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Fracture toughness of irradiated and recovered vessel steels
Nuclear Engineering and Design 182 (1998) 131 – 140
Fracture toughness of irradiated and recovered vessel steels F. Perosanz a, A. Valiente b,*, J. Lapen˜a a b
a Energy, En6ironment and Technology Research Center (CIEMAT), A6. Complutense 2, 28040 -Madrid, Spain Materials Science Department, Polytechnic Uni6ersity of Madrid, E.T.S. Ing. Caminos, Ciudad Uni6ersitaria s/n, 28040 -Madrid, Spain
Abstract This paper presents the fracture toughness measurements carried out on three vessel steels in an irradiated condition and after a post-irradiation recovery treatment. A statistical approach and the fracture parameters corresponding to two theoretical models of the fracture tests are used for evaluating toughness. Test results show that the neutron fluence gradually transforms the fracture behaviour of the vessel steels from ductile to brittle and seriously reduces their fracture toughness. The effectiveness of the recovery treatment, as evaluated from the toughness measurements, is confirmed, although the efficiency is not the same for the steels and depends on the evaluation parameter except in the case of almost complete recovery. The recovery effect increases with the received neutron fluence if the toughness values after treatment are compared with those in the irradiated condition rather than those in the as received condition. © 1998 Elsevier Science S.A. All rights reserved.
1. Introduction Fracture mechanics assessment is a requirement of vital importance for the structural integrity of the reactor pressure vessel of a nuclear power plant. The safety criterion most usually employed to limit the risk of non-ductile fracture is based on the linear elastic fracture mechanics methodology and consists of the comparison of calculated stress intensity factor values with experimental estimates of the fracture toughness of the steel. The application of the criterion involves a periodic modification of the limit conditions for reactor operation in accordance with the updated knowledge of the
* Corresponding author. Tel.: + 34 1 3366679; fax: + 34 1 3366680.
material properties. Data obtained by testing the specimens of surveillance capsules provide measurements of these properties, and particularly Charpy tests provide indirect measurements of toughness through extensively accepted correlation equations (ASME, 1993). However, these correlations are empirical so a wide safety margin must be adopted, and the toughness values determined by this method are of necessity conservative. The direct measurement of fracture toughness might help to suppress the excessive conservatism in non-ductile fracture prevention. It would also be an important step in the optimization of surveillance programmes, especially as it may soon be necessary to extend the scope of these programmes in order to include surveillance of regenerated vessels.
0029-5493/98/$19.00 © 1998 Elsevier Science S.A. All rights reserved. PII S0029-5493(97)00356-7
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F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
This paper presents part of the results of the Spanish contribution to an international research programme promoted and coordinated by IAEA (1993) with the following aims: optimizing of the fracture resistance measurement procedures, correlation of the results between the different mechanical tests to evaluate the effects of neutron irradiation, and development of methods to reduce these effects. The main purpose of the work was the direct measurement of fracture toughness on irradiated vessel steels as well as on vessel steels regenerated by recovery post-irradiation heat treatments. The materials used in the investigation, the testing conditions, and the test methodology are described previously to present the toughness values obtained by direct measurement. On the basis of these experimental results, the embrittlement effect of neutron irradiation and the efficacy of the recovery treatment are examined, and the direct measurement of toughness as an evaluation method for irradiation damage is assessed.
2. Materials and testing procedure The experiments were carried out with three ASTM A 533 grade B class 1 ferritic steels, classified as vessel steels in accordance with the ASME specifications. All three were supplied as hot rolled, heat treated plates. Steel F was a 30 mm thick normalized (880°C) air cooled and tempered (670°C for 80 min) plate made as a laboratory product. Steels J and Q were industrially made plates with thicknesses of 250 and 220 mm, respectively. Plate J was normalized (880°C), wa-
Table 1 Chemical composition (% wt) of the steels %
Steel F
Steel J
Steel Q
C Mn Ni Mo P Cu Si S
0.18 1.49 0.62 0.55 0.020 0.16 0.27 0.001
0.18 1.43 0.63 0.48 0.004 0.05 0.23 0.002
0.19 1.41 0.84 0.50 0.019 0.14 0.25 0.004
ter quenched and tempered (660°C for 8.7 h), whereas a treatment including normalizing (900°C), water quenching, tempering (665°C for 12 h) and stress relieving (620°C for 40 h) was applied to plate Q. Their chemical compositions and mechanical properties are shown in Tables 1 and 2. The selection of the steels was made in terms of their sensibility to neutron irradiation as evaluated from their Cu, Ni and P content. According to these contents, greater sensitivity was expected for Q and F and less so for J. The fracture specimen used was the standardized CT type, 12.7 mm thick without side grooves. All the specimens were fatigue precracked in the as received condition until reaching an uncracked ligament of 10 mm. The crack front and crack growth direction were chosen to coincide with the thickness and rolling direction of the plate respectively. The tests were performed within a hot cell at room temperature, under controlled rate of the crack opening displacement as measured at the load line. Integral J was calculated by ASTM E
Table 2 Mechanical properties of the steels Properties
Steel F
Steel J
Steel Q
Young’s modulus (GPa) 0.2% Yield strength (MPa) Tensile strength (MPa) Total elongation (%) Reduction of area (%) Ramberg-Osgood constants (s = s0o 1/n)
195 499 625 23 71 6.4 990
210 412 569 32 73 5.1 990
190 482 630 26 77 6.6 1015
n s0 (MPa)
F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
813 (ASTM, 1987) standard and crack extension was measured using the compliance technique. The tensile properties relevant to the analysis of fracture data were directly measured. So tensile tests of the three steels were carried out for each one of the test conditions considered in the fracture tests.
