COMPARISON OF THE SCIENTIFIC BASIS OF RUSSIAN AND EUROPEAN APPROACHES FOR EVALUATING IRRADIATION EFFECTS IN REACTOR PRESSURE VESSELS

Transcription

1 LIMITED DISTRIBUTION COMPARISON OF THE SCIENTIFIC BASIS OF RUSSIAN AND EUROPEAN APPROACHES FOR EVALUATING IRRADIATION EFFECTS IN REACTOR PRESSURE VESSELS Kim Wallin European Network on Ageing Materials Evaluation and Studies Espoo, December 1994 VTT Manufacturing Technology P.O. Box 1704, FIN VTT, Finland Tel , Telefax

2 ABSTRACT Irradiation affects the properties of reactor pressure vessel material. Different countries use different approaches for evaluating these irradiation effects. Because the approaches are different, also the accuracy of them is likely to vary. In this report the scientific basis of Russian and European approaches for evaluating irradiation effects in reactor pressure vessels have been compared. As representative European approaches, the ASME and French methodologies have been examined. It can be concluded that the Russian and European approaches have similar scientific bases, but mechanistically neither approach is optimal for evaluating the irradiation effects. The approaches are based upon the Charpy-V test that is not directly descriptive of the materials true fracture toughness. At the same time highly conservative reference fracture toughness curves are usually applied. The comparison emphasizes the need for a new improved approach for evaluating irradiation effects in reactor pressure vessels. The new approach should be based upon a direct determination of the fracture toughness, or a validated correlation based on a direct determination of the fracture toughness, combined with a mechanistic treatment of irradiation and material variables. This is the subject matter of Task Group C of AMES Project 1. 1

3 PREFACE This report has been compiled at the VTT Manufacturing Technology as part of a state-of-the-art review on irradiation embrittlement, surveillance and mitigation methods carried out by the European Network for Ageing Materials Evaluation and Studies (AMES). The work is financed by CEC DG XI through the contract "Centre d'etudes de Saclay Gif Sur Yvette Cedex, Ref B018750, M. Soulat". Additional financing was provided by the Finnish Centre for Radiation and Nuclear Safety (STUK) and VTT. Associated with this subject, reports are prepared under the auspice of AMES on thermal annealing of irradiation effects and the evaluation of the available and possible mitigation methods. The author acknowledges his gratitude to Acad. Myrddin Davies, Dr. Colin English, Dr. Karen Gott and Dr. Pierre Petrequin for conducting a peer review for this report. 2

4 CONTENTS ABSTRACT...1 PREFACE INTRODUCTION RUSSIAN APPROACH Background Fracture toughness reference curves Critical brittleness temperature Chemistry factor Surveillance methodology EUROPEAN APPROACHES Background Fracture toughness reference curves Reference temperature Chemistry factor Surveillance methodology MECHANISTIC INTERPRETATION OF THE APPROACHES Fracture toughness reference curves Reference temperature shift Chemistry factor CONCLUSIONS...30 REFERENCES

5 1 INTRODUCTION Neutron irradiation affects the properties of reactor pressure vessel material. The material changes that may affect the use of the pressure vessel are connected to the materials fracture resistance. As a rule irradiation reduces the materials fracture resistance and therefore it is important to know the rate of reduction as a function of irradiation dose. Different countries apply different approaches for evaluating these irradiation effects. Because the approaches are different, also the accuracy of them is likely to vary. The largest differences between approaches for evaluating irradiation effects in pressure vessels are between Eastern (Russian) and Western (European) approaches. The safety of Central and Eastern European reactors have recently been the subject of close scrutiny by international experts. The scientific basis of the different approaches for evaluating irradiation effects in reactor pressure vessels need to be compared. 2 RUSSIAN APPROACH The Russian approach for evaluating irradiation effects in reactor pressure vessels is applied in the case of all VVER-440 and VVER-1000 type reactor pressure wessels in the Central and Eastern European countries [1]. 2.1 Background The safety assessment of reactor pressure vessels is based upon linear elastic fracture mechanics (LEFM). The reactor pressure vessel lifetime is thus determined solely by it's resistance to brittle fracture. The fracture mechanical material property describing the (LEFM) fracture resistance is denoted fracture toughness (KIC). KIC describes the materials resistance towards brittle fracture initiation under static or "quasistatic" loading. KIC is related to temperature, and irradiation changes this relation. The assessment thus requires knowledge regarding the change of KIC as a function of both temperature and irradiation. In the Russian approach the KIC for the assessment is not determined directly, but instead a reference curve methodology is used. It is assumed that the temperature dependence of fracture toughness is not affected by irradiation enabling the fracture toughness temperature dependence to be described by a single curve. Irradiation is assumed only to shift the location of the curve to higher temperatures. Thus the irradiation effects in pressure vessels are evaluated by an estimation of this temperature shift. In the Russian approach the shift is either determined from Charpy-V impact tests or from the chemical composition of the material, applying a special chemistry factor. 4

