17
Validation and application of the Master Curve and reconstitution techniques to a Spanish nuclear vessel D. Ferreño a, * , M. Scibetta b , I. Gorrochategui c , R. Lacalle a , E. van Walle b , F. Gutiérrez-Solana a a University of Cantabria, ETS Ingenieros de Caminos, Av. Los Castros s/n, 39005 Santander, Spain b SCK-CEN, Boeretang 200, 2400 Mol, Belgium c CENTRO TECNOLÓGICO DE COMPONENTES (CTC), E.T.S. de Ingenieros Industriales y Telecomunicaciones, CDTUC, Universidad de Cantabria, Av. Los Castros s/n, 39005 Santander, Spain article info Article history: Received 25 March 2009 Received in revised form 13 August 2009 Accepted 21 August 2009 Available online 28 August 2009 Keywords: Reconstitution Master Curve Spanish nuclear vessel abstract In this work, a methodology based on the Master Curve approach and the reconstitution of specimens is validated and applied to the vessel base metal of the currently in service Spanish boiling water reactor of the Santa María de Garoña nuclear power plant. The exten- sive experimental program performed consisted in the characterisation of the ductile to brittle transition region with standard and reconstituted specimens using subsized com- pact tension and Pre-Cracked Charpy V-notch specimens, under non-irradiated and irradi- ated conditions. Experimental results validated the reconstitution technique down to inserts of 10 mm which allows specimen reorientation and therefore, the comparison of LT and TL material orientations. The 110 specimens tested in this program allowed the Master Curve approach to be validated for the base steel of Santa María de Garoña nuclear power plant. By comparing the results for the compact and Pre-Cracked Charpy V-notch specimens, the existence of a systematic bias between these two geometries has been ana- lysed. Although the neutron irradiation effect on the condition studied is predicted to be very limited (due to the small fluence), the results did allow the irradiation-induced shift of the ductile to brittle transition temperature to be detected. Comparison between the directly measured fracture toughness and the conventional semi-empirical approach pro- posed by the ASME Code reveals the overconservatism of the latter approach. Ó 2009 Elsevier Ltd. All rights reserved. 1. Introduction Nuclear reactor pressure vessel steels are degraded due to various causes during plant operation. Several embrittlement phenomena contribute to the degradation, neutron irradiation being the most relevant. This embrittlement leads to an in- crease in strength, a decrease in toughness and, as a consequence, a shift in the ductile to brittle transition temperature. Therefore, it is necessary to know in advance the evolution of material properties with irradiation to avoid the in service fail- ure of the vessel. For this reason, all nuclear utilities have a surveillance program in place that consists of placing, attached to the inside vessel wall in the beltline region (that is, the general area of the reactor vessel near the core midplane where radi- ation dose rates are relatively high), capsules holding specimens fabricated with the same steel as that of the vessel. Thus, the specimen irradiation history duplicates the neutron spectrum, temperature history, and maximum neutron fluence experi- enced by the reactor vessel inner surface. When the surveillance capsule lead factor (i.e., the ratio of the neutron fluence rate, E > 1 MeV, at the specimens in a surveillance capsule to the neutron fluence rate, E > 1 MeV, at the reactor pressure vessel 0013-7944/$ - see front matter Ó 2009 Elsevier Ltd. All rights reserved. doi:10.1016/j.engfracmech.2009.08.010 * Corresponding author. E-mail address: [email protected] (D. Ferreño). Engineering Fracture Mechanics 76 (2009) 2495–2511 Contents lists available at ScienceDirect Engineering Fracture Mechanics journal homepage: www.elsevier.com/locate/engfracmech

Validation and application of the Master Curve and reconstitution techniques to a Spanish nuclear vessel

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Page 1: Validation and application of the Master Curve and reconstitution techniques to a Spanish nuclear vessel

Engineering Fracture Mechanics 76 (2009) 2495–2511

Contents lists available at ScienceDirect

Engineering Fracture Mechanics

journal homepage: www.elsevier .com/locate /engfracmech

Validation and application of the Master Curve and reconstitutiontechniques to a Spanish nuclear vessel

D. Ferreño a,*, M. Scibetta b, I. Gorrochategui c, R. Lacalle a, E. van Walle b, F. Gutiérrez-Solana a

a University of Cantabria, ETS Ingenieros de Caminos, Av. Los Castros s/n, 39005 Santander, Spainb SCK-CEN, Boeretang 200, 2400 Mol, Belgiumc CENTRO TECNOLÓGICO DE COMPONENTES (CTC), E.T.S. de Ingenieros Industriales y Telecomunicaciones, CDTUC, Universidad de Cantabria,Av. Los Castros s/n, 39005 Santander, Spain

a r t i c l e i n f o

Article history:Received 25 March 2009Received in revised form 13 August 2009Accepted 21 August 2009Available online 28 August 2009

Keywords:ReconstitutionMaster CurveSpanish nuclear vessel

0013-7944/$ - see front matter � 2009 Elsevier Ltddoi:10.1016/j.engfracmech.2009.08.010