3. Irradiation conditions and recovery treatment The test conditions for fracture testing were chosen with the purpose of evaluating both the embrittlement effect on toughness of the neutron irradiation and the recovery post-irradiation heat treatments. Different neutron irradiation fluences were used, each one followed by the same recovery treatment for some of the specimens. The irradiation temperature was typical for a PWR reactor (290°C) and the neutron fluences were comparable to those of a commercial vessel after a number of years in operation. Steel Q, used as the reference material because of its high content of Cu, Ni and P, was exposed to the following neutron fluences, which only concerns neutrons with energy above 1 MeV: A, (3.6×1018 n cm − 2 under a neutron flux of 1.5× 1011 n cm − 2 s − 1); B, (0.9 ×1019 n cm − 2 under a neutron flux of 0.8 × 1012 n cm − 2 s − 1); and C, (2.5×1019 n cm − 2 under a neutron flux of 7.7× 1012 n cm − 2 s − 1). The other two steels, F and J, were exposed only to a fluence D of 3.1×1019 n cm − 2, the neutron flux being 3.2×1012 n cm − 2 s − 1. Irradiation A was carried out in a commercial nuclear plant and took a year to complete. The capsule containing the specimens was placed in a space left vacant by a sample previously removed in the surveillance programme. Irradiations B, C and D were carried out in an experimental reactor with an accelerated flux. The duration and temperature of the recovery treatment were determined by positron annihilation and microhardness tests. The final values adopted (168 h and 400°C) were the combination of time–temperature values for which the microhardness of the recovered material and the mean life time of the positrons inside its crystal lattice attained the values of the as received condition
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(Pareja et al., 1993). This recovery treatment was applied to all steels in all irradiation conditions, so that for each irradiation state a recovery state to measure toughness was added to the testing programme. 4. Experimental results The load–COD curve obtained from a toughness test is the first indication of the fracture behaviour of the material. The different types of curves obtained in the tests can be classified in one of the four categories shown in Fig. 1. These curves represent four different fracture processes ranging from brittle fracture by cleavage cracking without macroscopic plastic deformation (curve I) to stable crack extension by ductile tearing (curve II) and include unstable fracture by cleavage cracking after macroscopic plastic deformation, with no previous stable crack extension (curve III) or with it (curve IV). The fracture behaviour exhibited in curve I is that of a brittle material for which unstable crack propagation due to cleavage is preceded by only a small amount of plastic deformation confined to the crack tip. Curve II also involves an abrupt failure due to cleavage, but is preceded by a large amount of plastic deformation which spreads across the ligament of the specimen. If prior to such cleavage failure, plastic strain at the crack tip becomes sufficiently high, a ductile crack extension may initiate, giving rise to a curve of type III. Finally, curve IV shows the typical behaviour of a highly ductile material in which the cleavage mechanism is not triggered and fracture occurs by stable crack extension and continuous tearing of the specimen. The number and conditions of the tests resulting in each type of curve are summarized in Table 3. As shown in this Table, the curves obtained for the as received material of steels J and F were of type IV, whereas for steel Q there was an roughly equal split between types I and II. The Table also indicates that the ductile behaviour exhibited by all three steels in the as received condition is significantly altered by neutron irradiation. This was clearly shown by the fact that in all the tests performed with irradiated specimens the resulting curves were of type I or II and not III or IV.
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F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
Fig. 1. Types of load – COD curves registered in the fracture tests.
In addition to this, it was observed that the three steels showed a general tendency to return to the behaviour associated with type III and IV curves after being subjected to the recovery heat treatment. However, the results were not identical for all three: steel F exhibited a behaviour producing curves of type IV, the curves obtained for steel J were a mixture of types II, III and IV, whilst steel Q barely managed to pass from type I to type II. For each of the tests, a critical value JQ of the J integral was determined according to ASTM E 813: at failure instability for curves I and II and at the onset of stable crack extension for curves III and IV. Table 4 shows the average value and S.D. of JQ for the whole series of fracture tests corresponding to each material and test condition, together with the strain hardening exponent and the yield and tensile strength as measured in the tension tests. It also includes a column indicating the predominant type of initial failure: ductile tearing (DT) for curves III and IV and cleavage instability (CI) for curves I and II. The last column is the limit value of JQ as derived from the specimen size requirements of ASTM E 813:
!
JQ B JQL Min
b s0.2 + su B s0.2 + su , 2 25 2 25
"
(1)
where B is the specimen thickness, b the uncracked ligament length, s0.2 the yield strength, and su the tensile strength.
5. Discussion The qualitative effects of irradiation and the recovery treatment have already been pointed out. The change in failure mode from ductile tearing to cleavage instability suffered by the three steels is a sure proof of embrittlement due to irradiation, whereas the opposed change produced by the recovery treatment is a clear sign of effectiveness, though with a different level of efficacy for each material. A quantitative assessment of the same effects requires some reflection on the meaning of the fracture tests. The ASTM E 813 size requirement given by Eq. (1) are fulfilled by the 1/2 inch CT specimens used for the fracture tests only under the test conditions which lead to a cleavage instability failure mode. Since J values below this limit ensure J
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Table 3 Qualitative results of the toughness tests Steel
Test condition (20°C)
Type of Curve I
II
III
IV
F
As Received Irradiated D Irradiated D and recovered
— 3 —
— 1 —
— — —
4 — 3
J
As received Irradiated D Irradiated D and recovered
— 1 —
— 3 1
— — 1
4 — 1
Q
As received Irradiated A Irradiated B Irradiated C Irradiated A and recovered Irradiated B and recovered Irradiated C and recovered
— 3 2 4 — — —
— 1 2 — 5 4 1
5 — — — — — —
5 — — — — 1 —
integral dominance of the stress and strain fields near the crack tip, the activation of the failure initiation mechanism below this limit ensures a JQ-value independent of specimen size. In the case of E 813, whose aim is restricted to quantifying the resistance to initiation of stable crack growth in metals, a measured JQ value fulfilling Eq. (1) would be an intrinsic ductile tearing fracture toughness of the material, but as already stated, the specimen dimensions only verify Eq. (1) for the test conditions that do not give rise to this type of failure. However, other standards for fracture testing such as ESIS (1992) or Schwalve et al. (1994) admit J integral as a measurement for fracture resistance characterization even if no ductile crack growth occurs prior to instability failure. This justifies the use of the JQ values of Table 4 for quantitative comparison of embrittlement in the different test conditions, although with the following considerations: —The size requirements for J dominance at fracture instability are similar to that of Eq. (1), so the toughness values which do not verify Eq. (1) cannot be considered as intrinsic toughness measurements of the material, but as these measurements related to specimen geometry. However, given that all the tests were carried out using identical specimen geometry, they can be
used for the purpose of quantitative comparison. —To compare toughness measurements corresponding to different fracture modes might not be adequate. However, instability fracture due to cleavage can occur after ductile crack growth as the curves of type III registered in the tests demonstrate. So each value of J integral measured at the onset of ductile tearing may be considered as a lower bound for the J integral value at the onset of cleavage instability since ductile tearing comes first. More restrictive size requirements than that given by Eq. (1) has been proposed for critical J values at the onset of cleavage fracture. The limits proposed by Wallin (1989) and Anderson and Dodss (1991) divide the limit of Eq. (1) by 2 and 8 respectively, though on the basis of a finite element calculation the latter authors (1993) then found that another parameter depending on specimen geometry and integral J dominates the driving force of cleavage (the maximum principal stress) near the crack tip. This parameter J* approaches J integral as the size requirement is fulfilled and its critical value, JA, substitutes that of J, JQ, as cleavage toughness measurement for specimens not fulfilling the size requirement. According to Anderson and Dodss (1993), the parameter for the E 813 specimens is as follows:
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Table 4 Quantitative results of the toughness tests Steel
Test condition (20°C)
s0.2 (MPa)
su (MPa)
Initial failure
n
JQ (kJ m−2) Mean
JQL (kJ m−2)
S.D.