6 2.2 Fracture toughness reference curves The Russian approach applies material specific reference curves (different for VVER-440 base material, VVER-1000 base material and their welds). The reference curves are based upon KIC test results corresponding to the non-irradiated material state, measured by comparatively large test specimens. Each material has different curves for 1) normal operating conditions, 2) operational occurrences and hydraulic tests and 3) emergency situations. The initial reference curve, corresponding to emergency situations, is based on an "eye-ball" lower envelope of experimental data. The curves for normal operating conditions and operational occurrences are derived as the lower envelope of two curves determined on the basis of the initial curve. One of the curves is derived by dividing the ordinates of the initial curve by a safety factor nk {1) nk = 2, 2) nk = 1.5}, while the other is derived by shifting the initial curve along the x-axis by an amount T {1) & 2) T = 30 C}. The method of construction forces the 1) and 2) curves together at higher toughness levels. The curves are given in terms of an effective temperature (T-Tk), where Tk denotes a critical brittleness temperature. The three different initial curves (emergency situations) are: VVER-440 base materials (12X2MΦPA, 15X2MΦA and 15X2MΦA-A) [KI]3 = exp{0.02 (T-Tk)} VVER-1000 base materials (15X2HMΦA and 15X2HMΦA-A) [KI]3 = exp{ (T-Tk)} VVER-440 and VVER-1000 beltline welds (15X2MΦA, 15X2MΦA-A, 15X2HMΦA and 15X2HMΦA-A) [KI]3 = exp{ (T-Tk)} The Russian reference fracture toughness curves are presented graphically in Figs Critical brittleness temperature The critical brittleness temperature Tk forms the essence of the Russian approach for evaluating irradiation effects in reactor pressure vessels. The temperature is determined from Charpy-V impact test results. The definition for Tk is not however fixed, but varies as a function of the material's true room temperature yield stress. The yield stress relates to the average value for three or more tensile tests and the maximum value for only two tests. Impact tests are performed at different temperatures close to the expected critical brittleness temperature. Conservative estimates of Tk for the different materials are given in the code [1]. The criterion for Tk is determined by a combination of absorbed energy required (impact toughness) 5

7 and 50% ductile fracture appearance determined from the broken specimen fracture surface. A preliminary determination of Tk is first performed using specimens at a few test temperatures. Based upon the preliminary estimate, additional tests are performed at surrounding temperatures in order to make the estimation more accurate. 6

8 Fig. 1 Russian approach reference fracture toughness curves for VVER-440 base materials. Temperature normalized by critical brittleness temperature. Fig. 2 Russian approach reference fracture toughness curves for VVER-1000 base materials. Temperature normalized by critical brittleness temperature. 7

9 Fig. 3 Russian approach reference fracture toughness curves for VVER-440 and VVER-1000 welds. Temperature normalized by critical brittleness temperature. A minimum of three tests is required at each temperature. If any of the requirements set for the minimum value is not met on one of the three test specimens, one is allowed to test three additional specimens at that temperature. Thus if five tests out of 6 fulfil the requirements, the results are acceptable. Tk is taken to be the lowest test temperature at (and above) which a fulfilment of the criteria in table 1 are achieved. Table 1. Criteria for the definition of Tk. Yield stress 20 C (MPa) Mean impact toughness (J) Minimum impact toughness (J) Tk+30 o C mean impact toughness (J) Tk+30 o C min. impact toughness (J) Tk+30 o C minimum ductile (%) < < < <

10 If the number of available specimens is, in the case of irradiated specimens, inadequate for a detailed determination of Tk, a simplified approach is allowed, but at least 12 specimens must be tested. In this case the test data is fitted, applying the least square method, by the equation where KCV is the impact energy, A is the average impact energy between the upper shelf energy (KCVmax) and the lower shelf energy (KCVmin), B = (KCVmax-KCVmin)/2, T0 is the temperature corresponding to A and C is a constant describing the steepness of the impact energy temperature dependence. For the determination of Tk, the mean impact toughness criteria presented in table 1 is used. Tk is then used together with the appropriate fracture toughness reference curves to estimate the materials fracture toughness. If the neutron fluency to which Tk refers, is relevant, the experimentally determined Tk can be used as such. Otherwise the change of Tk as a function of neutron fluency has to be evaluated. Irradiation shift The shift of the critical brittleness temperature due to irradiation may be determined by the equation where TkF is the critical brittleness temperature after irradiation and TkI is the initial critical brittleness temperature. The irradiation shift is needed to determine the radiation embrittlement coefficient. Radiation embrittlement coefficient KCV = A+ B tanh T F =T kf - T T ki - T C 0 Based upon the irradiation shift ( TF) the radiation embrittlement coefficient AF can be determined from: AF = T F F F where Fn is the neutron fluency with E > 0.5 MeV, F0 = neutrons/m 2 and n is a constant. If data corresponding to several different fluences is available, the power n can be determined experimentally, otherwise a value of 1/3 is assumed. The validity of the above equation is limited to a fluency range of < Fn < neutrons/m 2. The radiation embrittlement coefficient enables the determination of TF and thus Tk for any fluency within the validity range of the equation. n 0 n 9

