T91 at 170 C was cleavage. • FE analysis allowed to calculate stress and strain fields and to determine r1. These results indicated that brittle fracture occurred in the implanted specimens at lower r1 than in the case of un-implanted samples. • Both FE results and SEM observations demonstrate that the implanted helium has induced a decrease in the critical stress for intergranular fracture. This embrittlement is linked to the implantation-induced microstructure evolution and in particular to the accumulation of helium at grain boundary, which could lower grain boundary cohesion. • Moreover, based on the FE analyses it was proposed that brittle fracture mechanism is controlled by a double criteria depending on implantation temperature and helium content: a nucleation criterion if (req P ry) and a propagation criterion if (r1 P rc). For our experiments, the mechanism is controlled by the nucleation criterion at 250 C, whatever the helium content, and for the un-implanted material and the implanted one with 0.25 at.% He at 400 C, governing criterion is propagation. • Finally, the application of the Beremin model in the cases of brittle fracture mechanism controlled by propagation criterion allowed to determine m and ru parameters. Using the obtained values for m and ru, the evolution of toughness with helium has been evaluated. The toughness of the T91 steel at room temperature decreases from more than 200 MPa Æ m0.5 according to the literature to about 16 MPa Æ m0.5 after implantation with 0.25 at.% He at 400 C, which shows the detrimental effect of helium implantation. Acknowledgements Financial support from the EUROTRANS project (6th EU Framework Programme) is gratefully acknowledged. References [1] Schroeder H, Ullmaier H. Helium and hydrogen effects on the embrittlement of iron- and nickel-based alloys. J Nucl Mater 1991;179– 181:118–24. [2] Jung P, Henry J, Chen J, Brachet J-C. Effect of implanted helium on tensile properties and hardness of 9%Cr martensitic stainless steels. J Nucl Mater 2003;318:241–8. [3] Henry J, Mathon M-H, Jung P. Microstructural analysis of 9%Cr martensitic steels containing 05. at.% helium. J Nucl Mater 2003;318:249–59. [4] Odette GR, Yamamoto T, Kishimoto H. Fusion materials semiannual report, 2004. [5] Klueh R, Harries D. High-Chromium ferritic and martensitic steels for nuclear applications. ASTM Monogr Mono3 2001:156–62. [6] Lindau R, Mo¨slang A, Preininger D, Rieth M, Ro¨hrig HD. Influence of helium on impact properties of reduced-activation ferritic/ martensitic Cr-steels. J Nucl Mater 1999;271–272:450–4. [7] Henry J, Vincent L, Averty X, Marini B, Jung P. Bending tests on T91 samples implanted with 0.25at.% helium : experiments and mechanical analysis. J Nucl Mater 2006;356:78–87. [8] Yamato T, Odette GR, Kishimoto H, Rensman J-W, Miao P. On the effects of irradiation and helium on the yield stress changes and hardening and non-hardening embrittlement of 8Cr tempered martensitic steels: compilation and analysis of existing data. J Nucl Mater 2006;356:27–49.
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[9] Dai Y, Jia XJ, Farrell K. Mechanical properties of modified 9Cr–1Mo (T91) irradiated at 6300 C in SINQ Target-3. J Nucl Mater 2003;318:192–9. [10] Jia X, Dai Y. Small punch tests on martensitic/ferritic steels F82H, T91 and Optimax-A irradiated in SINQ Target-3. J Nucl Mater 2003;323:360–7. [11] Tetelman, Mc Evily. Fracture of structural materials. John Wiley and Sons, Inc., 1967. p. 234–84. [12] Pineau A. Advances in Fracture Research, vol. 2. Pergamon Press; 1981. p. 553–77. [13] ESIS Procedure P6-94. Draft procedure to measure and calculate material parameters for the local approach to fracture using notched tensile specimens.