F
As received Irradiated D Irradiated D and recovered
499 670 530
625 784 657
6.4 7.9 6.5
DT CI DT
907 92 832
25 56 7
225 291 237
J
As received Irradiated D Irradiated D and recovered
412 497 426
569 626 577
5.1 6.3 5.7
DT CI DT
904 60 220
34 24 51
194 225 201
Q
As received Irradiated A Irradiated B Irradiated C Irradiated A and recovered Irradiated B and recovered Irradiated C and recovered
484 534 537 620 497 483 493
622 679 688 760 639 623 631
6.6 6.4 6.1 6.5 5.9 6.4 6.4
DT CI CI CI CI CI CI
624 97 83 29 129 397 226
237 30 47 6 96 239 —
222 254 259 276 227 221 276
J bs 0.2 Á for J 5 g &g
F(g − 1) Ã1+ F J bs 0.2 J*= Í Ãg−1 bs 0.2 bs 0.2 for J ] &g &g Ä g F(g −1)
F(g − 1)
(2)
F=0.8425n 2.262 g= 1.126+ 0.01925n −0.00008333n 2
(3)
where b is the uncracked ligament length, s0.2 is the yield strength, and n is the strain hardening exponent as defined in Table 2. The statistical nature of cleavage is present in most of the micromechanical models developed in recent years to explain the fracture toughness values obtained in the brittle – ductile transition region. This explains the large scatter shown by the toughness measurements when fracture is due to cleavage instability, as is confirmed by the column of standard deviations in Table 4. As a consequence, fracture toughness must be considered as a random variable rather than as a material constant and must be characterized by a probability function. Then comparison of fracture toughness in different test conditions should be done through the parameters that determine this probability function. For ferritic steels, one of the
most widely used is that derived by Wallin (1985) on the basis of the weakest link theory with empirical corrections: the toughness distribution is a three parameter Weibull’ one with a fixed value for two of the parameters (see, for instance Wallin, 1993). The fracture toughness G, as evaluated by the critical value of J or J* at cleavage instability, has the following probability function: 4
P[G5 G]= 1−e − ( G − Gm/ G0)
(4)
where Gm = 1.82 kJ m − 2 and G0 is the parameter to be fitted from experimental data. The calculation with the probability function of Eq. (4) of the mean value G( of toughness gives: G(
&
G
Gm
dP dG dG
= 0.8862G0 + 1.8128 G0Gm + Gm
(5)
which allows the mean value G( to be estimated from any estimate of G0. With this aim the application of the maximum likelihood method to a series of N experimental toughness measurements Gi leads to (Moskovic, 1993): G0 =
'
1 N % [ Gi − Gm]4 Ni = 1
(6)
F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
137
Table 5 Statistical results of the toughness tests Steel
Test Condition (20°C)
JQ (kJ m−2)
JA (kJ m−2)
Sample mean
Sample S.D. (%)
Mean toughness Sample mean
Sample S.D. (%)
Mean toughness
F
As received Irradiated D Irradiated D and recovered
907 92 832
2.7 61.2 0.8
796 91 729
120 61 124
0.0 49.7 0.0
103 58 107
J
As received Irradiated D Irradiated D and recovered
904 60 220
3.8 39.9 23.0
793 55 195
143 48 103
0.0 33.5 8.7
124 43 89
Q
As received Irradiated A Irradiated B Irradiated C Irradiated A and recovered Irradiated B and recovered Irradiated C and recovered
624 97 83 29 129
37.9 30.7 56.1 21.9 74.4
583 87 81 25 138
109 69 61 27 76
4.0 20.3 47.0 20.1 44.4
94 60 57 23 71
397
60.3
398
105
14.0
92
226
—
196
102
—
88
Experimental toughness values obtained in the tests for the different test conditions have been quantitative compared in terms of the mean toughness given by Eq. (5) after estimating it by means of Eqs. (5) and (6). By assuming that toughness measurements follow the probability law of Eq. (4), this approach is equivalent to estimating the mean value of toughness corresponding to an infinite test series from the measurements of a sample of this series. The critical values, JQ and JA, at fracture instability or at the onset of ductile tearing of both the J integral and the J* parameter of Eq. (2) were used as a series of toughness measurements, the values at the onset of ductile tearing as lower bounds in substitution of the true values. Table 5 shows the two mean toughnesses,J( Q and J( A, for a given steel and test condition as derived from the JQ and JA values of the corresponding sample. The table also includes the mean value and the S.D. of JQ and JA in each sample. The toughness values coming from JQ are
higher than their equivalents for JA in all the cases, since J* is always less than J. Furthermore, according to Eqs. (2) and (3), J* becomes saturated as J increases, this being the reason why the S.D. almost vanishes for the test conditions in which ductile tearing failure prevails. Wallin (1989) also pointed out the saturation of the parameter that controls cleavage fracture, whose occurrence before the onset of ductile tearing reinforces the idea of using the JA value at the onset in substitution of that of cleavage instability. However, this latter JA value does not necessarily become saturated since the ductile crack growth previous to potential cleavage could modify the saturation value of J*. The quantitative effect of neutron irradiation and recovery treatment on toughness is well shown in Figs. 2 and 3, that illustrate the numerical results of Table 5. It is clear that the change in the failure mode, induced by neutron irradiation, is accompanied by a significant reduction in toughness. For the highest neutron fluence this
F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
138
Fig. 2. Effect of irradiation and heat recovery on toughness as evaluated through J( Q.
