LIMITED DISTRIBUTION

COMPARISON OF THE SCIENTIFIC BASISOF RUSSIAN AND EUROPEAN

APPROACHES FOR EVALUATINGIRRADIATION EFFECTS IN

REACTOR PRESSURE VESSELS

Kim Wallin

European Network onAgeing Materials Evaluation and Studies

Espoo, December 1994

VTT Manufacturing TechnologyP.O. Box 1704, FIN-02044 VTT, Finland

Tel. 90-4561, Telefax 90-456 7002

1

ABSTRACT

Irradiation affects the properties of reactor pressure vessel material. Differentcountries use different approaches for evaluating these irradiation effects. Becausethe approaches are different, also the accuracy of them is likely to vary. In thisreport the scientific basis of Russian and European approaches for evaluatingirradiation effects in reactor pressure vessels have been compared. Asrepresentative European approaches, the ASME and French methodologies havebeen examined.

It can be concluded that the Russian and European approaches have similarscientific bases, but mechanistically neither approach is optimal for evaluating theirradiation effects. The approaches are based upon the Charpy-V test that is notdirectly descriptive of the materials true fracture toughness. At the same timehighly conservative reference fracture toughness curves are usually applied.

The comparison emphasizes the need for a new improved approach for evaluatingirradiation effects in reactor pressure vessels. The new approach should be basedupon a direct determination of the fracture toughness, or a validated correlationbased on a direct determination of the fracture toughness, combined with amechanistic treatment of irradiation and material variables. This is the subjectmatter of Task Group C of AMES Project 1.

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PREFACE

This report has been compiled at the VTT Manufacturing Technology as part of astate-of-the-art review on irradiation embrittlement, surveillance and mitigationmethods carried out by the European Network for Ageing Materials Evaluationand Studies (AMES). The work is financed by CEC DG XI through the contract"Centre d'Etudes de Saclay 91191 Gif Sur Yvette Cedex, Ref. 7220 3B018750, M.Soulat". Additional financing was provided by the Finnish Centre for Radiation andNuclear Safety (STUK) and VTT.

Associated with this subject, reports are prepared under the auspice of AMES onthermal annealing of irradiation effects and the evaluation of the available andpossible mitigation methods. The author acknowledges his gratitude to Acad.Myrddin Davies, Dr. Colin English, Dr. Karen Gott and Dr. Pierre Petrequin forconducting a peer review for this report.

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CONTENTS

ABSTRACT..........................................................................................................................1

PREFACE .............................................................................................................................2

1 INTRODUCTION.........................................................................................................4

2 RUSSIAN APPROACH..................................................................................................42.1 Background...........................................................................................................42.2 Fracture toughness reference curves .................................................................52.3 Critical brittleness temperature..........................................................................52.4 Chemistry factor.................................................................................................102.5 Surveillance methodology.................................................................................10

3 EUROPEAN APPROACHES........................................................................................10

3.1 Background.........................................................................................................113.2 Fracture toughness reference curves ...............................................................113.3 Reference temperature ......................................................................................133.4 Chemistry factor.................................................................................................143.5 Surveillance methodology.................................................................................17

4 MECHANISTIC INTERPRETATION OF THE APPROACHES...................................17

4.1 Fracture toughness reference curves ...............................................................174.2 Reference temperature shift..............................................................................274.3 Chemistry factor.................................................................................................30

5 CONCLUSIONS .........................................................................................................30REFERENCES.....................................................................................................................31

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1 INTRODUCTION

Neutron irradiation affects the properties of reactor pressure vessel material. Thematerial changes that may affect the use of the pressure vessel are connected to thematerials fracture resistance. As a rule irradiation reduces the materials fractureresistance and therefore it is important to know the rate of reduction as a function ofirradiation dose. Different countries apply different approaches for evaluating theseirradiation effects. Because the approaches are different, also the accuracy of them islikely to vary. The largest differences between approaches for evaluating irradiationeffects in pressure vessels are between Eastern (Russian) and Western (European)approaches. The safety of Central and Eastern European reactors have recently beenthe subject of close scrutiny by international experts. The scientific basis of thedifferent approaches for evaluating irradiation effects in reactor pressure vesselsneed to be compared.

2 RUSSIAN APPROACH

The Russian approach for evaluating irradiation effects in reactor pressure vessels isapplied in the case of all VVER-440 and VVER-1000 type reactor pressure wesselsin the Central and Eastern European countries [1].

2.1 Background

The safety assessment of reactor pressure vessels is based upon linear elastic fracturemechanics (LEFM). The reactor pressure vessel lifetime is thus determined solely byit's resistance to brittle fracture. The fracture mechanical material propertydescribing the (LEFM) fracture resistance is denoted fracture toughness (KIC). KICdescribes the materials resistance towards brittle fracture initiation under static or"quasistatic" loading.