Engineering Fracture Mechanics 75 (2008) 3570–3580 www.elsevier.com/locate/engfracmech
Characterization of He embrittlement of a 9Cr–1Mo steel using local approach of brittle fracture J. Malaplate a b
a,*
, L. Vincent a, X. Averty b, J. Henry a, B. Marini
a
CEA Saclay, DEN-DANS/DMN/SRMA 91191, Gif-sur-Yvette Cedex, France CEA Saclay, DEN-DANS/DMN/SEMI 91191, Gif-sur-Yvette Cedex, France
Received 22 September 2006; received in revised form 22 December 2006; accepted 12 February 2007 Available online 27 February 2007
Abstract Ferritic-martensitic steels are prime candidate materials for future reactors. We present here the results of a study on the effects of helium implantation on the fracture behavior of 9Cr (T91) martensitic steels. Three-points static bending tests were performed at room temperature on implanted specimens and at 170 C on un-implanted material. All these tests led to brittle fracture. Based on a mechanical analysis of the tests results using Finite Element calculations, we have proposed that the mechanism of brittle fracture is controlled by a double criterion depending on implantation temperature and helium content. Furthermore, by applying the Beremin model, the toughness of helium implanted steel has been evaluated. 2007 Elsevier Ltd. All rights reserved. Keywords: Helium embrittlement; 9Cr martensitic steels; Brittle fracture; Bending tests; Finite element analysis
1. Introduction Ferritic/martensitic steels are prime candidate materials for future fusion reactors and spallation neutron sources due to their excellent thermo-mechanical properties and radiation resistance at temperature above 400 C in a fission spectrum. However, high helium contents will be produced in these materials during operation, which can have a detrimental effect on mechanical properties. Up to now, it has been well established that ferritic/martensitic steels are relatively immune to the so called ‘‘high temperature’’ helium embrittlement phenomenon (see for instance [1]). As regards helium effects at lower temperature on hardening and embrittlement, data are scarce and this topic is still a matter of controversy (see for instance [2,3] for tensile properties and [4–6] for fracture properties). Therefore, the aim of this work is to advance the characterization and understanding of intrinsic effect of helium at low temperature on fracture properties of 9Cr martensitic steel. To this end, three-points static bending tests have been performed at room temperature on 9Cr–1Mo–V–Nb martensitic steel (commercially called
*
Corresponding author. E-mail address: [email protected] (J. Malaplate).
0013-7944/$ - see front matter 2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.engfracmech.2007.02.019
J. Malaplate et al. / Engineering Fracture Mechanics 75 (2008) 3570–3580
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Nomenclature ry req reqmin reqmax r1 rf rfmin rfmax f r rc C r Pr ru þ r u ; ru m m, m+ rw Vp V0 Kj
yield stress equivalent von Mises stress equivalent von Mises stress for the lowest value of the opening displacement equivalent von Mises stress for the highest value of the opening displacement maximum principal stress maximum stress at fracture (‘‘pop-in’’) maximum stress at ‘‘pop-in’’ for the lowest value of the opening displacement maximum stress at ‘‘pop-in’’ for the highest value of the opening displacement average maximum stress at fracture critical fracture stress average critical fracture stress failure probability material constant defined in the Beremin model confidence interval of ru Weibull modulus confidence interval of m Weibull stress plastic volume representative elementary volume stress intensity factor
T91) implanted with helium at different contents and/or temperatures. Bending tests have also been performed on pristine materials at low temperature. The goal was to compare the fracture properties of T91 loaded with helium to those of un-implanted T91 in the brittle domain. Results of tests have been analyzed by Finite Element (FE). A preliminary microstructural characterization has also been carried out on miniature tensile specimens by Transmission Electron Microscopy (TEM). 2. Experimental 2.1. Materials, specimens and helium implantation The studied steel is a mod-9Cr–1Mo steel, i.e. 9Cr–1Mo–V–Nb commercially referred to as T91, received in the normalized and tempered metallurgical condition (solution annealed for 1 h at 1040 C, fast cool, reheated 1 h at 760 C, fast cool). The chemical composition is given in Table 1. Tensile, CT 12.5 and sub-size Charpy specimens were machined from a 15 mm thick plate of this T91. Subsize Charpy specimens had a KLST type geometry, with notch radius of 100 lm for un-implanted material and 60 lm for implanted specimens (see Fig. 1). This smaller notch radius was chosen in order to ensure that crack initiation would happened in the implanted zone of the specimens. Miniature tensile specimens were obtained from an other T91 heat of almost the same composition. More details about treatments, geometries and preparations can be found in [7]. All the implantations (sub-size Charpy and tensile specimens) were carried out at the cyclotron of the Forschungzentrum Ju¨lich. The implantation conditions have already been described in previous articles [2,7]. Due Table 1 Composition of the T91 steel (in weight %) C
Cr
Mo
V
Nb
Ni
Mn
N
P
Si
0.1
8.73
0.99
0.19
0.031
0.023
0.43
0.029
0.021
0.32
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Fig. 1. Dimensions (in mm) of the sub-size KLST Charpy specimens. Note that for the helium implanted specimens the root notch radius value was 60 lm.