reduction becomes drastic as evaluated through J( Q, with toughness losses of about 90% or even greater in the three steels for the highest neutron fluence. The rapid decrease in toughness for increasing fluences can be observed in Fig. 4, where the two toughness measurements of steel Q, J( Q and J( A, have been plotted against neutron fluence. The resulting curves suggest the following relationships between these variables: J( %Q =
J( Q
J( %A =
J( A
Fig. 4. Effect of neutron fluence on toughness and recovered toughness of steel Q.
duced by the fluence F and FQ and FA are constants with respective values of 0.8× 1018 and 9.5×1018 n cm − 2. As Fig. 4 shows, for steel Q the recovered toughness does not seem to be a monotone function of neutron fluence. However, this lack of regularity disappears if the recovery effect is evaluated in relative terms, say, by using the quotient between the pre- and post-irradiation toughnesses J( % and J( ¦ respectively (Fig. 5). The results can be fitted with the following expressions:
(7) F F 1+ 1+ FQ FA where J( %Q and J( %A are the mean toughnesses pro-
F J( ¦Q = 3.31 ln J( %Q F0
Fig. 3. Effect of irradiation and heat recovery on toughness as evaluated through J( A.
Fig. 5. Relative effect of neutron fluence on recovery of steel Q.
J( ¦A F = 1.49 ln J( %A F0
(8)
again F being neutron fluence and F0 a new constant with a value of 2.2× 1018 n cm − 2.
F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
Combining the above Eqs. (7) and (8), the recovered toughness of steel Q can be expressed as a function of fluence: J( ¦Q =
3.31J( Q F ln F F0 1+ FQ
J( ¦A =
3.31J( A F ln F F0 1+ FA
(9)
A whole examination of the experimental results indicates that all three steels exhibit similar behaviour when irradiated at the highest fluence, with toughness losses in the region of 90% (96% for steel Q, 93% for steel J and 81% for steel F) as evaluated with J( Q. By making the evaluation with J( A, the losses are also important but lower, and the differences between the three steels show the same trend but are more pronounced (44% for steel F, 65% for steel J and 36% for steel Q). A similar analysis of the effectiveness of the recovery treatment and the sensitivity to it of steels confirms the ability of the treatment for recovering the toughness of irradiated materials, but provides different conclusions as to its efficacy for each steel depending on the parameter adopted for evaluating toughness. The recovery percentages are 91% in steel F, 24% in steel J and 42% in steel Q (as averaged for the three fluences) for J( Q, but becomes respectively 100%, 71% and 88% for J( A. Only the qualitative conclusion that steels F and J are more and less sensitive to the treatment, respectively, the former showing an almost complete recovery capacity, is common to the two evaluation methods. Finally, regarding the role of chemical composition, the test results do not seem to substantiate any relevant influence of the more important elements from the point of view of irradiation (Ni, Cu and P). Steel Q, which has the highest P, Ni and Cu content shows significant toughness losses, but they are similar to those of Steel J, with comparable Ni content and with much lower P and Cu content. Steel F, on the other hand, with the same Ni content as the J and with similar Cu and P content to Steel Q, shows slightly smaller losses and great toughness improvement due to the recovery treatment, in contrast to steels Q and J.
6. Concluding remarks Fracture tests of three vessel steels in irradiated
139
and recovered post-irradiated states were carried out at room temperature. Two fracture parameters were used for evaluating fracture toughness. According to the test results and regardless of the toughness evaluation parameter, it can be concluded that the irradiation received in amounts comparable to that of a working vessel alters the failure mechanism and considerably reduces the toughness, the reduction rate decreasing as the neutron fluence increases. The applied post-irradiation treatment produces different toughness recovery rates for each steel, the conclusions about the level of recovery being dependent on the evaluation parameter except in the only case of complete recovery. However, what stands out for the two parameters used is that such a recovery, when evaluated in relative terms, can be expressed as a simple function of the fluence received. Finally, no simple correlation between chemical composition and toughness loss and recovery was found, as might have been expected from the Ni, P and Cu content of the steels selected for testing. Acknowledgements This study was made possible by the financial support of ENDESA and Grant PR84-511 from the Spanish Office for Scientific and Technological Research (CICYT), whom the authors wish to thank. They also express their gratitude to IAEA. References Anderson, T.L., Dodss, R.H., 1991. Specimen size requirements for fracture toughness testing in the ductile – brittle transition region. J. Test. Eval. 19, 123 – 134. Anderson, T.L., Dodss, R.H., 1993. Simple constraint correction for subsize fracture toughness specimens. In: ASTM STP 1204. ASTM, Philadelphia, PA, pp. 93 – 105. ASME, 1993. Protection against nonductile fracture. Boiler and Pressure Vessel Code, Section III, Appendix G. ASME. ASTM, 1987. ASTM E 813-87 Standard Test Method for JIC, a Measure of Fracture Toughness. ASTM, Philadelphia, PA. ESIS, 1992. ESIS P2-92, Procedure for Determining the Fracture Behaviour of Materials. ESIS, Delft. International Atomic Energy Agency, 1985 – 1993. Optimizing of Reactors Pressure Vessel Surveillance Programmes and Their Analysis. IAEA Coordinated Research Programme Phase III.
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Moskovic, R., 1993. Statistical analysis of censored fracture toughness data in the ductile to brittle transition temperature region. Eng. Fract. Mech. 44, 21–41. Pareja, R., de Diego, N., de la Cruz, R., del Rio, J., 1993. Postirradiation recovery of a reactor pressure vessel steel investigated by positron annihilation and microhardness measurements. Nucl. Technol. 14, 52–63. Schwalve, K.H., Neale, B.K., Heerens, J., 1994. EFAM 94 (The GKSS Test Procedure for Determining the Fracture Behaviour of Materials). GKKS, Geesthacht.
. .
Wallin, K., 1985. The size effect in KIC results. Eng. Fract. Mech. 22, 149 – 163. Wallin, K., 1989. Fracture toughness testing in the ductile– brittle transition region. In: Salama, K. et al. (Ed), Advances in Fracture Research, vol. 1. Pergamon Press, Oxford, pp. 267 – 276. Wallin, K., 1993. Irradiation damage effects on the fracture toughness transition curves shape for reactor pressure vessel steels. Int. J. Press. Vessel. Pip. 55, 61– 79.