KIC is related to temperature, and irradiation changes this relation. The assessmentthus requires knowledge regarding the change of KIC as a function of bothtemperature and irradiation. In the Russian approach the KIC for the assessment isnot determined directly, but instead a reference curve methodology is used. It isassumed that the temperature dependence of fracture toughness is not affected byirradiation enabling the fracture toughness temperature dependence to be describedby a single curve. Irradiation is assumed only to shift the location of the curve tohigher temperatures. Thus the irradiation effects in pressure vessels are evaluated byan estimation of this temperature shift. In the Russian approach the shift is eitherdetermined from Charpy-V impact tests or from the chemical composition of thematerial, applying a special chemistry factor.

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2.2 Fracture toughness reference curves

The Russian approach applies material specific reference curves (different forVVER-440 base material, VVER-1000 base material and their welds). The referencecurves are based upon KIC test results corresponding to the non-irradiated materialstate, measured by comparatively large test specimens. Each material has differentcurves for 1) normal operating conditions, 2) operational occurrences and hydraulictests and 3) emergency situations. The initial reference curve, corresponding toemergency situations, is based on an "eye-ball" lower envelope of experimental data.The curves for normal operating conditions and operational occurrences are derivedas the lower envelope of two curves determined on the basis of the initial curve. Oneof the curves is derived by dividing the ordinates of the initial curve by a safetyfactor nk {1) → nk = 2, 2) → nk = 1.5}, while the other is derived by shifting theinitial curve along the x-axis by an amount ∆T {1) & 2) → ∆T = 30°C}. Themethod of construction forces the 1) and 2) curves together at higher toughnesslevels.

The curves are given in terms of an effective temperature (T-Tk), where Tk denotes acritical brittleness temperature. The three different initial curves (emergencysituations) are:

VVER-440 base materials (12X2MΦPA, 15X2MΦA and 15X2MΦA-A) →→→→

[KI]3 = 35 + 45·exp{0.02·(T-Tk)} →→→→

VVER-1000 base materials (15X2HMΦA and 15X2HMΦA-A) →→→→

[KI]3 = 74 + 11·exp{0.0385·(T-Tk)}

VVER-440 and VVER-1000 beltline welds (15X2MΦA, 15X2MΦA-A,15X2HMΦA and 15X2HMΦA-A) →→→→

[KI]3 = 35 + 53·exp{0.0217·(T-Tk)}

The Russian reference fracture toughness curves are presented graphically in Figs.1-3.

2.3 Critical brittleness temperature

The critical brittleness temperature Tk forms the essence of the Russian approach forevaluating irradiation effects in reactor pressure vessels. The temperature isdetermined from Charpy-V impact test results. The definition for Tk is not howeverfixed, 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 themaximum value for only two tests. Impact tests are performed at differenttemperatures close to the expected critical brittleness temperature. Conservativeestimates of Tk for the different materials are given in the code [1]. The criterion forTk is determined by a combination of absorbed energy required (impact toughness)

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and 50% ductile fracture appearance determined from the broken specimen fracturesurface. A preliminary determination of Tk is first performed using specimens at afew test temperatures. Based upon the preliminary estimate, additional tests areperformed at surrounding temperatures in order to make the estimation moreaccurate.

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Fig. 1 Russian approach reference fracture toughness curves for VVER-440base materials. Temperature normalized by critical brittlenesstemperature.

Fig. 2 Russian approach reference fracture toughness curves for VVER-1000base materials. Temperature normalized by critical brittlenesstemperature.

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A minimum of three tests is required at each temperature. If any of the requirementsset for the minimum value is not met on one of the three test specimens, one isallowed to test three additional specimens at that temperature. Thus if five tests outof 6 fulfil the requirements, the results are acceptable. Tk is taken to be the lowesttest temperature at (and above) which a fulfilment of the criteria in table 1 areachieved.

Table 1. Criteria for the definition of Tk.

Yieldstress20°C(MPa)

Meanimpacttoughness(J)

Minimumimpacttoughness(J)

Tk+30oCmean impacttoughness(J)

Tk+30oCmin. impacttoughness(J)

Tk+30oCminimumductile(%)

< 304 23 16 35 25 50

< 402 31 22 47 33 50

< 549 39 27 59 41 50

< 687 47 33 71 50 50

Fig. 3 Russian approach reference fracture toughness curves for VVER-440and VVER-1000 welds. Temperature normalized by critical brittlenesstemperature.

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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, butat least 12 specimens must be tested. In this case the test data is fitted, applying theleast square method, by the equation

where KCV is the impact energy, A is the average impact energy between the uppershelf energy (KCVmax) and the lower shelf energy (KCVmin), B =(KCVmax-KCVmin)/2, T0 is the temperature corresponding to A and C is a constantdescribing the steepness of the impact energy temperature dependence. For thedetermination 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 toestimate the materials fracture toughness. If the neutron fluency to which Tk refers,is relevant, the experimentally determined Tk can be used as such. Otherwise thechange 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 determinedby the equation

where TkF is the critical brittleness temperature after irradiation and TkI is the initialcritical brittleness temperature. The irradiation shift is needed to determine theradiation embrittlement coefficient.