* Corresponding author.E-mail address: [email protected] (D. Ferreño)

a b s t r a c t

In this work, a methodology based on the Master Curve approach and the reconstitution ofspecimens is validated and applied to the vessel base metal of the currently in serviceSpanish boiling water reactor of the Santa María de Garoña nuclear power plant. The exten-sive experimental program performed consisted in the characterisation of the ductile tobrittle transition region with standard and reconstituted specimens using subsized com-pact tension and Pre-Cracked Charpy V-notch specimens, under non-irradiated and irradi-ated conditions. Experimental results validated the reconstitution technique down toinserts of 10 mm which allows specimen reorientation and therefore, the comparison ofLT and TL material orientations. The 110 specimens tested in this program allowed theMaster Curve approach to be validated for the base steel of Santa María de Garoña nuclearpower plant. By comparing the results for the compact and Pre-Cracked Charpy V-notchspecimens, the existence of a systematic bias between these two geometries has been ana-lysed. Although the neutron irradiation effect on the condition studied is predicted to bevery limited (due to the small fluence), the results did allow the irradiation-induced shiftof the ductile to brittle transition temperature to be detected. Comparison between thedirectly measured fracture toughness and the conventional semi-empirical approach pro-posed by the ASME Code reveals the overconservatism of the latter approach.

� 2009 Elsevier Ltd. All rights reserved.

1. Introduction

Nuclear reactor pressure vessel steels are degraded due to various causes during plant operation. Several embrittlementphenomena contribute to the degradation, neutron irradiation being the most relevant. This embrittlement leads to an in-crease in strength, a decrease in toughness and, as a consequence, a shift in the ductile to brittle transition temperature.Therefore, it is necessary to know in advance the evolution of material properties with irradiation to avoid the in service fail-ure of the vessel. For this reason, all nuclear utilities have a surveillance program in place that consists of placing, attached tothe inside vessel wall in the beltline region (that is, the general area of the reactor vessel near the core midplane where radi-ation dose rates are relatively high), capsules holding specimens fabricated with the same steel as that of the vessel. Thus, thespecimen irradiation history duplicates the neutron spectrum, temperature history, and maximum neutron fluence experi-enced by the reactor vessel inner surface. When the surveillance capsule lead factor (i.e., the ratio of the neutron fluence rate,E > 1 MeV, at the specimens in a surveillance capsule to the neutron fluence rate, E > 1 MeV, at the reactor pressure vessel

. All rights reserved.

.

Page 2: Validation and application of the Master Curve and reconstitution techniques to a Spanish nuclear vessel

Nomenclature

a crack depthB specimen thickness or crack front lengthB0 reference thickness for Master Curve analysis (B0 = 25.4 mm)b ligament sizeE Young’s modulusJ J integralJc critical value of J integralK stress intensity factorK0 normalizing fracture toughness corresponding to 63.2% probabilityKIC plane strain (quasi-static) fracture toughness; ASME curve describing a lower envelope of quasi-static fracture

toughness dataKIR ASME curve describing a lower envelope of dynamic and crack arrest fracture toughness dataKJ stress intensity factor determined from J integralKJc elastic–plastic stress intensity factor (determined from J integral)KJc (0.01) 1% lower bound Master Curve fracture toughnessKJc (0.05) 5% lower bound Master Curve fracture toughnessKJc (0.95) 95% upper bound Master Curve fracture toughnessKJc (0.99) 99% upper bound Master Curve fracture toughnessKJc,1T fracture toughness related to reference thickness B0

KJc (lim) maximum KJc capacity of a specimenKJc (med) median (50% cumulative failure probability) Master Curve fracture toughnessKJc,Pf Master Curve fracture toughness corresponding to a cumulative failure probability PfKmin Master Curve threshold fracture toughnessM margin to be added (Regulatory Guide 1.99) to obtain a conservative estimate of RTNDT

m Master Curve Weibull exponentPf cumulative failure probabilityr number of valid (non censored) dataRTNDT ASME reference temperature – nil ductility temperatureRTT0 Code Case N629/N631 T0-based reference temperaturesY yield stresssu ultimate stressT temperatureT0 Master Curve reference temperatureT28J temperature corresponding to the Charpy transition curve indexed at 28 JT41J temperature corresponding to the Charpy transition curve indexed at 41 JW specimen nominal widthb value used to obtain the standard deviation in T0

di =1.0 if the datum is valid or zero if the datum is a dummy substitute value, for T0 MML estimationm Poisson’s ratiorT0 standard deviation on the estimate of T0 associated with the use of only a few specimens to establish T0

DRTNDT shift in the reference temperature RTNDT due to irradiation

Acronyms3 PB (single edge notched) 3 point bend specimensASME American Society of Mechanical EngineersASTM American Society for Testing and MaterialsASW arc stud-weldingBWR boiling water reactorCT compact tension specimenCRP copper rich precipitationCVN Charpy V notchDBT ductile to brittle transitionDBTT ductile to brittle transition temperatureEBW electron-beam weldingEPFM elastic–plastic fracture mechanicsHAZ heat affected zoneHV hardness VickersLEFM linear-elastic fracture mechanicsMC Master Curve