Fracture toughness of irradiated and recovered vessel steels F. Perosanz a, A. Valiente b,*, J. Lapen˜a a b
a Energy, En6ironment and Technology Research Center (CIEMAT), A6. Complutense 2, 28040 -Madrid, Spain Materials Science Department, Polytechnic Uni6ersity of Madrid, E.T.S. Ing. Caminos, Ciudad Uni6ersitaria s/n, 28040 -Madrid, Spain
Abstract This paper presents the fracture toughness measurements carried out on three vessel steels in an irradiated condition and after a post-irradiation recovery treatment. A statistical approach and the fracture parameters corresponding to two theoretical models of the fracture tests are used for evaluating toughness. Test results show that the neutron fluence gradually transforms the fracture behaviour of the vessel steels from ductile to brittle and seriously reduces their fracture toughness. The effectiveness of the recovery treatment, as evaluated from the toughness measurements, is confirmed, although the efficiency is not the same for the steels and depends on the evaluation parameter except in the case of almost complete recovery. The recovery effect increases with the received neutron fluence if the toughness values after treatment are compared with those in the irradiated condition rather than those in the as received condition. © 1998 Elsevier Science S.A. All rights reserved.
1. Introduction Fracture mechanics assessment is a requirement of vital importance for the structural integrity of the reactor pressure vessel of a nuclear power plant. The safety criterion most usually employed to limit the risk of non-ductile fracture is based on the linear elastic fracture mechanics methodology and consists of the comparison of calculated stress intensity factor values with experimental estimates of the fracture toughness of the steel. The application of the criterion involves a periodic modification of the limit conditions for reactor operation in accordance with the updated knowledge of the
* Corresponding author. Tel.: + 34 1 3366679; fax: + 34 1 3366680.
material properties. Data obtained by testing the specimens of surveillance capsules provide measurements of these properties, and particularly Charpy tests provide indirect measurements of toughness through extensively accepted correlation equations (ASME, 1993). However, these correlations are empirical so a wide safety margin must be adopted, and the toughness values determined by this method are of necessity conservative. The direct measurement of fracture toughness might help to suppress the excessive conservatism in non-ductile fracture prevention. It would also be an important step in the optimization of surveillance programmes, especially as it may soon be necessary to extend the scope of these programmes in order to include surveillance of regenerated vessels.
0029-5493/98/$19.00 © 1998 Elsevier Science S.A. All rights reserved. PII S0029-5493(97)00356-7
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F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
This paper presents part of the results of the Spanish contribution to an international research programme promoted and coordinated by IAEA (1993) with the following aims: optimizing of the fracture resistance measurement procedures, correlation of the results between the different mechanical tests to evaluate the effects of neutron irradiation, and development of methods to reduce these effects. The main purpose of the work was the direct measurement of fracture toughness on irradiated vessel steels as well as on vessel steels regenerated by recovery post-irradiation heat treatments. The materials used in the investigation, the testing conditions, and the test methodology are described previously to present the toughness values obtained by direct measurement. On the basis of these experimental results, the embrittlement effect of neutron irradiation and the efficacy of the recovery treatment are examined, and the direct measurement of toughness as an evaluation method for irradiation damage is assessed.
2. Materials and testing procedure The experiments were carried out with three ASTM A 533 grade B class 1 ferritic steels, classified as vessel steels in accordance with the ASME specifications. All three were supplied as hot rolled, heat treated plates. Steel F was a 30 mm thick normalized (880°C) air cooled and tempered (670°C for 80 min) plate made as a laboratory product. Steels J and Q were industrially made plates with thicknesses of 250 and 220 mm, respectively. Plate J was normalized (880°C), wa-
Table 1 Chemical composition (% wt) of the steels %
Steel F
Steel J
Steel Q
C Mn Ni Mo P Cu Si S
0.18 1.49 0.62 0.55 0.020 0.16 0.27 0.001
0.18 1.43 0.63 0.48 0.004 0.05 0.23 0.002
0.19 1.41 0.84 0.50 0.019 0.14 0.25 0.004
ter quenched and tempered (660°C for 8.7 h), whereas a treatment including normalizing (900°C), water quenching, tempering (665°C for 12 h) and stress relieving (620°C for 40 h) was applied to plate Q. Their chemical compositions and mechanical properties are shown in Tables 1 and 2. The selection of the steels was made in terms of their sensibility to neutron irradiation as evaluated from their Cu, Ni and P content. According to these contents, greater sensitivity was expected for Q and F and less so for J. The fracture specimen used was the standardized CT type, 12.7 mm thick without side grooves. All the specimens were fatigue precracked in the as received condition until reaching an uncracked ligament of 10 mm. The crack front and crack growth direction were chosen to coincide with the thickness and rolling direction of the plate respectively. The tests were performed within a hot cell at room temperature, under controlled rate of the crack opening displacement as measured at the load line. Integral J was calculated by ASTM E
Table 2 Mechanical properties of the steels Properties
Steel F
Steel J
Steel Q
Young’s modulus (GPa) 0.2% Yield strength (MPa) Tensile strength (MPa) Total elongation (%) Reduction of area (%) Ramberg-Osgood constants (s = s0o 1/n)
195 499 625 23 71 6.4 990
210 412 569 32 73 5.1 990
190 482 630 26 77 6.6 1015
n s0 (MPa)
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813 (ASTM, 1987) standard and crack extension was measured using the compliance technique. The tensile properties relevant to the analysis of fracture data were directly measured. So tensile tests of the three steels were carried out for each one of the test conditions considered in the fracture tests.