Radiation embrittlement coefficient

Based upon the irradiation shift (∆TF) the radiation embrittlement coefficient AF

can be determined from:

where Fn is the neutron fluency with E > 0.5 MeV, F0 = 1022 neutrons/m2 and n is aconstant. If data corresponding to several different fluences is available, the power ncan be determined experimentally, otherwise a value of 1/3 is assumed. The validityof the above equation is limited to a fluency range of 1022 < Fn < 1024 neutrons/m2.The radiation embrittlement coefficient enables the determination of ∆TF and thusTk for any fluency within the validity range of the equation.

•

CT-T B+A = KCV 0tanh

T-T = T kIkFF∆

•∆

FF T = A

0

nn

FF

10

2.4 Chemistry factor

The radiation embrittlement coefficient AF can also be estimated indirectly basedupon the materials chemistry. In this context the radiation embrittlement coefficientis also called the chemistry factor and it is used for vessels that do not have aspecific surveillance program. For base materials and welds corresponding toVVER-1000 reactors, the Russian code gives fixed (material type dependent) valuesfor the chemistry factor. For VVER-440 type reactor welds, the chemistry factor hasbeen determined on the basis of an empirical treatment of Charpy-V impact data(∆TF), obtaining an estimate of the dependence of radiation embrittlement againstthe content of phosphorus and copper. The chemistry factor has for VVER-440 typereactor welds the form

AF = 800-(P[%] + 0.07-Cu[%]) for irradiation temperature 270°°°°C

and

AF = 800-(P[%] + 0.07-Cu[%]) + 8 for irradiation temperature 250°°°°C.

2.5 Surveillance methodology

Older vessels do not have a surveillance programme, and for them the chemistryfactor concept is applied. In this case characteristic (conservative) values are givenin the code both for TkI and AF [1].

The surveillance methodology for newer vessels is based upon Charpy-V and tensiletesting, by which the critical brittleness temperature is determined. No fracturetoughness testing is prescribed, only the use of the fracture toughness referencecurves.

3 EUROPEAN APPROACHES

Most European approaches for evaluating irradiation effects in reactor pressurevessels 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 analysisis carried out [4]. Many countries apply the ASME code directly, but in somecountries like France [5,6] the approaches have, however, experienced a nationalvariation.

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3.1 Background

The ASME code is based upon linear elastic fracture mechanics, similar to theRussian code. It too uses a reference curve methodology for estimating the materialsfracture toughness, but the scientific basis is somewhat different. The Frenchapproach originates from the ASME methodology, but it has been made moreflexible, to account for possible plasticity effects. Both approaches assume that thetemperature dependence of fracture toughness is not affected by irradiation enablingthe fracture toughness temperature dependence to be described by a single curve. Asin the Russian approach, the shift is either determined based upon Charpy-V impacttests 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 inthe Russian approach. It is additionally assumed that the difference between staticinitiation and crack arrest is constant. Both curves are given in the form of aneffective temperature (T-RTNDT), where RTNDT denotes the Nil-DuctilityReference Temperature (see 3.3). For the fracture mechanical assessment of normaloperation conditions the crack arrest curve also describes crack initiation and it isthen denoted KIR. The ASME code only gives the reference curves in a graphicalform, but the French approach gives descriptive equations for them as follows [6]

36.5+3.1·exp [0.036·(T-RTNDT+55.5)]KIC = min.

220 MPa√√√√m

and

29.43+1.355·exp [0.0261·(T-RTNDT+88.9)]KIR = min.

220 MPa√√√√m

where units are in MPa√m and °C.

The reference curves have been developed based upon empirical "eye ball" lowerenvelope curve fitting to experimental LEFM unirradiated fracture toughness data.

The KIR curve is intended to describe normal operation conditions and the KIC curveemergency and faulted conditions. In addition to the different reference curves,safety factors are applied. For normal operation conditions a safety factor of 10 is

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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 2for the allowable crack size or √2 for KI or KIC. The ASME reference curves arepresented graphically in Fig. 4.

The French approach

The French approach applies essentially the ASME fracture toughness referencecurves, but the safety factors are different and also the definition of "upper shelf"differs from ASME. The approach recognizes three different situations withdifferent safety factors: 1) LEVEL A; normal and upset, 2) LEVEL C; emergencyand 3) LEVEL D; faulted conditions. The upper shelf toughness denoted as KJC is, ifnot determined directly from an elastic plastic JIC-test, 165 MPa√m up to atemperature of +150 °C and 150 Mpa√m in the temperature range +150°C...+350°C

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}.

Fig. 4 The ASME fracture toughness reference curves and curvescorresponding to normal and emergency situations. Temperaturenormalized by the Nil-Ductility Reference Temperature.

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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-∆acurves (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.

3.3 Reference temperature

The Nil-Ductility Reference Temperature (RTNDT) is determined from thenil-ductility temperature (NDT) and the Charpy-V impact test. The NDTtemperature is determined in accordance with the ASTM Method for ConductingDrop-Weight Test to Determine Nil-Ductility Transition Temperature of Ferritic

Fig. 5 The French fracture toughness reference curves. Temperaturenormalized by the Nil-Ductility Reference Temperature.