to the thickness of the sub-size Charpy specimens, only a depth of 240 lm (equal to the range of the 34 MeV 3 He particles) was homogeneously implanted with helium, contrary to the miniature tensile specimens, for which totality of the volume was implanted. Three implantation conditions were chosen: 250 C up to a concentration of 0.125 at.% He or of 0.25 at.% He, and 400 C up to a concentration of 0.25 at.% He. For each condition, six specimens were implanted and subsequently shipped to us to be tested. 2.2. Mechanical tests Three-points static bending tests were performed on implanted sub-size Charpy specimens in a glove box in our hot laboratories, because the samples had been activated by implantation. All tests were carried out at room temperature in crosshead displacement control mode at a crosshead speed of 0.1 mm/min. A high frequency data acquisition system also equipped the testing machine in order to allow the detection of possible ‘‘pop-in’’ phenomena inside the implanted zone. The miniature tensile specimens were tested at the Ju¨lich Forschungzentrum. More details about tensile tests can be found in [2]. Tensile, toughness and three-points static bending tests on un-implanted specimens were carried out on a servo-hydraulic machine in a conventional laboratory. These tests were performed at 170 C. The specimens were brought down to 170 C in a chamber mounted around the specimens and acclimatized for about 30 min prior to testing. A more detailed description of mechanical tests can be found in [7]. 2.3. Microstructural analyses The fracture surfaces of some sub-size Charpy specimens were observed by Scanning Electron Microscopy (SEM) on a JEOL 5400. Implantation microstructures were characterized by TEM. Discs of 2 mm diameter were cut from the gauge section of the miniature tensile specimens implanted at 250 C and tested at room temperature, and subsequently electropolished. Observations were carried out on a JEOL 2010-F operated at 200 kV. 3. Results 3.1. Implanted microstructures The preliminary TEM observations on tensile specimens implanted with 0.25 at.% He at 250 C revealed the presence of a high density of small helium bubbles as shown in Fig. 2a. Their density was evaluated to be about 3–4 · 1023 m3. No precise bubble size determination was carried out, however the average detect-
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Fig. 2. TEM micrographs showing the microstructure in mod 9Cr–1Mo steel implanted to 0.25 at.% He at 250 C (a) Bright Field image of the bubble microstructure, underfocus imaging conditions (d f = 2000 nm) and (b) Weak Beam Dark Field image: (g-3g) condition.
able bubble radius appears to be less than 1 nm. Besides, implantation of helium also induced the formation of dislocation loops and small defects clusters (‘‘black dots’’) (Fig. 2a and b). 3.2. Bending tests Thirteen un-implanted sub-size Charpy specimens were submitted to three-points bending tests at 170 C. The fracture surfaces of 6 broken sub-size Charpy specimens were observed by SEM. As expected, the failure mode was predominantly brittle cleavage, even if some intergranular fracture zones were also detected. Among the received implanted specimens, all 6 specimens implanted at 250 C with 0.25 at.% He, five specimens implanted at 400 C with 0.25 at.% He and three at 250 C with 0.125 at.% He were tested in bending. All the implanted specimens tested at room temperature, whatever the implantation conditions, displayed the same behavior. On all the curves, a ‘‘pop-in’’ phenomenon was detected as shown in Fig. 3.