3. Irradiation conditions and recovery treatment The test conditions for fracture testing were chosen with the purpose of evaluating both the embrittlement effect on toughness of the neutron irradiation and the recovery post-irradiation heat treatments. Different neutron irradiation fluences were used, each one followed by the same recovery treatment for some of the specimens. The irradiation temperature was typical for a PWR reactor (290°C) and the neutron fluences were comparable to those of a commercial vessel after a number of years in operation. Steel Q, used as the reference material because of its high content of Cu, Ni and P, was exposed to the following neutron fluences, which only concerns neutrons with energy above 1 MeV: A, (3.6×1018 n cm − 2 under a neutron flux of 1.5× 1011 n cm − 2 s − 1); B, (0.9 ×1019 n cm − 2 under a neutron flux of 0.8 × 1012 n cm − 2 s − 1); and C, (2.5×1019 n cm − 2 under a neutron flux of 7.7× 1012 n cm − 2 s − 1). The other two steels, F and J, were exposed only to a fluence D of 3.1×1019 n cm − 2, the neutron flux being 3.2×1012 n cm − 2 s − 1. Irradiation A was carried out in a commercial nuclear plant and took a year to complete. The capsule containing the specimens was placed in a space left vacant by a sample previously removed in the surveillance programme. Irradiations B, C and D were carried out in an experimental reactor with an accelerated flux. The duration and temperature of the recovery treatment were determined by positron annihilation and microhardness tests. The final values adopted (168 h and 400°C) were the combination of time–temperature values for which the microhardness of the recovered material and the mean life time of the positrons inside its crystal lattice attained the values of the as received condition
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(Pareja et al., 1993). This recovery treatment was applied to all steels in all irradiation conditions, so that for each irradiation state a recovery state to measure toughness was added to the testing programme. 4. Experimental results The load–COD curve obtained from a toughness test is the first indication of the fracture behaviour of the material. The different types of curves obtained in the tests can be classified in one of the four categories shown in Fig. 1. These curves represent four different fracture processes ranging from brittle fracture by cleavage cracking without macroscopic plastic deformation (curve I) to stable crack extension by ductile tearing (curve II) and include unstable fracture by cleavage cracking after macroscopic plastic deformation, with no previous stable crack extension (curve III) or with it (curve IV). The fracture behaviour exhibited in curve I is that of a brittle material for which unstable crack propagation due to cleavage is preceded by only a small amount of plastic deformation confined to the crack tip. Curve II also involves an abrupt failure due to cleavage, but is preceded by a large amount of plastic deformation which spreads across the ligament of the specimen. If prior to such cleavage failure, plastic strain at the crack tip becomes sufficiently high, a ductile crack extension may initiate, giving rise to a curve of type III. Finally, curve IV shows the typical behaviour of a highly ductile material in which the cleavage mechanism is not triggered and fracture occurs by stable crack extension and continuous tearing of the specimen. The number and conditions of the tests resulting in each type of curve are summarized in Table 3. As shown in this Table, the curves obtained for the as received material of steels J and F were of type IV, whereas for steel Q there was an roughly equal split between types I and II. The Table also indicates that the ductile behaviour exhibited by all three steels in the as received condition is significantly altered by neutron irradiation. This was clearly shown by the fact that in all the tests performed with irradiated specimens the resulting curves were of type I or II and not III or IV.
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F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
Fig. 1. Types of load – COD curves registered in the fracture tests.
In addition to this, it was observed that the three steels showed a general tendency to return to the behaviour associated with type III and IV curves after being subjected to the recovery heat treatment. However, the results were not identical for all three: steel F exhibited a behaviour producing curves of type IV, the curves obtained for steel J were a mixture of types II, III and IV, whilst steel Q barely managed to pass from type I to type II. For each of the tests, a critical value JQ of the J integral was determined according to ASTM E 813: at failure instability for curves I and II and at the onset of stable crack extension for curves III and IV. Table 4 shows the average value and S.D. of JQ for the whole series of fracture tests corresponding to each material and test condition, together with the strain hardening exponent and the yield and tensile strength as measured in the tension tests. It also includes a column indicating the predominant type of initial failure: ductile tearing (DT) for curves III and IV and cleavage instability (CI) for curves I and II. The last column is the limit value of JQ as derived from the specimen size requirements of ASTM E 813:
!
JQ B JQL Min
b s0.2 + su B s0.2 + su , 2 25 2 25
"
(1)
where B is the specimen thickness, b the uncracked ligament length, s0.2 the yield strength, and su the tensile strength.
5. Discussion The qualitative effects of irradiation and the recovery treatment have already been pointed out. The change in failure mode from ductile tearing to cleavage instability suffered by the three steels is a sure proof of embrittlement due to irradiation, whereas the opposed change produced by the recovery treatment is a clear sign of effectiveness, though with a different level of efficacy for each material. A quantitative assessment of the same effects requires some reflection on the meaning of the fracture tests. The ASTM E 813 size requirement given by Eq. (1) are fulfilled by the 1/2 inch CT specimens used for the fracture tests only under the test conditions which lead to a cleavage instability failure mode. Since J values below this limit ensure J
F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
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Table 3 Qualitative results of the toughness tests Steel
Test condition (20°C)
Type of Curve I
II
III
IV
F
As Received Irradiated D Irradiated D and recovered
— 3 —
— 1 —
— — —
4 — 3
J
As received Irradiated D Irradiated D and recovered
— 1 —
— 3 1
— — 1
4 — 1
Q
As received Irradiated A Irradiated B Irradiated C Irradiated A and recovered Irradiated B and recovered Irradiated C and recovered
— 3 2 4 — — —
— 1 2 — 5 4 1
5 — — — — — —
5 — — — — 1 —
integral dominance of the stress and strain fields near the crack tip, the activation of the failure initiation mechanism below this limit ensures a JQ-value independent of specimen size. In the case of E 813, whose aim is restricted to quantifying the resistance to initiation of stable crack growth in metals, a measured JQ value fulfilling Eq. (1) would be an intrinsic ductile tearing fracture toughness of the material, but as already stated, the specimen dimensions only verify Eq. (1) for the test conditions that do not give rise to this type of failure. However, other standards for fracture testing such as ESIS (1992) or Schwalve et al. (1994) admit J integral as a measurement for fracture resistance characterization even if no ductile crack growth occurs prior to instability failure. This justifies the use of the JQ values of Table 4 for quantitative comparison of embrittlement in the different test conditions, although with the following considerations: —The size requirements for J dominance at fracture instability are similar to that of Eq. (1), so the toughness values which do not verify Eq. (1) cannot be considered as intrinsic toughness measurements of the material, but as these measurements related to specimen geometry. However, given that all the tests were carried out using identical specimen geometry, they can be
used for the purpose of quantitative comparison. —To compare toughness measurements corresponding to different fracture modes might not be adequate. However, instability fracture due to cleavage can occur after ductile crack growth as the curves of type III registered in the tests demonstrate. So each value of J integral measured at the onset of ductile tearing may be considered as a lower bound for the J integral value at the onset of cleavage instability since ductile tearing comes first. More restrictive size requirements than that given by Eq. (1) has been proposed for critical J values at the onset of cleavage fracture. The limits proposed by Wallin (1989) and Anderson and Dodss (1991) divide the limit of Eq. (1) by 2 and 8 respectively, though on the basis of a finite element calculation the latter authors (1993) then found that another parameter depending on specimen geometry and integral J dominates the driving force of cleavage (the maximum principal stress) near the crack tip. This parameter J* approaches J integral as the size requirement is fulfilled and its critical value, JA, substitutes that of J, JQ, as cleavage toughness measurement for specimens not fulfilling the size requirement. According to Anderson and Dodss (1993), the parameter for the E 813 specimens is as follows:
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Table 4 Quantitative results of the toughness tests Steel
Test condition (20°C)
s0.2 (MPa)
su (MPa)
Initial failure
n
JQ (kJ m−2) Mean
JQL (kJ m−2)
S.D.