11 2.4 Chemistry factor The radiation embrittlement coefficient AF can also be estimated indirectly based upon the materials chemistry. In this context the radiation embrittlement coefficient is also called the chemistry factor and it is used for vessels that do not have a specific surveillance program. For base materials and welds corresponding to VVER-1000 reactors, the Russian code gives fixed (material type dependent) values for the chemistry factor. For VVER-440 type reactor welds, the chemistry factor has been determined on the basis of an empirical treatment of Charpy-V impact data ( TF), obtaining an estimate of the dependence of radiation embrittlement against the content of phosphorus and copper. The chemistry factor has for VVER-440 type reactor welds the form and AF = 800-(P[%] Cu[%]) for irradiation temperature 270 C AF = 800-(P[%] Cu[%]) + 8 for irradiation temperature 250 C. 2.5 Surveillance methodology Older vessels do not have a surveillance programme, and for them the chemistry factor concept is applied. In this case characteristic (conservative) values are given in the code both for TkI and AF [1]. The surveillance methodology for newer vessels is based upon Charpy-V and tensile testing, by which the critical brittleness temperature is determined. No fracture toughness testing is prescribed, only the use of the fracture toughness reference curves. 3 EUROPEAN APPROACHES Most European approaches for evaluating irradiation effects in reactor pressure vessels are originally derived from the ASME code [2,3]. In some countries, like e.g. UK, no accepted standard approach exists, but a case by case best estimate analysis is carried out [4]. Many countries apply the ASME code directly, but in some countries like France [5,6] the approaches have, however, experienced a national variation. 10

12 3.1 Background The ASME code is based upon linear elastic fracture mechanics, similar to the Russian code. It too uses a reference curve methodology for estimating the materials fracture toughness, but the scientific basis is somewhat different. The French approach originates from the ASME methodology, but it has been made more flexible, to account for possible plasticity effects. Both approaches assume that the temperature dependence of fracture toughness is not affected by irradiation enabling the fracture toughness temperature dependence to be described by a single curve. As in the Russian approach, the shift is either determined based upon Charpy-V impact tests or from the chemical composition of the material, applying a chemistry factor. 3.2 Fracture toughness reference curves The ASME approach The ASME code section XI includes both a static fracture initiation reference curve, as well as a crack arrest reference curve. The curves are not material dependent as in the Russian approach. It is additionally assumed that the difference between static initiation and crack arrest is constant. Both curves are given in the form of an effective temperature (T-RTNDT), where RTNDT denotes the Nil-Ductility Reference Temperature (see 3.3). For the fracture mechanical assessment of normal operation conditions the crack arrest curve also describes crack initiation and it is then denoted KIR. The ASME code only gives the reference curves in a graphical form, but the French approach gives descriptive equations for them as follows [6] KIC = min exp [0.036 (T-RTNDT+55.5)] 220 MPa m and KIR = min exp [ (T-RTNDT+88.9)] 220 MPa m where units are in MPa m and C. The reference curves have been developed based upon empirical "eye ball" lower envelope curve fitting to experimental LEFM unirradiated fracture toughness data. The KIR curve is intended to describe normal operation conditions and the KIC curve emergency and faulted conditions. In addition to the different reference curves, safety factors are applied. For normal operation conditions a safety factor of 10 is 11

13 applied either upon the allowable crack size, or a safety factor of 10 upon KI or KIa, (KIR). The corresponding safety factors for emergency and faulted conditions are 2 for the allowable crack size or 2 for KI or KIC. The ASME reference curves are presented graphically in Fig. 4. The French approach The French approach applies essentially the ASME fracture toughness reference curves, but the safety factors are different and also the definition of "upper shelf" differs from ASME. The approach recognizes three different situations with different safety factors: 1) LEVEL A; normal and upset, 2) LEVEL C; emergency and 3) LEVEL D; faulted conditions. The upper shelf toughness denoted as KJC is, if not determined directly from an elastic plastic JIC-test, 165 MPa m up to a temperature of +150 C and 150 Mpa m in the temperature range +150 C C Fig. 4 The ASME fracture toughness reference curves and curves corresponding to normal and emergency situations. Temperature normalized by the Nil-Ductility Reference Temperature. The level A fracture toughness (KcpA) is calculated from: T-RTNDT 50 C => KcpA = minimum of {0.4-KIC or 0.7 KIa} T RTNDT > 50 C => KcpA = minimum of {0.7-KIa or 0.7 KJC}. 12

14 The level C fracture toughness (KcpC) is calculated from: T-RTNDT 50 C => KcpC = minimum of {0.5-KIC or 0.85 KIa} T RTNDT > 50 C => KcpC = minimum of {0.85-KIC or 0.85 KJC}. The level D fracture toughness (KcpC) is calculated from: T-RTNDT < 100 C => KcpD = minimum of {0.8-KIC* or 0.9 KJC} T-RTNDT > 100 C => KcpD is determined based upon direct measurement of J- a curves (no brittle fracture). The French fracture toughness reference curves are presented graphically in Fig. 5. *Crack arrest is assumed if KI < 0.8 KIa before crack reaches 3/4 wall thickness. Fig. 5 The French fracture toughness reference curves. Temperature normalized by the Nil-Ductility Reference Temperature. 3.3 Reference temperature The Nil-Ductility Reference Temperature (RTNDT) is determined from the nil-ductility temperature (NDT) and the Charpy-V impact test. The NDT temperature is determined in accordance with the ASTM Method for Conducting Drop-Weight Test to Determine Nil-Ductility Transition Temperature of Ferritic 13