1000
Load (N)
800 600 400 200 0 0
10
20
30
40
50
Opening displacement (µm) Fig. 3. Applied load as a function of opening displacement for T91 implanted at 250 C with 0.25 at.% He tested in three-points bending at room temperature. For all specimens, a ‘‘pop-in’’ (indicated by the arrow) occurred during testing. Although, bending curves corresponding to the other implantation conditions are not shown for clarity purposes, they also displayed the same phenomenon. Finite Element calculated curve for T91 implanted at 400 C with 0.25 at.% He is plotted as well (thick curve).
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This behavior corresponds to the initiation and propagation of a brittle crack in the implanted zone, followed by crack arrest in the un-implanted area of the sample. As can be seen in the SEM micrograph of Fig. 4, the fracture surface consists of two different zones. The first one, just below the notch root, has a fully brittle fracture appearance, mixture of cleavage (Fig. 4b) and intergranular (Fig. 4c) fracture modes. The thickness of this zone is equal to 240 lm, which corresponds exactly to the a particles implantation depth. The second zone presents a fully ductile appearance, as expected for T91 submitted to static bending tests at room temperature. 3.3. Mechanical analysis of the tests Three-dimensional FE calculations of the bending tests were performed to determine the stress and strain fields and the stress at the onset of failure. One quarter of the specimen was meshed as can be seen in Fig. 5, which represents 10,120 elements and 11,511 nodes. The constitutive behaviors used as input data in the FE calculations have been modeled based on the results of tensile tests performed at room temperature or 170 C on cylindrical tensile specimens for the unimplanted material, and at room temperature on miniature tensile specimens for the different implantation conditions. All the curves have been plotted in Fig. 6. For the specimens implanted at 250 C, the fracture in tensile tests happened at yield stress, so the plastic behavior of these materials have been extrapolated by
Fig. 4. SEM micrographs showing the fracture surface of a T91 specimen implanted with 0.25 at.% He after three-points bending test at room temperature. (a) general view of the broken specimen; (b) and (c) details of the fracture surface in the brittle helium implanted zone.
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Fig. 5. Mesh of the bending test. (a) general view of the quarter of the test sample and (b) detail of the notch area, dimension of the mesh in the crack propagation direction in the implanted zone: 25 lm.
Un-implanted, RT (1) Un-implanted, -170˚C (2) 1250 appm / 250˚C, RT - Considere (3) 1250 appm / 250˚C, RT - Rigid-plastic (4) 2500 appm / 250˚C, RT - Considere (5) 2500 appm / 250˚C, RT - Rigid-plastic (6) 2500 appm / 400˚C, RT (7)
True stress (MPa)
2500 2000
(5) (3) (2) (1) (7) (6) (4)
1500 1000 500 0 0
20
40
60 True strain (%)
80
100
Fig. 6. Curves used as input data for the FE calculations.
two models, a rigid-plastic behavior and a ‘‘Considere’’ behavior (necking happens at the yield stress dr/de = ry). In the case of un-implanted T91, only one elastic–plastic constitutive behavior (obtained from the tensile tests at 170 C) was used as input data, whereas two constitutive behaviors were used for the implanted specimens, one for calculations in the un-implanted zone (obtained from the tensile tests at room temperature on un-implanted specimen) and the other for calculations in the helium implanted area (obtained from the tensile tests at room temperature on implanted T91). The applied load as a function of opening displacement for the bending tests on implanted and un-implanted material was calculated and good agreement with experimental data was found, at least up to load at which the ‘‘pop-in’’ occurred for the implanted material as can be seen in Fig. 3. For the materials implanted at 250 C, the two models (rigid-plastic and ‘‘Considere’’) give almost the same results.