F
As received Irradiated D Irradiated D and recovered
499 670 530
625 784 657
6.4 7.9 6.5
DT CI DT
907 92 832
25 56 7
225 291 237
J
As received Irradiated D Irradiated D and recovered
412 497 426
569 626 577
5.1 6.3 5.7
DT CI DT
904 60 220
34 24 51
194 225 201
Q
As received Irradiated A Irradiated B Irradiated C Irradiated A and recovered Irradiated B and recovered Irradiated C and recovered
484 534 537 620 497 483 493
622 679 688 760 639 623 631
6.6 6.4 6.1 6.5 5.9 6.4 6.4
DT CI CI CI CI CI CI
624 97 83 29 129 397 226
237 30 47 6 96 239 —
222 254 259 276 227 221 276
J bs 0.2 Á for J 5 g &g
F(g − 1) Ã1+ F J bs 0.2 J*= Í Ãg−1 bs 0.2 bs 0.2 for J ] &g &g Ä g F(g −1)
F(g − 1)
(2)
F=0.8425n 2.262 g= 1.126+ 0.01925n −0.00008333n 2
(3)
where b is the uncracked ligament length, s0.2 is the yield strength, and n is the strain hardening exponent as defined in Table 2. The statistical nature of cleavage is present in most of the micromechanical models developed in recent years to explain the fracture toughness values obtained in the brittle – ductile transition region. This explains the large scatter shown by the toughness measurements when fracture is due to cleavage instability, as is confirmed by the column of standard deviations in Table 4. As a consequence, fracture toughness must be considered as a random variable rather than as a material constant and must be characterized by a probability function. Then comparison of fracture toughness in different test conditions should be done through the parameters that determine this probability function. For ferritic steels, one of the
most widely used is that derived by Wallin (1985) on the basis of the weakest link theory with empirical corrections: the toughness distribution is a three parameter Weibull’ one with a fixed value for two of the parameters (see, for instance Wallin, 1993). The fracture toughness G, as evaluated by the critical value of J or J* at cleavage instability, has the following probability function: 4
P[G5 G]= 1−e − ( G − Gm/ G0)
(4)
where Gm = 1.82 kJ m − 2 and G0 is the parameter to be fitted from experimental data. The calculation with the probability function of Eq. (4) of the mean value G( of toughness gives: G(
&
G
Gm
dP dG dG
= 0.8862G0 + 1.8128 G0Gm + Gm
(5)
which allows the mean value G( to be estimated from any estimate of G0. With this aim the application of the maximum likelihood method to a series of N experimental toughness measurements Gi leads to (Moskovic, 1993): G0 =
'
1 N % [ Gi − Gm]4 Ni = 1
(6)
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Table 5 Statistical results of the toughness tests Steel
Test Condition (20°C)
JQ (kJ m−2)
JA (kJ m−2)
Sample mean
Sample S.D. (%)
Mean toughness Sample mean
Sample S.D. (%)
Mean toughness
F
As received Irradiated D Irradiated D and recovered
907 92 832
2.7 61.2 0.8
796 91 729
120 61 124
0.0 49.7 0.0
103 58 107
J
As received Irradiated D Irradiated D and recovered
904 60 220
3.8 39.9 23.0
793 55 195
143 48 103
0.0 33.5 8.7
124 43 89
Q
As received Irradiated A Irradiated B Irradiated C Irradiated A and recovered Irradiated B and recovered Irradiated C and recovered
624 97 83 29 129
37.9 30.7 56.1 21.9 74.4
583 87 81 25 138
109 69 61 27 76
4.0 20.3 47.0 20.1 44.4
94 60 57 23 71
397
60.3
398
105
14.0
92
226
—
196
102
—
88
Experimental toughness values obtained in the tests for the different test conditions have been quantitative compared in terms of the mean toughness given by Eq. (5) after estimating it by means of Eqs. (5) and (6). By assuming that toughness measurements follow the probability law of Eq. (4), this approach is equivalent to estimating the mean value of toughness corresponding to an infinite test series from the measurements of a sample of this series. The critical values, JQ and JA, at fracture instability or at the onset of ductile tearing of both the J integral and the J* parameter of Eq. (2) were used as a series of toughness measurements, the values at the onset of ductile tearing as lower bounds in substitution of the true values. Table 5 shows the two mean toughnesses,J( Q and J( A, for a given steel and test condition as derived from the JQ and JA values of the corresponding sample. The table also includes the mean value and the S.D. of JQ and JA in each sample. The toughness values coming from JQ are
higher than their equivalents for JA in all the cases, since J* is always less than J. Furthermore, according to Eqs. (2) and (3), J* becomes saturated as J increases, this being the reason why the S.D. almost vanishes for the test conditions in which ductile tearing failure prevails. Wallin (1989) also pointed out the saturation of the parameter that controls cleavage fracture, whose occurrence before the onset of ductile tearing reinforces the idea of using the JA value at the onset in substitution of that of cleavage instability. However, this latter JA value does not necessarily become saturated since the ductile crack growth previous to potential cleavage could modify the saturation value of J*. The quantitative effect of neutron irradiation and recovery treatment on toughness is well shown in Figs. 2 and 3, that illustrate the numerical results of Table 5. It is clear that the change in the failure mode, induced by neutron irradiation, is accompanied by a significant reduction in toughness. For the highest neutron fluence this
F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
138
Fig. 2. Effect of irradiation and heat recovery on toughness as evaluated through J( Q.