15 Steels (E 208). If the minimum impact energy and lateral expansion in the Charpy-V test (3 specimens) are at least 68 J and 0.9 mm respectively, at a temperature equal to NDT + 33 C, the NDT temperature is taken to represent RTNDT. If the Charpy-V properties do not meet the above criterion, RTNDT is taken as the temperature at which the requirements are reached, minus 33 C. Normally, for modern steels RTNDT is equal to NDT. The ASME approach for determining irradiation shift The NDT test is not included in the surveillance programs. The irradiation induced shift of the RTNDT reference temperature is defined as the shift in the 41 J impact energy transition temperature TK41J. It is assumed that the true static fracture toughness shift and crack arrest shift is equal to or less than TK 4IJ ( RT NDT). The U.S. Regulatory Guide 1.99, Revision 2 [7], requires the use of an additional margin in the determination of RT NDT. For welds the margin is 31 C and for base metal 19 C. In both cases the margin is not, however, more than RT NDT/2. This margin is put on top of the experimental RT NDT. The margin is not directly prescribed in the ASME approach, nor in European approaches equivalent to ASME. Based upon RT NDT the fluence dependence can be calculated by the equation [7] ( log f) RT NDT = (CF) f where CF is the radiation embrittlement coefficient and f is the neutron fluence ( n/cm 2 ) with E > 1 MeV. When two or more credible surveillance data sets are available, the radiation embrittlement coefficient is determined from the experimental data by least square fitting. The French approach for determining irradiation shift The French approach differs from the ASME methodology. The irradiation induced shift of the RTNDT reference temperature is defined as the shift in the 56 J impact energy transition temperature TK56J or the shift in the 0.9 mm lateral expansion transition temperature TK0.9mm, whichever is greater. The lateral expansion corresponding to a certain energy level, is affected by the material yield strength. Increasing the yield stress, makes plastic deformation of the specimen more difficult. Therefore, RTNDT is generally controlled by TK 0.9mm. The French approach does not apply additional safety margins. The French codes [5,6] allow the fluence dependence of RT NDT to be determined experimentally, but do not give any recommendations for type of expression to be used. In this respect the French approach is more flexible than other approaches. 3.4 Chemistry factor 14

16 The application of the chemistry factor is similar to the Russian approach. The European approaches also apply the chemistry factor concept in an effort to determine the irradiation shift directly from the steel chemistry. For the ASME methodology type steels, the U.S. Regulatory Guide 1.99, Revision 2 [7] gives different chemistry factors for welds and base metal in a tabulated form. The tables are presented graphically in Figs. 6 and 7. The tables include only the effect of copper and nickel. This does not mean that investigations would have shown phosphorus and sulphur to have no effect, but that the materials for which the chemistry factors have been developed have had a comparatively constant level of these elements. This means that the applicability of the chemistry factor is restricted to a special population of materials. Even when RTNDT is calculated with the chemistry factor, one is required to use the additional safety margin prescribed in the Regulatory Guide [7]. Fig. 6 Reg. Guide 1.99, Rev. 2 Chemistry Factor for welds. 15

17 Fig. 7 Reg. Guide 1.99, Rev. 2 Chemistry Factor for base metal. The French codes from 1988 contain a chemistry factor equation of the form RT NDT = [ (%Cu ) (%P )) [f / 10 where the terms (%P-0.008) and (%Cu-0.08) become zero for P < % and Cu < 0.08 %. The equation is applicable in a fluence range n/cm 2 and for irradiation temperatures between 275 C C. The equation originates from the U.S. Regulatory Guide 1.99, Revision 1 [8]. Essentially the same equation is used (in a graphical form having a fixed % P) in the German KTA code [9], but only for selecting materials to be included into the surveillance program. The French approach includes also a newer so called FIS equation giving an upper bound estimate of the chemistry factor [5] 19 ] 1/2 o 2 φ FIS ( C)= 8 +( (%P )+ 238(%Cu )+191 Ni Cu) where the terms (%P-0.008) and (%Cu-0.08) become zero for P < % and Cu < 0.08 %. 16

18 3.5 Surveillance methodology The surveillance methodology in Europe varies. Usually, the surveillance methodology is based upon Charpy-V testing, from which the RTNDT is determined. The French methodology includes also fracture toughness specimens, but they are mainly used to determine the effect of irradiation on the ductile fracture properties. 4 MECHANISTIC INTERPRETATION OF THE APPROACHES Mechanistically the Russian and European approaches for evaluating irradiation effects in reactor pressure vessels are very similar. All the different approaches rely on the Charpy-V impact test to determine the effect of irradiation on the materials fracture toughness. In none of the approaches is the fracture toughness determined directly. The reason for this is that originally there was no fracture toughness testing standard that would be suitable for surveillance testing. Therefore the different approaches have had to apply the Charpy-V test, combined with reference temperature and reference fracture toughness concepts. Unfortunately, these concepts have never been comprehensively, experimentally or theoretically, validated for irradiation embrittlement. This does of course not mean that the concepts would necessarily lead to an unconservative estimate of the materials fracture toughness. 4.1 Fracture toughness reference curves All the fracture toughness reference curves are based on linear-elastic fracture mechanical tests on relevant materials. In all cases, however, the fracture toughness has been determined only for unirradiated material. The crack arrest and dynamic fracture toughness tests used to determine the ASME KIR curve (KIa) were not performed according to any testing standard, because there did not exist one at the time. This was not considered a problem at the time because the reference fracture toughness curves are essentially intended for the description of static brittle fracture initiation toughness. This toughness was determined according to the ASTM standard E 399 called, Standard Test Method for Plane-Strain Fracture Toughness of Metallic Materials. The Russian fracture toughness tests, used in their reference curve development, were performed and analysed according to the corresponding, practically identical, Russian GOST standard. Thus the mechanical basis for the fracture toughness reference curves is the same. The standard E 399, denotes the valid measured entity as plane strain fracture toughness. The definition of plane strain in the standard is not, however, directly related to the actual stress state in front of the crack tip of the specimen. Classically, the plane strain fracture toughness is supposed to represent a lower bound value. 17