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Table 2 Results of the 3D FE calculations; N is the number of specimens tested [He] (appm) 0 0 1250 2300 2500
Timpl (C)
ry
rfmin
rfmax
r f
reqmin/ry
reqmax/ry
N
250 250 400
501 (RT) 815 (170 C) 1003 1090 737
1870 1690 1660 1490
2180 2020 1760 1790
2025 1855 1710 1640
1.33 1.02 1.01 1.15
1.78 1.04 1.01 1.27
13 3 6 5
2500
RT
σ (MPa)
2000
σf
-170˚C
1500
Timpl = 250˚C σy at RT
1000 500
Timpl = 400˚C 0 0
500
1000
1500
2000
2500
3000
Helium content (appm) f as a function of helium content. Fig. 7. Evolution of yield stress ry and average maximum principal stress at fracture r
The main results obtained from these 3D FE calculations have been gathered in Table 2. For the maximum principal stress r1, the maximum stress values, calculated at the onset of the fracture, have been reported in this table and referred to as rf. In each case a minimum (rfmin) and a maximum (rfmax) value corresponding, respectively to the lowest and highest values of the opening displacement at failure (or at the onset of the ‘‘pop-in’’) are indicated together with the average value ð rf Þ. As expected, the maximum stress values were located in the implanted zone on the plan at the notch root referred as A in Fig. 5. Calculated values indicated that failure happened at lower r1 values for implanted specimens tested at room temperature than for unimplanted samples tested at 170 C. The ratio of the equivalent von Mises stress (req) to the yield stress (ry), which gives a qualitative indication of the local amount of plastic deformation, is also indicated (as in the case of the maximum principal stress at fracture, reqmin and reqmax were calculated from the lowest and the highest values of the opening displacement at the onset of the ‘‘pop-in’’). Obtained values for this ratio indicated that there was very little plastic deformation in specimens implanted at 250 C, contrary to un-implanted specimens and samples implanted at 400 C. In Fig. 7 are reported the evolution of measured yield stress and calculated maximum principal stress at rupture as a function of helium content. The two ry curves show the hardening associated with helium implantation, whereas the other curve illustrates the embrittlement effect. 4. Discussion The experiments performed in this study clearly show a drastic embrittlement of T91 steel by the implanted helium. Indeed, this helium implantation leads to a decrease of the value of the brittle fracture stress, since for the same geometry, the brittle fracture occurs for un-implanted specimens for values of the maximum principal stress r1, calculated from three-points static bending tests at 170 C, in the range 1870–2180 MPa, whereas for implanted specimens tested at room temperature, the values of r1 are almost all below
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1850 MPa. Obviously, the implanted helium lowers the grain boundary cohesion, since there is a significant amount of intergranular fracture zones (40–50% of the brittle fracture area) on the fracture surface of implanted specimens, whereas fracture mode on un-implanted T91 is fully cleavage. These results, particularly the notable amount of intergranular fracture areas, seem to indicate that a ‘‘Non-Hardening Embrittlement’’ (NHE) effect is added to the Hardening Embrittlement effect linked to the implantation, i.e. the increase of the yield stress as can be seen in Fig. 7 [4]. Yamato and co-workers [8] have analyzed the published data sets on irradiation-induced embrittlement. On the one hand, results suggest that up to concentrations several hundred appm (400–600 appm), embrittlement is dominated by a hardening mechanism at irradiation temperature below about 400 C. On the other hand, for higher concentrations, the scarce data [9,10] are consistent with the hypothesis that accumulation of helium sufficiently weakens grain boundaries to the point where a NHE effect emerges as signaled by intergranular fracture. In addition, the FE calculations indicate that for the specimens implanted with helium at 250 C, ‘‘pop-in’’ crack propagation can be triggered only if some plastic deformation is introduced in the material (ratio von Mises stress to yield stress just above 1). This behavior can be summarized by the following criterion: failure happens if (req P ry), which could mean that flaws or microcracks have to be created by plasticity before failure. This suggests that a limiting step in the fracture mechanism is the nucleation of a crack. On the contrary, for un-implanted specimens and specimens implanted at 400 C, the apparition of