reduction becomes drastic as evaluated through J( Q, with toughness losses of about 90% or even greater in the three steels for the highest neutron fluence. The rapid decrease in toughness for increasing fluences can be observed in Fig. 4, where the two toughness measurements of steel Q, J( Q and J( A, have been plotted against neutron fluence. The resulting curves suggest the following relationships between these variables: J( %Q =
J( Q
J( %A =
J( A
Fig. 4. Effect of neutron fluence on toughness and recovered toughness of steel Q.
duced by the fluence F and FQ and FA are constants with respective values of 0.8× 1018 and 9.5×1018 n cm − 2. As Fig. 4 shows, for steel Q the recovered toughness does not seem to be a monotone function of neutron fluence. However, this lack of regularity disappears if the recovery effect is evaluated in relative terms, say, by using the quotient between the pre- and post-irradiation toughnesses J( % and J( ¦ respectively (Fig. 5). The results can be fitted with the following expressions:
(7) F F 1+ 1+ FQ FA where J( %Q and J( %A are the mean toughnesses pro-
F J( ¦Q = 3.31 ln J( %Q F0
Fig. 3. Effect of irradiation and heat recovery on toughness as evaluated through J( A.
Fig. 5. Relative effect of neutron fluence on recovery of steel Q.
J( ¦A F = 1.49 ln J( %A F0
(8)
again F being neutron fluence and F0 a new constant with a value of 2.2× 1018 n cm − 2.
F. Perosanz et al. / Nuclear Engineering and Design 182 (1998) 131–140
Combining the above Eqs. (7) and (8), the recovered toughness of steel Q can be expressed as a function of fluence: J( ¦Q =
3.31J( Q F ln F F0 1+ FQ
J( ¦A =
3.31J( A F ln F F0 1+ FA
(9)
A whole examination of the experimental results indicates that all three steels exhibit similar behaviour when irradiated at the highest fluence, with toughness losses in the region of 90% (96% for steel Q, 93% for steel J and 81% for steel F) as evaluated with J( Q. By making the evaluation with J( A, the losses are also important but lower, and the differences between the three steels show the same trend but are more pronounced (44% for steel F, 65% for steel J and 36% for steel Q). A similar analysis of the effectiveness of the recovery treatment and the sensitivity to it of steels confirms the ability of the treatment for recovering the toughness of irradiated materials, but provides different conclusions as to its efficacy for each steel depending on the parameter adopted for evaluating toughness. The recovery percentages are 91% in steel F, 24% in steel J and 42% in steel Q (as averaged for the three fluences) for J( Q, but becomes respectively 100%, 71% and 88% for J( A. Only the qualitative conclusion that steels F and J are more and less sensitive to the treatment, respectively, the former showing an almost complete recovery capacity, is common to the two evaluation methods. Finally, regarding the role of chemical composition, the test results do not seem to substantiate any relevant influence of the more important elements from the point of view of irradiation (Ni, Cu and P). Steel Q, which has the highest P, Ni and Cu content shows significant toughness losses, but they are similar to those of Steel J, with comparable Ni content and with much lower P and Cu content. Steel F, on the other hand, with the same Ni content as the J and with similar Cu and P content to Steel Q, shows slightly smaller losses and great toughness improvement due to the recovery treatment, in contrast to steels Q and J.
6. Concluding remarks Fracture tests of three vessel steels in irradiated
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and recovered post-irradiated states were carried out at room temperature. Two fracture parameters were used for evaluating fracture toughness. According to the test results and regardless of the toughness evaluation parameter, it can be concluded that the irradiation received in amounts comparable to that of a working vessel alters the failure mechanism and considerably reduces the toughness, the reduction rate decreasing as the neutron fluence increases. The applied post-irradiation treatment produces different toughness recovery rates for each steel, the conclusions about the level of recovery being dependent on the evaluation parameter except in the only case of complete recovery. However, what stands out for the two parameters used is that such a recovery, when evaluated in relative terms, can be expressed as a simple function of the fluence received. Finally, no simple correlation between chemical composition and toughness loss and recovery was found, as might have been expected from the Ni, P and Cu content of the steels selected for testing. Acknowledgements This study was made possible by the financial support of ENDESA and Grant PR84-511 from the Spanish Office for Scientific and Technological Research (CICYT), whom the authors wish to thank. They also express their gratitude to IAEA. References Anderson, T.L., Dodss, R.H., 1991. Specimen size requirements for fracture toughness testing in the ductile – brittle transition region. J. Test. Eval. 19, 123 – 134. Anderson, T.L., Dodss, R.H., 1993. Simple constraint correction for subsize fracture toughness specimens. In: ASTM STP 1204. ASTM, Philadelphia, PA, pp. 93 – 105. ASME, 1993. Protection against nonductile fracture. Boiler and Pressure Vessel Code, Section III, Appendix G. ASME. ASTM, 1987. ASTM E 813-87 Standard Test Method for JIC, a Measure of Fracture Toughness. ASTM, Philadelphia, PA. ESIS, 1992. ESIS P2-92, Procedure for Determining the Fracture Behaviour of Materials. ESIS, Delft. International Atomic Energy Agency, 1985 – 1993. Optimizing of Reactors Pressure Vessel Surveillance Programmes and Their Analysis. IAEA Coordinated Research Programme Phase III.
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Moskovic, R., 1993. Statistical analysis of censored fracture toughness data in the ductile to brittle transition temperature region. Eng. Fract. Mech. 44, 21–41. Pareja, R., de Diego, N., de la Cruz, R., del Rio, J., 1993. Postirradiation recovery of a reactor pressure vessel steel investigated by positron annihilation and microhardness measurements. Nucl. Technol. 14, 52–63. Schwalve, K.H., Neale, B.K., Heerens, J., 1994. EFAM 94 (The GKSS Test Procedure for Determining the Fracture Behaviour of Materials). GKKS, Geesthacht.
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Wallin, K., 1985. The size effect in KIC results. Eng. Fract. Mech. 22, 149 – 163. Wallin, K., 1989. Fracture toughness testing in the ductile– brittle transition region. In: Salama, K. et al. (Ed), Advances in Fracture Research, vol. 1. Pergamon Press, Oxford, pp. 267 – 276. Wallin, K., 1993. Irradiation damage effects on the fracture toughness transition curves shape for reactor pressure vessel steels. Int. J. Press. Vessel. Pip. 55, 61– 79.