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Steels (E 208). If the minimum impact energy and lateral expansion in the Charpy-Vtest (3 specimens) are at least 68 J and 0.9 mm respectively, at a temperature equalto NDT + 33 °C, the NDT temperature is taken to represent RTNDT. If the Charpy-Vproperties do not meet the above criterion, RTNDT is taken as the temperature atwhich the requirements are reached, minus 33°C. Normally, for modern steelsRTNDT is equal to NDT.

The ASME approach for determining irradiation shift

The NDT test is not included in the surveillance programs. The irradiation inducedshift of the RTNDT reference temperature is defined as the shift in the 41 J impactenergy transition temperature TK41J. It is assumed that the true static fracturetoughness shift and crack arrest shift is equal to or less than ∆TK4IJ (∆RTNDT). TheU.S. Regulatory Guide 1.99, Revision 2 [7], requires the use of an additional marginin the determination of ∆RTNDT. For welds the margin is 31°C and for base metal 19°C. In both cases the margin is not, however, more than ∆RTNDT/2. This margin isput on top of the experimental ∆RTNDT. The margin is not directly prescribed in theASME approach, nor in European approaches equivalent to ASME.

Based upon ∆RTNDT the fluence dependence can be calculated by the equation [7]

∆RTNDT = (CF)·f(0.28-0.10·log f)

where CF is the radiation embrittlement coefficient and f is the neutron fluence(·10-19 n/cm2) with E > 1 MeV. When two or more credible surveillance data sets areavailable, the radiation embrittlement coefficient is determined from theexperimental data by least square fitting.

The French approach for determining irradiation shift

The French approach differs from the ASME methodology. The irradiation inducedshift of the RTNDT reference temperature is defined as the shift in the 56 J impactenergy transition temperature TK56J or the shift in the 0.9 mm lateral expansiontransition temperature TK0.9mm, whichever is greater. The lateral expansioncorresponding 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 ∆TK0.9mm. The French approach doesnot apply additional safety margins.

The French codes [5,6] allow the fluence dependence of ∆RTNDT to be determinedexperimentally, but do not give any recommendations for type of expression to beused. In this respect the French approach is more flexible than other approaches.

3.4 Chemistry factor

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The application of the chemistry factor is similar to the Russian approach. TheEuropean approaches also apply the chemistry factor concept in an effort todetermine the irradiation shift directly from the steel chemistry. For the ASMEmethodology type steels, the U.S. Regulatory Guide 1.99, Revision 2 [7] givesdifferent chemistry factors for welds and base metal in a tabulated form. The tablesare presented graphically in Figs. 6 and 7. The tables include only the effect ofcopper and nickel. This does not mean that investigations would have shownphosphorus and sulphur to have no effect, but that the materials for which thechemistry factors have been developed have had a comparatively constant level ofthese elements. This means that the applicability of the chemistry factor is restrictedto a special population of materials. Even when ∆RTNDT is calculated with thechemistry factor, one is required to use the additional safety margin prescribed in theRegulatory Guide [7].

Fig. 6 Reg. Guide 1.99, Rev. 2 Chemistry Factor for welds.

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The French codes from 1988 contain a chemistry factor equation of the form

where the terms (%P-0.008) and (%Cu-0.08) become zero for P < 0.008 % and Cu <0.08 %.

The equation is applicable in a fluence range 1018...6·1019 n/cm2 and for irradiationtemperatures between 275°C...300°C. The equation originates from the U.S.Regulatory Guide 1.99, Revision 1 [8]. Essentially the same equation is used (in agraphical form having a fixed % P) in the German KTA code [9], but only forselecting materials to be included into the surveillance program.

The French approach includes also a newer so called FIS equation giving an upperbound estimate of the chemistry factor [5]

where the terms (%P-0.008) and (%Cu-0.08) become zero for P < 0.008 % and Cu <0.08 %.

Fig. 7 Reg. Guide 1.99, Rev. 2 Chemistry Factor for base metal.

]10/ [f0.008))-(%P2278+0.08)-(%Cu556+[22 = RT 1/219NDT •••∆

•

10Cu)Ni191+0.08)-238(%Cu+0.008)-1537(%P+(24+8 = C)( FIS

19

0.352o φ

17

3.5 Surveillance methodology

The surveillance methodology in Europe varies. Usually, the surveillancemethodology is based upon Charpy-V testing, from which the RTNDT is determined.The French methodology includes also fracture toughness specimens, but they aremainly used to determine the effect of irradiation on the ductile fracture properties.

4 MECHANISTIC INTERPRETATION OF THEAPPROACHES

Mechanistically the Russian and European approaches for evaluating irradiationeffects in reactor pressure vessels are very similar. All the different approaches relyon the Charpy-V impact test to determine the effect of irradiation on the materialsfracture toughness. In none of the approaches is the fracture toughness determineddirectly. The reason for this is that originally there was no fracture toughness testingstandard that would be suitable for surveillance testing. Therefore the differentapproaches have had to apply the Charpy-V test, combined with referencetemperature and reference fracture toughness concepts. Unfortunately, theseconcepts have never been comprehensively, experimentally or theoretically,validated for irradiation embrittlement. This does of course not mean that theconcepts would necessarily lead to an unconservative estimate of the materialsfracture toughness.