plasticity is not the limiting step. In this case, the criterion of failure is (r1 P rc), where rc is the critical fracture stress. The critical step in the fracture process is here the propagation of the microcracks. Hence it is proposed that the brittle failure of the specimens is controlled by a double criterion depending on helium content and implantation temperature. This double criterion has already been proposed for other materials [11,12]. In the cases of the crack propagation controlled mechanism, for which the critical fracture stress can be measured, the values of rc are equal to the rf values calculated by FE. It concerns the un-implanted T91 and the material implanted at 400 C with 0.25 at.% He. The experimental associated failure probabilities, Pr, were determined according to the following equation: Pr ¼
ði 0:5Þ N
ð1Þ
where N stands for the number of experiments and i for the rank of a given result (the experiments were ranked from 1 to N based on increasing opening displacement). The Beremin model of brittle fracture was then applied to these two materials. In this model, the failure probability is described by a Weibull distribution, which can be written as follows: m rw P r ¼ 1 exp ð2Þ ru where m and ru are material constants characterizing the fracture behavior. The Weibull stress, rw, is defined as: Z 1=m m dV rw ¼ rc ð3Þ V0 VP with Vp the volume plastically deformed and V0 a representative elementary volume, chosen as 25 lm3. The parameters m and ru of this model were adjusted to fit to the experimental results. Simulated curves and experimental results of the un-implanted material and implanted T91 at 400 C are plotted in Fig. 8. The obtained values of the fitting parameters are gathered in Table 3, as well as 95% confidence intervals values, calculated according to [13]. The m values of the two cases are almost equal, which means that the scattering effect is the same. On the contrary, a significant effect of helium implantation is noticeable on ru values. From the m and ru material parameters values, the Weibull stress rw, and then the failure probability, can be computed by FE simulations for a given mechanical test and specimen geometry. This was done, based on the un-implanted material parameters, in the case of toughness test on CT 12.5 specimens. From a CT 12.5 specimen mesh, the stress and strain fields have been determined and then the stress intensity factors have been calculated and the failure probabilities have been computed by applying formulae (2) and (3) with
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Failure probability Pr
1.0 0.8 0.6 0.4
2500 appm / 400˚C - Experiment 2500 appm / 400˚C - Simulation
0.2
Un-implanted - Experiment Un-implanted - Simulation
0.0 0.00
0.05
0.10
0.15
0.20
0.25
0.30
Opening displacement (mm) Fig. 8. Experimental (dots) and calculated (lines) failure probabilities as a function of opening displacement for T91 tested in three-points bending (empty dots for un-implanted specimens and full dots for specimens implanted at 400 C).
Table 3 Values of the fracture parameters m and ru for T91 un-implanted and 0.25 at.% He implanted at 400 C, as defined in the Beremin model, þ þ identified using static bending tests (respectively at 170 C and room temperature). Confidence intervals ðr u ; ru ; m ; m Þ corresponding to a confidence level of 95% are indicated as well
Un-implanted Timpl = 400 C
ru (MPa)
r u
rþ u
m
m
m+
3997 2433
3787 2186
4221 2716
13.2 14.8
7.2 4.2
18.8 24.5
the m and ru values obtained from the charpy tests. Experimental and calculated curves at 170 C are plotted in Fig. 9. There is a small shift between simulated and experimental results. However, only 4 tests were carried out. Likewise, the failure probability for a toughness test on CT 12.5 specimen with m and ru parameters values determined for the material implanted at 400 C was computed. The obtained curves are also plotted in Fig. 9. 1.0
Implanted
Failure probability Pr
0.9
Un-implanted
0.8 0.7 0.6 Experiment, -170˚C Pr inf, -170˚C Pr mean, -170˚C Pr sup, -170˚C Pr inf, impl, RT Pr mean, impl, RT Pr sup, impl, RT
0.5 0.4 0.3 0.2 0.1 0.0 0
10
20
30
40
50
60
70
Kj (MPa.m 0.5) Fig. 9. Experimental and calculated failure probabilities as a function of stress intensity factor for T91 CT specimens tested at 170 C. Failure probabilities calculated for T91 CT implanted with 0.25 at.% He at 400 C and tested at room temperature are also indicated. Pr þ inf and Pr sup correspond to curves calculated using ðm; r u Þ resp. ðm; ru Þ values reported in Table 3. Pr mean was calculated using the optimum value (m, ru).