19 Originally, in the development of the test standard, high strength steel and aluminium and titanium alloys were used. These materials fail typically by a ductile, strain controlled mechanism having an increasing tearing resistance as a function of crack growth. For these materials, typically, specimens with small in-plane dimensions yield lower toughness values than specimens with large in-plane dimension (Fig. 8 [10]). The specimen thickness does not have a significant effect for these materials (Fig. 8 [10]). Plane-strain was in the standard development defined as a criterion guaranteeing essentially specimen size insensitive toughness values. Usually, for those materials however, the value was not a minimum but a maximum. This does not directly imply that valid KIC values for pressure vessel steel displaying brittle fracture, would not represent plane-strain. It only means that the definition in the standard is unrelated to the true stress state. Most importantly, valid KIC values guarantee that the specimen behaves macroscopically in a linear-elastic manner. The problem with valid KIC testing of pressure vessel steels is that the valid specimen size is dependent on the fracture toughness. This has led to the use of small specimens in the low toughness temperature regime and large specimens in the higher temperature regime. This would not be a problem if the valid fracture toughness were also specimen size independent also in the case of brittle fracture. Unfortunately, pressure vessel steels, showing brittle fracture, were not included in the original test development work. New experimental and theoretical work have shown that brittle fracture initiation toughness in reality shows a specimen size effect, such that large specimens yield lower values than small specimens. 18

20 Fig. 8 Plane-strain fracture toughness of 75 mm thick 2219-T851 plate (material typical for the development of E 399) [10]. Results indicate increasing fracture toughness with increasing specimen in-plane dimensions. This size effect is active also for valid KIC values. Fig. 9 shows KIC data for the HSST 02 plate (A533B Cl.l) originally forming the major part of the data set leading to the ASME KIC reference curve [11]. For clarity the smallest specimens (tested at lower shelf temperatures), and the largest specimens (showing inhomogeneity effects) have been omitted. A clear difference between the 50 mm thick specimens and the mm thick specimens can be seen. The temperature T0 reported in the figure, denotes the temperature where the mean fracture toughness is 100 MPa m. Mechanistically the size effect can be explained by a statistical sampling effect due to an increased crack front length (specimen thickness). The statistical size effect can be expressed in the form [12] K = K +( K ) B B where B1 and B2 are two different specimen thicknesses and Kmin is a lower limiting fracture toughness, that for pressure vessel steels can be approximated by Kmin = 20 MPa m. Fig. 10 shows the effect of the size correction on the HSST 02 data from Fig K B2 min B1 min 1 2 1/4 19

21 Fig. 9 KIC results of the HSST 02 plate originally used for the derivation of the ASME KIC-reference curve [11]. Data show clear specimen size effect. Fig. 10 KIC results of the HSST 02 plate originally used for the derivation of the ASME KIC-reference curve [11]. Data size corrected to 25 mm specimen thickness. 20

22 In the construction of both the ASME reference curves as well as the Russian reference curves, the statistical size effect has not been accounted for. Due to this, it has not been possible to get a proper description of the true temperature dependence of the fracture toughness. New research indicates that a single expression (often called the "Master Curve") can be used to describe the fracture toughness temperature dependence of most ferritic structural steels, including both Russian and ASME type pressure vessel steels [13-15]. The results so far appear convincing, but more validation may be necessary to gain complete acceptance of the master curve. Another problem in both approaches is the reference curve indexing method. There is no scientific basis to use the indexing based on Tk or RTNDT for the description of static brittle fracture initiation toughness. The RTNDT temperature is generally, for the unirradiated materials, equal to the nil-ductility temperature NDT determined by the Pellini drop weight test. The NDT temperature is a measure of the materials crack arrest properties and therefore "mechanistically" it should form an effective indexing for crack arrest toughness. The ASME KIR curve is of course actually a lower bound crack arrest curve and for that purpose RTNDT is appropriate. However, the structural integrity analysis is performed with respect to static brittle fracture initiation and for indexing this, RTNDT is not as good [16]. The ASME approach assumes that the relation between static initiation toughness and crack arrest is fixed. In reality this is not necessarily the case. The relation is not only material dependent, it is also affected by irradiation as shown in Figs. 11 and 12 [17,18]. The Russian approach is based on a Charpy-V transition temperature indexation of the static fracture toughness. The indexation criterion is not fixed, but is dependent upon the materials yield strength. For higher strength materials a higher energy transition temperature is prescribed. In reality, the materials yield strength has only a minute effect upon the relation between static fracture toughness transition temperature and the Charpy-V transition temperature [16]. Based on a theoretical examination, of the differences in the fracture toughness test and the Charpy-V test, a commonly applicable relation, including also irradiated material, between the 28 J Charpy-V transition temperature and the 100 MPa m static fracture toughness temperature has been developed [16]. The correlation has an effective standard deviation of 13 C and it is the same for low and high strength steels. The correlation enables the calculation of the fracture toughness as a function of Charpy-V 28 J transition temperature, specimen thickness or crack front length and desired total cumulative failure probability. It can be expressed as K IC 20+ [11+77 exp{0.019(t -TK 28J < P >)}] ln 1- P where temperature is in C, thickness is in mm and fracture toughness is in MPa m. P is the desired cumulative total failure probability accounting also for the uncertainty in the Charpy-V - fracture toughness correlation. 1/4 25 B 1/4 21