4.1 Fracture toughness reference curves

All the fracture toughness reference curves are based on linear-elastic fracturemechanical tests on relevant materials. In all cases, however, the fracture toughnesshas been determined only for unirradiated material. The crack arrest and dynamicfracture toughness tests used to determine the ASME KIR curve (KIa) were notperformed according to any testing standard, because there did not exist one at thetime. This was not considered a problem at the time because the reference fracturetoughness curves are essentially intended for the description of static brittle fractureinitiation toughness. This toughness was determined according to the ASTMstandard E 399 called, Standard Test Method for Plane-Strain Fracture Toughness ofMetallic Materials. The Russian fracture toughness tests, used in their referencecurve development, were performed and analysed according to the corresponding,practically identical, Russian GOST standard. Thus the mechanical basis for thefracture toughness reference curves is the same.

The standard E 399, denotes the valid measured entity as plane strain fracturetoughness. The definition of plane strain in the standard is not, however, directlyrelated 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.

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Originally, in the development of the test standard, high strength steel andaluminium and titanium alloys were used. These materials fail typically by a ductile,strain controlled mechanism having an increasing tearing resistance as a function ofcrack growth. For these materials, typically, specimens with small in-planedimensions yield lower toughness values than specimens with large in-planedimension (Fig. 8 [10]). The specimen thickness does not have a significant effectfor these materials (Fig. 8 [10]). Plane-strain was in the standard developmentdefined as a criterion guaranteeing essentially specimen size insensitive toughnessvalues. Usually, for those materials however, the value was not a minimum but amaximum. This does not directly imply that valid KIC values for pressure vesselsteel displaying brittle fracture, would not represent plane-strain. It only means thatthe definition in the standard is unrelated to the true stress state. Most importantly,valid KIC values guarantee that the specimen behaves macroscopically in alinear-elastic manner.

The problem with valid KIC testing of pressure vessel steels is that the validspecimen size is dependent on the fracture toughness. This has led to the use ofsmall specimens in the low toughness temperature regime and large specimens inthe higher temperature regime. This would not be a problem if the valid fracturetoughness were also specimen size independent also in the case of brittle fracture.Unfortunately, pressure vessel steels, showing brittle fracture, were not included inthe original test development work. New experimental and theoretical work haveshown that brittle fracture initiation toughness in reality shows a specimen sizeeffect, such that large specimens yield lower values than small specimens.

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This size effect is active also for valid KIC values. Fig. 9 shows KIC data for theHSST 02 plate (A533B Cl.l) originally forming the major part of the data set leadingto the ASME KIC reference curve [11]. For clarity the smallest specimens (tested atlower shelf temperatures), and the largest specimens (showing inhomogeneityeffects) have been omitted. A clear difference between the 50 mm thick specimensand the 100-150 mm thick specimens can be seen. The temperature T0 reported inthe figure, denotes the temperature where the mean fracture toughness is 100MPa√m. Mechanistically the size effect can be explained by a statistical samplingeffect due to an increased crack front length (specimen thickness). The statisticalsize effect can be expressed in the form [12]

where B1 and B2 are two different specimen thicknesses and Kmin is a lower limitingfracture toughness, that for pressure vessel steels can be approximated by Kmin = 20MPa√m. Fig. 10 shows the effect of the size correction on the HSST 02 data fromFig. 9.

Fig. 8 Plane-strain fracture toughness of 75 mm thick 2219-T851 plate(material typical for the development of E 399) [10]. Results indicateincreasing fracture toughness with increasing specimen in-planedimensions.

•

BB )K-K( + K = K

2

11/4

BB 12 minmin

20

Fig. 9 KIC results of the HSST 02 plate originally used for the derivation ofthe ASME KIC-reference curve [11]. Data show clear specimen sizeeffect.

Fig. 10 KIC results of the HSST 02 plate originally used for the derivation ofthe ASME KIC-reference curve [11]. Data size corrected to 25 mmspecimen thickness.

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In the construction of both the ASME reference curves as well as the Russianreference curves, the statistical size effect has not been accounted for. Due to this, ithas not been possible to get a proper description of the true temperature dependenceof the fracture toughness. New research indicates that a single expression (oftencalled the "Master Curve") can be used to describe the fracture toughnesstemperature dependence of most ferritic structural steels, including both Russian andASME type pressure vessel steels [13-15]. The results so far appear convincing, butmore validation may be necessary to gain complete acceptance of the master curve.

Another problem in both approaches is the reference curve indexing method. Thereis no scientific basis to use the indexing based on Tk or RTNDT for the description ofstatic brittle fracture initiation toughness.

The RTNDT temperature is generally, for the unirradiated materials, equal to thenil-ductility temperature NDT determined by the Pellini drop weight test. The NDTtemperature 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 forthat purpose RTNDT is appropriate. However, the structural integrity analysis isperformed with respect to static brittle fracture initiation and for indexing this,RTNDT is not as good [16]. The ASME approach assumes that the relation betweenstatic initiation toughness and crack arrest is fixed. In reality this is not necessarilythe case. The relation is not only material dependent, it is also affected by irradiationas shown in Figs. 11 and 12 [17,18].