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From the calculated failure probabilities, the mean toughness for the un-implanted material tested at 170 C and for the material implanted at 400 C and tested at room temperature are deduced. These values are 30 MPa Æ m0.5 and 16 MPa Æ m0.5, respectively. 5. Summary/Conclusion The effects of helium implantation on the fracture properties of a mod-9Cr–1Mo martensitic steel were studied by three-points static bending tests on implanted and un-implanted specimens, performed at room temperature and 170 C, respectively. Results were analyzed by FE coupled with SEM observations of fracture surfaces. The main results are summarized below: • A ‘‘pop-in’’ phenomenon was systematically noted during the bending tests at room temperature on implanted samples. The SEM observations of their fracture surfaces revealed that the failure mode in the implanted zone was fully brittle with both cleavage and intergranular fracture (and fully ductile in the un-implanted zone). By contrast, the fracture mode for pristine T91 at 170 C was cleavage. • FE analysis allowed to calculate stress and strain fields and to determine r1. These results indicated that brittle fracture occurred in the implanted specimens at lower r1 than in the case of un-implanted samples. • Both FE results and SEM observations demonstrate that the implanted helium has induced a decrease in the critical stress for intergranular fracture. This embrittlement is linked to the implantation-induced microstructure evolution and in particular to the accumulation of helium at grain boundary, which could lower grain boundary cohesion. • Moreover, based on the FE analyses it was proposed that brittle fracture mechanism is controlled by a double criteria depending on implantation temperature and helium content: a nucleation criterion if (req P ry) and a propagation criterion if (r1 P rc). For our experiments, the mechanism is controlled by the nucleation criterion at 250 C, whatever the helium content, and for the un-implanted material and the implanted one with 0.25 at.% He at 400 C, governing criterion is propagation. • Finally, the application of the Beremin model in the cases of brittle fracture mechanism controlled by propagation criterion allowed to determine m and ru parameters. Using the obtained values for m and ru, the evolution of toughness with helium has been evaluated. The toughness of the T91 steel at room temperature decreases from more than 200 MPa Æ m0.5 according to the literature to about 16 MPa Æ m0.5 after implantation with 0.25 at.% He at 400 C, which shows the detrimental effect of helium implantation. Acknowledgements Financial support from the EUROTRANS project (6th EU Framework Programme) is gratefully acknowledged. References [1] Schroeder H, Ullmaier H. Helium and hydrogen effects on the embrittlement of iron- and nickel-based alloys. J Nucl Mater 1991;179– 181:118–24. [2] Jung P, Henry J, Chen J, Brachet J-C. Effect of implanted helium on tensile properties and hardness of 9%Cr martensitic stainless steels. J Nucl Mater 2003;318:241–8. [3] Henry J, Mathon M-H, Jung P. Microstructural analysis of 9%Cr martensitic steels containing 05. at.% helium. J Nucl Mater 2003;318:249–59. [4] Odette GR, Yamamoto T, Kishimoto H. Fusion materials semiannual report, 2004. [5] Klueh R, Harries D. High-Chromium ferritic and martensitic steels for nuclear applications. ASTM Monogr Mono3 2001:156–62. [6] Lindau R, Mo¨slang A, Preininger D, Rieth M, Ro¨hrig HD. Influence of helium on impact properties of reduced-activation ferritic/ martensitic Cr-steels. J Nucl Mater 1999;271–272:450–4. [7] Henry J, Vincent L, Averty X, Marini B, Jung P. Bending tests on T91 samples implanted with 0.25at.% helium : experiments and mechanical analysis. J Nucl Mater 2006;356:78–87. [8] Yamato T, Odette GR, Kishimoto H, Rensman J-W, Miao P. On the effects of irradiation and helium on the yield stress changes and hardening and non-hardening embrittlement of 8Cr tempered martensitic steels: compilation and analysis of existing data. J Nucl Mater 2006;356:27–49.
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[9] Dai Y, Jia XJ, Farrell K. Mechanical properties of modified 9Cr–1Mo (T91) irradiated at 6300 C in SINQ Target-3. J Nucl Mater 2003;318:192–9. [10] Jia X, Dai Y. Small punch tests on martensitic/ferritic steels F82H, T91 and Optimax-A irradiated in SINQ Target-3. J Nucl Mater 2003;323:360–7. [11] Tetelman, Mc Evily. Fracture of structural materials. John Wiley and Sons, Inc., 1967. p. 234–84. [12] Pineau A. Advances in Fracture Research, vol. 2. Pergamon Press; 1981. p. 553–77. [13] ESIS Procedure P6-94. Draft procedure to measure and calculate material parameters for the local approach to fracture using notched tensile specimens.