23 22

24 Fig. 11 KJC and KIa for non-irradiated 73W [17,18]. KJC values size corrected to 25 mm specimen thickness. Results show clear difference between KJC and KIa. Fig. 12 KJC and KIa for irradiated 73W [17,18]. KJC values size corrected to 25 mm specimen thickness. Results show small difference between KJC and KIa. 23

25 When the relation between RTNDT, Tk and TK28J is combined with the above equation it is possible to compare the different reference fracture toughness curves and to study their intrinsic level of safety. Figs contain a comparison of the Russian approach reference fracture toughness curves with the statistically defined failure probability curves based on the commonly applicable correlation between Charpy-V (CVN) and fracture toughness (KIC). The failure probability curves have been calculated, corresponding to a crack front length of 100 mm and they include the uncertainty in the correlation. The failure probability curves have been shifted from TK28J, to the reference temperature Tk estimating the mean difference to be 10 C. The standard deviation of the difference is not expected to be more than 5 C. This additional uncertainty has not, however, been included in Figs For all three material cases, the normal operating conditions curves are close to, or below, the 1 % cumulative failure probability curve. Only in the case of the VVER-1000 base material, the lower shelf fracture toughness assumption appear somewhat high. The emergency situations curves lie between the % cumulative failure probability curves, again with the exception of the VVER-1000 lower shelf estimate. As a whole, the reference fracture toughness curves corresponding to welded joints appear to be least conservative. Fig. 13 Comparison of Russian approach reference fracture toughness curves (VVER-440 base material) with statistically defined failure probability curves based on a commonly applicable CVN-KIC correlation. 24

26 Fig. 14 Comparison of Russian approach reference fracture toughness curves (VVER-1000 base material) with failure probability curves based on a commonly applicable CVN-KIC correlation. Fig. 15 Comparison of Russian approach reference fracture toughness curves (welded joints) with statistically defined failure probability curves based on a commonly applicable CVN-KIC correlation. 25

27 Figs. 16 and 17 contain a comparison of the ASME and the French reference fracture toughness curves with the statistically defined failure probability curves based on the commonly applicable correlation between Charpy-V (CVN) and fracture toughness (KIC). The failure probability curves have been calculated, corresponding to a crack front length of 100 mm and they include the uncertainty in the correlation. The failure probability curves have been shifted from TK28J to the reference temperature RTNDT estimating the mean difference to be 7 C. The standard deviation of the difference is estimated to be more than 15 C. This additional uncertainty has not, however, been included into the figures. The ASME KIR curve lies close to, or below, the 1% cumulative failure probability curve. In this respect it is quite similar to the Russian approach normal operating conditions curves. On the lower shelf the ASME KIC reference curve is more conservative than the Russian approach emergency situations curves, but in the higher temperature regime the toughness response to temperature appear to be unrealistically high, thus decreasing the conservatism. The French approach reference fracture toughness curves are based upon the ASME reference curves, but applying additional safety factors. Thus they are more conservative than the normal ASME curves. If the safety factors applied in the ASME code would be used for the ASME reference curves (Fig. 4), these would become more conservative. Fig. 16 Comparison of the ASME reference fracture toughness curves with statistically defined failure probability curves based on a commonly applicable CVN-KIC correlation. 26

28 Fig. 17 Comparison of the French approach reference fracture toughness curves with statistically defined failure probability curves based on a commonly applicable CVN-KIC correlation. The above described method is the only one enabling the construction of statistically defined fracture toughness reference curves. The comparison reveal that the reference curves in the different national approaches are usually very conservative. 4.2 Reference temperature shift In all approaches, the reference temperature shift is based on the Charpy-V test. The Charpy-V test differs in two major aspects from the static fracture toughness test. First, it is a dynamic test and second, it does not measure a fracture initiation event, but the difficulty to cause complete fracture of the specimens. In mechanistic terms the test measures the combined effect of initiation and propagation. Thus, the Charpy-V transition temperature is affected partly by the materials dynamic initiation properties and partly by the materials dynamic ductile tearing resistance. The higher the transition temperature energy criterion, the more the result will be affected by the ductile tearing properties. Since irradiation affects both the relation between static and dynamic toughness as well as the ductile tearing properties, it is quite clear that the Charpy-V transition temperature shift cannot be relied on as an indicator of the static fracture toughness shift. Several investigations have revealed the discrepancies between the Charpy-V shift and the fracture toughness shift, the first and probably the most extensive, being the 27