The Russian approach is based on a Charpy-V transition temperature indexation ofthe static fracture toughness. The indexation criterion is not fixed, but is dependentupon the materials yield strength. For higher strength materials a higher energytransition temperature is prescribed. In reality, the materials yield strength has only aminute effect upon the relation between static fracture toughness transitiontemperature and the Charpy-V transition temperature [16]. Based on a theoreticalexamination, of the differences in the fracture toughness test and the Charpy-V test,a commonly applicable relation, including also irradiated material, between the 28 JCharpy-V transition temperature and the 100 MPa√m static fracture toughnesstemperature has been developed [16]. The correlation has an effective standarddeviation of 13 °C and it is the same for low and high strength steels. Thecorrelation enables the calculation of the fracture toughness as a function ofCharpy-V 28 J transition temperature, specimen thickness or crack front length anddesired total cumulative failure probability. It can be expressed as

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 theuncertainty in the Charpy-V - fracture toughness correlation.

≈

B25

P-11 >)}]P-0.5<13-18+TK-{0.019(T77+[11+20 K

1/41/4

28JIC lnexp

22

23

Fig. 11 KJC and KIa for non-irradiated 73W [17,18]. KJC values size correctedto 25 mm specimen thickness. Results show clear difference betweenKJC and KIa.

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

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When the relation between RTNDT, Tk and TK28J is combined with the aboveequation it is possible to compare the different reference fracture toughness curvesand to study their intrinsic level of safety.

Figs. 13-15 contain a comparison of the Russian approach reference fracturetoughness curves with the statistically defined failure probability curves based on thecommonly applicable correlation between Charpy-V (CVN) and fracture toughness(KIC). The failure probability curves have been calculated, corresponding to a crackfront length of 100 mm and they include the uncertainty in the correlation. Thefailure probability curves have been shifted from TK28J, to the reference temperatureTk estimating the mean difference to be 10°C. The standard deviation of thedifference is not expected to be more than 5°C. This additional uncertainty has not,however, been included in Figs. 13-15. For all three material cases, the normaloperating conditions curves are close to, or below, the 1 % cumulative failureprobability curve. Only in the case of the VVER-1000 base material, the lower shelffracture toughness assumption appear somewhat high. The emergency situationscurves lie between the 10 - 25 % cumulative failure probability curves, again withthe exception of the VVER-1000 lower shelf estimate. As a whole, the referencefracture toughness curves corresponding to welded joints appear to be leastconservative.

Fig. 13 Comparison of Russian approach reference fracture toughness curves(VVER-440 base material) with statistically defined failure probabilitycurves based on a commonly applicable CVN-KIC correlation.

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Fig. 14 Comparison of Russian approach reference fracture toughness curves(VVER-1000 base material) with failure probability curves based on acommonly applicable CVN-KIC correlation.

Fig. 15 Comparison of Russian approach reference fracture toughness curves(welded joints) with statistically defined failure probability curvesbased on a commonly applicable CVN-KIC correlation.

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Figs. 16 and 17 contain a comparison of the ASME and the French referencefracture toughness curves with the statistically defined failure probability curvesbased on the commonly applicable correlation between Charpy-V (CVN) andfracture toughness (KIC). The failure probability curves have been calculated,corresponding to a crack front length of 100 mm and they include the uncertainty inthe correlation. The failure probability curves have been shifted from TK28J to thereference temperature RTNDT estimating the mean difference to be 7 °C. Thestandard deviation of the difference is estimated to be more than 15 °C. Thisadditional uncertainty has not, however, been included into the figures.

The ASME KIR curve lies close to, or below, the 1% cumulative failure probabilitycurve. In this respect it is quite similar to the Russian approach normal operatingconditions curves. On the lower shelf the ASME KIC reference curve is moreconservative than the Russian approach emergency situations curves, but in thehigher temperature regime the toughness response to temperature appear to beunrealistically high, thus decreasing the conservatism. The French approachreference fracture toughness curves are based upon the ASME reference curves, butapplying additional safety factors. Thus they are more conservative than the normalASME curves. If the safety factors applied in the ASME code would be used for theASME reference curves (Fig. 4), these would become more conservative.

Fig. 16 Comparison of the ASME reference fracture toughness curves withstatistically defined failure probability curves based on a commonlyapplicable CVN-KIC correlation.

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The above described method is the only one enabling the construction of statisticallydefined fracture toughness reference curves. The comparison reveal that thereference 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. TheCharpy-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 termsthe test measures the combined effect of initiation and propagation.

Thus, the Charpy-V transition temperature is affected partly by the materialsdynamic initiation properties and partly by the materials dynamic ductile tearingresistance. The higher the transition temperature energy criterion, the more the resultwill be affected by the ductile tearing properties. Since irradiation affects both therelation between static and dynamic toughness as well as the ductile tearingproperties, it is quite clear that the Charpy-V transition temperature shiftcannot be relied on as an indicator of the static fracture toughness shift.