29 one by Hiser [19]. In Fig. 18 the Hiser results are presented [19]. He found that the relation shows a considerable scatter. Also material type appeared to have an effect on the relation. Most importantly, the Hiser investigation shows that there appears to be an overall trend for the Charpy-V test to underestimate the fracture toughness shift. Underpredictions of more than 40 C are possible. The Hiser investigation indicate that welds are more likely to be conservatively predicted by the Charpy-V shift, but even for them the accuracy is not very good. Also the results from the HSSI 5 irradiation embrittlement research programme [18] indicated that for the welds tested (72W and 73W), the Charpy-V tests were capable of producing a conservative estimate of the static fracture toughness shift only when the additional temperature margin prescribed by the U. S. Regulatory Guide 1.99, Revision 2 [7] was added to the Charpy-V shift. Fig. 18 Comparison of transition temperature shifts from Charpy-V 41 J and KJC = 100 MPa m. Underprediction of TKJC can be larger than 40 C for TK41J [19]. The Russian approach prescribes the use of a higher energy criteria for higher yield strength material. If the irradiation induced yield strength increase is sufficient, the result will be the use of a higher energy criteria for the irradiated material than for reference material. This will result in a larger shift than if a constant energy criteria, like in the ASME approach, would be used. If the irradiation induced yield strength 28

30 increase is not sufficient, the result will be essentially identical to the ASME approach. The French approach leads to a result similar to the Russian approach. The French approach uses the shift in the 56 J impact energy transition temperature TK56J or the shift in the 0.9 mm lateral expansion transition temperature TK0.9mm, whichever is greater. The lateral expansion corresponding to a certain energy level, is affected by the material yield strength. Increasing the yield stress, makes plastic deformation of the specimen more difficult. Therefore, for the irradiated material TK0.9mm will correspond to a higher energy level than for the reference material. The result will be similar to the Russian approach. The French approach appears preferable, because it will always react to the yield strength increase. In the United States, the U.S. Regulatory Guide 1.99, Revision 2 [7], requires the use of an additional margin in the determination of RTNDT in order to make corrections for the inaccuracy of the Charpy-V shift. Such a margin is not prescribed in the European or the Russian approaches. An interesting detail may be that the Regulatory Guide prescribes a larger margin for the welds than for the base material and yet, the Hiser study (Fig. 18) showed base material to be more likely to be underpredicted. Despite the fact that the Russian and the French approaches may have a trend to yield less unconservative estimates of the fracture toughness shift than the ASME approach, their uncertainty in terms of the standard deviation is still anticipated to be approximately σ T 20 C. This uncertainty is not accounted for (no additional safety margins) in either the Russian or the European approaches. Because the fracture toughness irradiation shift is, in all approaches, estimated with far from optimum type of tests, it is practically impossible to quantitatively assess their accuracy and degree of safety. In order to ensure their conservatism, additional margins should be added to the estimated fracture toughness shift, like in the U.S. Regulatory Guide 1.99, Revision 2. This will, however, unduly penalize the majority of materials. An other possibility would be to show that for a certain type of material, the Charpy-V shift is conservative. This will, however, also penalize the material. Mechanistically the only sound solution is to measure the fracture toughness directly for the irradiated material. Recent advances allow broken Charpy-V specimen halves to be reconstituted to new Charpy-V size three point bend specimens, which can be used for a direct measurement of the fracture toughness [15]. Fig. 19 shows an example for the HSSI 5 series weld 73W. The large specimens were tested at ORNL as a part of the HSSI 5 program [18], whereas the small Charpy-V size specimens (B = 1O mm) have been reconstituted from a broken half of a larger specimen and tested at VTT Manufacturing Technology. The analysis of the results was identical to the one presented in [15]. For the analysis, all fracture toughness results were corrected to correspond to 25 mm specimen thickness applying the statistical thickness correction. Fig. 19 show that the fracture toughness can be determined reliably with Charpy-V size specimens. Thus, the major reason for using the Charpy-V test 29

31 (applicability of small specimen) seems unwarranted. The same specimens can likely be used to determine the fracture toughness directly, thus getting rid of any necessity to prescribe additional safety margins. It appears that the constraint speculations usually used to questionize small specimen fracture toughness estimates are strongly exaggerated. Fig. 19 Fracture toughness transition temperature T0 for 73W weld. T0 correspond to 100 MPa m mean fracture toughness for 25 mm crack front length. B = 10 mm => reconstituted Charpy-V specimens. 4.3 Chemistry factor In all approaches the chemistry factors have been developed by an empirical correlation of the material composition to the Charpy-V based transition temperature shift. Thus the scientific basis for all the different approaches is equally poor. Presently a large effort is being expended into trying to model irradiation embrittlement in order to be able to develop a micromechanism based chemistry factor expression. This work has already given an improved understanding of the embrittlement mechanisms. However, as long as the simple Charpy-V transition temperature is used as reference, the chemistry factors can never be quantitatively reliable and much of the research effort may be performed in vain. 5 CONCLUSIONS Based on the comparison, of the scientific basis of Russian and European 30