Several investigations have revealed the discrepancies between the Charpy-V shiftand the fracture toughness shift, the first and probably the most extensive, being the

Fig. 17 Comparison of the French approach reference fracture toughnesscurves with statistically defined failure probability curves based on acommonly applicable CVN-KIC correlation.

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one by Hiser [19]. In Fig. 18 the Hiser results are presented [19]. He found that therelation shows a considerable scatter. Also material type appeared to have an effecton the relation. Most importantly, the Hiser investigation shows that there appears tobe an overall trend for the Charpy-V test to underestimate the fracture toughnessshift. Underpredictions of more than 40°C are possible. The Hiser investigationindicate that welds are more likely to be conservatively predicted by the Charpy-Vshift, but even for them the accuracy is not very good. Also the results from theHSSI 5 irradiation embrittlement research programme [18] indicated that for thewelds tested (72W and 73W), the Charpy-V tests were capable of producing aconservative estimate of the static fracture toughness shift only when the additionaltemperature margin prescribed by the U. S. Regulatory Guide 1.99, Revision 2 [7]was added to the Charpy-V shift.

The Russian approach prescribes the use of a higher energy criteria for higher yieldstrength material. If the irradiation induced yield strength increase is sufficient, theresult will be the use of a higher energy criteria for the irradiated material than forreference 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

Fig. 18 Comparison of transition temperature shifts from Charpy-V 41 J andKJC = 100 MPa√m. Underprediction of ∆TKJC can be larger than 40°Cfor ∆TK41J [19].

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increase is not sufficient, the result will be essentially identical to the ASMEapproach.

The French approach leads to a result similar to the Russian approach. The Frenchapproach uses the shift in the 56 J impact energy transition temperature TK56J or theshift in the 0.9 mm lateral expansion transition temperature TK0.9mm, whichever isgreater. The lateral expansion corresponding to a certain energy level, is affected bythe material yield strength. Increasing the yield stress, makes plastic deformation ofthe specimen more difficult. Therefore, for the irradiated material TK0.9mm willcorrespond to a higher energy level than for the reference material. The result will besimilar to the Russian approach. The French approach appears preferable, because itwill always react to the yield strength increase.

In the United States, the U.S. Regulatory Guide 1.99, Revision 2 [7], requires theuse of an additional margin in the determination of ∆RTNDT in order to makecorrections for the inaccuracy of the Charpy-V shift. Such a margin is not prescribedin the European or the Russian approaches. An interesting detail may be that theRegulatory Guide prescribes a larger margin for the welds than for the base materialand yet, the Hiser study (Fig. 18) showed base material to be more likely to beunderpredicted.

Despite the fact that the Russian and the French approaches may have a trend toyield less unconservative estimates of the fracture toughness shift than the ASMEapproach, their uncertainty in terms of the standard deviation is still anticipated to beapproximately σ∆T ≈ 20°C. This uncertainty is not accounted for (no additionalsafety margins) in either the Russian or the European approaches.

Because the fracture toughness irradiation shift is, in all approaches, estimated withfar from optimum type of tests, it is practically impossible to quantitatively assesstheir accuracy and degree of safety. In order to ensure their conservatism, additionalmargins 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 majorityof materials. An other possibility would be to show that for a certain type ofmaterial, the Charpy-V shift is conservative. This will, however, also penalize thematerial. Mechanistically the only sound solution is to measure the fracturetoughness directly for the irradiated material.

Recent advances allow broken Charpy-V specimen halves to be reconstituted to newCharpy-V size three point bend specimens, which can be used for a directmeasurement of the fracture toughness [15]. Fig. 19 shows an example for the HSSI5 series weld 73W. The large specimens were tested at ORNL as a part of the HSSI5 program [18], whereas the small Charpy-V size specimens (B = 1O mm) havebeen reconstituted from a broken half of a larger specimen and tested at VTTManufacturing Technology. The analysis of the results was identical to the onepresented in [15]. For the analysis, all fracture toughness results were corrected tocorrespond to 25 mm specimen thickness applying the statistical thicknesscorrection. Fig. 19 show that the fracture toughness can be determined reliably withCharpy-V size specimens. Thus, the major reason for using the Charpy-V test

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(applicability of small specimen) seems unwarranted. The same specimens canlikely be used to determine the fracture toughness directly, thus getting rid of anynecessity to prescribe additional safety margins. It appears that the constraintspeculations usually used to questionize small specimen fracture toughnessestimates are strongly exaggerated.

4.3 Chemistry factor

In all approaches the chemistry factors have been developed by an empiricalcorrelation of the material composition to the Charpy-V based transition temperatureshift. Thus the scientific basis for all the different approaches is equally poor.

Presently a large effort is being expended into trying to model irradiationembrittlement in order to be able to develop a micromechanism based chemistryfactor expression. This work has already given an improved understanding of theembrittlement mechanisms. However, as long as the simple Charpy-V transitiontemperature is used as reference, the chemistry factors can never be quantitativelyreliable 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

Fig. 19 Fracture toughness transition temperature T0 for 73W weld. T0correspond to 100 MPa√m mean fracture toughness for 25 mm crackfront length. B = 10 mm => reconstituted Charpy-V specimens.

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approaches for evaluating irradiation effects in reactor pressure vessels, thefollowing conclusions can be drawn:

- The Russian and European approaches have similar scientific bases, butmechanistically neither approach is optimal for evaluating the irradiation effects.In both cases the approaches are based on the Charpy-V test that is not directlydescriptive of the irradiated material's true fracture toughness. At the same timehighly conservative reference fracture toughness curves are usually applied.

- The comparison emphasizes the need for an improved approach for evaluatingirradiation effects in reactor pressure vessels. The new approach should be basedupon a direct determination of the fracture toughness, or a validated correlationbased on a direct determination of the fracture toughness, combined with amechanistic treatment of irradiation and material variables. This is the subjectmatter of Task Group C of AMES Project 1.

REFERENCES

1 Strength Analysis Standards for Equipment and Piping at Nuclear Power Plants,PNAE G-7-002-86.

2 Section III, Nuclear Power Plant Component, Division I, ASME Boiler and

Pressure Vessel Code, American Society of Mechanical Engineers, New York,1986.

3 Section XI, ASME Boiler and Pressure Vessel Code, Rules for InserviceInspection of Nuclear Power Plant Components, American Society ofMechanical Engineers, New York, 1986.

4 White Paper on Reactor Vessel Integrity Requirements for Level A and BConditions, EPRI TR-100251, 1993.

5 In-service Inspection Rules for Mechanical Components of PWR NuclearIslands, RSEM, EDF/afcen, 1990.

6 Design and Construction Rules for Mechanical Components of PWR NuclearIslands, RCC-M Code, 1988.

7 U. S. Nuclear Regulatory Commission, Regulatory Guide 1.99, RadiationEmbrittlement of Reactor Vessel Materials, Revision 2, Washington 1988.

8 U. S. Nuclear Regulatory Commission, Regulatory Guide 1.99, Effects ofResidual Elements on Predicted Radiation Damage to Reactor Vessel Materials,Revision 1, Washington 1977.

9 Sicherheitstechnische Regel des KTA, Überwachung der Strahlenversprödung

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von Werkstoffen des Reactordruckbehälters von Leichtwasserreaktoren, KTA3203, Fassung 3/84, 1984.

10 Kaufman, J. G., "Experience in Plane-Strain Fracture Toughness Testing PerASTM Method E 399," Developments in Fracture Mechanics Test MethodsStandardization, ASTM STP 632, W.F. Brown, Jr., and J.G. Kaufman, Eds.,American Society for Testing and Materials, Philadelphia, 1977, pp. 3-24.

11 Flaw Evaluation Procedures: Background and Application of ASME Section XIAppendix A, EPRI NP-719-SR, T.U. Marston, Ed., Electric Power ResearchInstitute, Palo Alto, 1978, pp. C8 - C12.

12 Wallin, K., The Size Effect in KIC Results, Engineering Fracture Mechanics, Vol.22, No. 1, 1985, pp. 149-163.

13 Wallin, K., Irradiation Damage Effects on the Fracture Toughness TransitionCurve Shape for Reactor Pressure Vessel Steels, International Journal ofPressure Vessels and Piping, Vol. 55, No. 1, 1993, pp. 61-79.

14 Wallin, K., Törrönen, K., Ahlstrand, R., Timofeev, B., Rybin, V., Nikolaev, V.and Morozov, A., Theory Based Statistical Interpretation of Brittle FractureToughness of Reactor Pressure Vessel Steel 15X2MΦA and its Welds, NuclearEngineering and Design, Vol. 135, 1992, pp. 239-246.

15 Wallin, K., Validity of Small Specimen Fracture Toughness EstimatesNeglecting Constraint Corrections, Constraint Effects in Fracture: Theory andApplications, ASTM STP 1244, M. Kirk and A. Bakker Eds., American Societyfor Testing and Materials, Philadelphia, 1994, In press.

16 Wallin, K., A Simple Theoretical Charpy-V - KIC Correlation for IrradiationEmbrittlement, Innovative Approaches to Irradiation Damage and FractureAnalysis, PVP-Vol. 170, D.L. Marriott et. al. Eds., The American Society ofMechanical Engineers, 1989, pp. 93-100.

17 Iskander, S. K., Corwin, W. R. and Nanstad, R. K., Effetcts of Irradiation onCrack-Arrest Toughness of Two High-Copper Welds, Effetcts of Radiation onMaterials: 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. 251-269.

18 Nanstad, R.K., Haggag, F.M., McCabe, D.E., Iskander, S.K., Bowman, K.O. andMenke, B.H., Irradiation Effects on Fracture Toughness of Two High-CopperSubmerged-Arc Welds, HSSI Series 5: Main Report and Appendices A, B, Cand D, NUREG/CR-5913, ORNL/TM-12165/V1, Vol. 1, Oak Ridge NationalLaboratory, Oak Ridge, 1992, 133 p. + 28 p. app.

19 Hiser, A.L., Correlation of Cv and KIC/KIC Transition Temperature Increases Dueto Irradiation, NUREG/CR-4395, MEA-2086, Materials Engineering Associates,Inc., 1985, 46 p. + app.