32 approaches for evaluating irradiation effects in reactor pressure vessels, the following conclusions can be drawn: - The Russian and European approaches have similar scientific bases, but mechanistically neither approach is optimal for evaluating the irradiation effects. In both cases the approaches are based on the Charpy-V test that is not directly descriptive of the irradiated material's true fracture toughness. At the same time highly conservative reference fracture toughness curves are usually applied. - The comparison emphasizes the need for an improved approach for evaluating irradiation effects in reactor pressure vessels. The new approach should be based upon a direct determination of the fracture toughness, or a validated correlation based on a direct determination of the fracture toughness, combined with a mechanistic treatment of irradiation and material variables. This is the subject matter of Task Group C of AMES Project 1. REFERENCES 1 Strength Analysis Standards for Equipment and Piping at Nuclear Power Plants, PNAE G Section III, Nuclear Power Plant Component, Division I, ASME Boiler and Pressure Vessel Code, American Society of Mechanical Engineers, New York, Section XI, ASME Boiler and Pressure Vessel Code, Rules for Inservice Inspection of Nuclear Power Plant Components, American Society of Mechanical Engineers, New York, White Paper on Reactor Vessel Integrity Requirements for Level A and B Conditions, EPRI TR , In-service Inspection Rules for Mechanical Components of PWR Nuclear Islands, RSEM, EDF/afcen, Design and Construction Rules for Mechanical Components of PWR Nuclear Islands, RCC-M Code, U. S. Nuclear Regulatory Commission, Regulatory Guide 1.99, Radiation Embrittlement of Reactor Vessel Materials, Revision 2, Washington U. S. Nuclear Regulatory Commission, Regulatory Guide 1.99, Effects of Residual Elements on Predicted Radiation Damage to Reactor Vessel Materials, Revision 1, Washington Sicherheitstechnische Regel des KTA, Überwachung der Strahlenversprödung 31

33 von Werkstoffen des Reactordruckbehälters von Leichtwasserreaktoren, KTA 3203, Fassung 3/84, Kaufman, J. G., "Experience in Plane-Strain Fracture Toughness Testing Per ASTM Method E 399," Developments in Fracture Mechanics Test Methods Standardization, ASTM STP 632, W.F. Brown, Jr., and J.G. Kaufman, Eds., American Society for Testing and Materials, Philadelphia, 1977, pp Flaw Evaluation Procedures: Background and Application of ASME Section XI Appendix A, EPRI NP-719-SR, T.U. Marston, Ed., Electric Power Research Institute, Palo Alto, 1978, pp. C8 - C Wallin, K., The Size Effect in KIC Results, Engineering Fracture Mechanics, Vol. 22, No. 1, 1985, pp Wallin, K., Irradiation Damage Effects on the Fracture Toughness Transition Curve Shape for Reactor Pressure Vessel Steels, International Journal of Pressure Vessels and Piping, Vol. 55, No. 1, 1993, pp Wallin, K., Törrönen, K., Ahlstrand, R., Timofeev, B., Rybin, V., Nikolaev, V. and Morozov, A., Theory Based Statistical Interpretation of Brittle Fracture Toughness of Reactor Pressure Vessel Steel 15X2MΦA and its Welds, Nuclear Engineering and Design, Vol. 135, 1992, pp Wallin, K., Validity of Small Specimen Fracture Toughness Estimates Neglecting Constraint Corrections, Constraint Effects in Fracture: Theory and Applications, ASTM STP 1244, M. Kirk and A. Bakker Eds., American Society for Testing and Materials, Philadelphia, 1994, In press. 16 Wallin, K., A Simple Theoretical Charpy-V - KIC Correlation for Irradiation Embrittlement, Innovative Approaches to Irradiation Damage and Fracture Analysis, PVP-Vol. 170, D.L. Marriott et. al. Eds., The American Society of Mechanical Engineers, 1989, pp Iskander, S. K., Corwin, W. R. and Nanstad, R. K., Effetcts of Irradiation on Crack-Arrest Toughness of Two High-Copper Welds, Effetcts of Radiation on Materials: 15th International Symposium, ASTM STP 1125, R. E. Stoller, A. S. Kumar and D. S. Gelles, Eds., American Society for Testing and Materials, Philadelphia 1992, pp Nanstad, R.K., Haggag, F.M., McCabe, D.E., Iskander, S.K., Bowman, K.O. and Menke, B.H., Irradiation Effects on Fracture Toughness of Two High-Copper Submerged-Arc Welds, HSSI Series 5: Main Report and Appendices A, B, C and D, NUREG/CR-5913, ORNL/TM-12165/V1, Vol. 1, Oak Ridge National Laboratory, Oak Ridge, 1992, 133 p p. app. 19 Hiser, A.L., Correlation of Cv and KIC/KIC Transition Temperature Increases Due to Irradiation, NUREG/CR-4395, MEA-2086, Materials Engineering Associates, Inc., 1985, 46 p. + app. 32