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MML maximum likelihoodNPP nuclear power plantPCCv Pre-Cracked Charpy V notchRPV reactor pressure vesselSEM scanning electron microscopySMD stable matrix damageUSE upper shelf energyWZ weld zone

D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511 2497

inside the surface peak fluence location) is less than three, it is assumed that the calculational uncertainties in extrapolatingthe surveillance measurements from the specimens to the reactor pressure wall are minimised. The objectives of a reactorvessel surveillance program are twofold: first, to monitor changes in the fracture toughness properties and second, to makeuse of the data obtained to determine the conditions under which the vessel can be operated throughout its service life.

Before the development of elastic–plastic fracture mechanics in the 1970s only the tools of linear-elastic fracturemechanics were available to characterise the fracture toughness of the steel and to ensure the structural integrity of the ves-sel. Nevertheless, two drawbacks arose when trying to implement LEFM in the surveillance programs to characterise thematerial toughness in the ductile to brittle transition region. First, very large specimens are required to obtain valid lin-ear-elastic fracture toughness values, because of the stringent size requirements imposed by the KIc test Standard [1]. Sec-ond, fracture toughness in the DBT region shows large scatter and an extremely large number of specimens would benecessary to properly describe the behaviour of the material. Due to the intrinsic nuclear reactor space limitations, LEFMis therefore not applicable in surveillance programs. These facts forced nuclear surveillance programs to adopt indirectmeans, i.e. correlations between Charpy V notch impact toughness and fracture toughness, to assess the fracture toughnessof RPV steels under irradiated condition. The use of this indirect process to quantify the effects of irradiation on toughness,not theoretically justified, was, therefore, motivated by the extremely large specimens required to obtain valid LEFM tough-ness values.

Advances in fracture mechanics technology have made it possible to improve this approach in three different aspects.First, the development of EPFM allows determining fracture toughness values using much smaller specimens and utilisingJ integral techniques, that is, measuring values of KJc instead of KIc. Second, from halves of specimens tested in the surveil-lance program, usually CVN, it is feasible to construct new small specimens to perform additional fracture toughness tests;this possibility allows the evolution of steel vessel toughness with irradiation to be monitored from the beginning of plantoperation. Finally, the analytical techniques for structural integrity assessment can now be expressed in terms of EPFM. Thefirst two issues will be relevant for the contents of this paper; for this reason a brief description of each one of them is in-cluded in Sections 2 and 3.

For this research, the base steel of the vessel of the Spanish NPP of Santa María de Garoña has been characterised in theDBT region. The experimental process was performed with Pre-Cracked Charpy V-notch standard specimens and two con-figurations of reconstituted samples: compact tension (0.4T-CT, 10 mm thickness) and PCCv specimens with inserts of10 mm. This insert size is considered to be the lowest limit for reliable PCCv reconstitution. Moreover, it allows specimenreorientation and therefore, the comparison of LT and TL (see the ASTM [1] nomenclature) material orientations. The exper-imental scope of this study included 110 fracture toughness tests performed on non-irradiated and irradiated specimens. Themain goals of the study are here summarised:

� To provide a Master Curve description, based on the reference temperature T0, of the fracture toughness in the DBT region,under both unirradiated and irradiated conditions and to compare it with the classical methodology which is based on thereference temperatures RTNDT or RTT0 as indexing parameters (see Section 2). Thus, the safety margin implicit in the con-ventional method can be properly analysed.

� To quantify the shift in T0 experienced by the steel with neutron irradiation (material embrittlement) and to compare thisresult with the available data from the surveillance program, based on CVN impact tests.

� To analyse the behaviour of the LT and TL oriented specimens in the DBT region with the MC approach and to comparewith the CVN data from the surveillance program. As stated above, the LT orientation was characterised through10 mm implant reconstituted PCCv specimens.

� To validate the techniques for reconstitution of the different fracture toughness configurations here analysed (0.4T-CT andPCCv), by comparison between the results obtained through standard and reconstituted specimens.

� To compare the fracture toughness results obtained from PCCv specimens with those coming from CT specimens in orderto detect any bias between these two experimental configurations.

2. Description of fracture toughness in the DBT region

The regulations requiring the imposition of pressure–temperature limits on the reactor coolant pressure boundary forSpanish nuclear power plants designed in the USA are given by the federal regulation 10CFR50 [2]. This law establishes that

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fracture toughness requirements for ferritic materials must fulfil the acceptance and performance criteria of Appendix G ofSection 3 of the ASME Boiler and Pressure Vessel Code [3]. The toughness behaviour of the vessel material in the DBT regionbefore irradiation is described by the reference temperature RTNDT(U) obtained from Charpy impact and Pellini drop weighttests. This parameter is used to index two generic curves, developed in 1973, provided by the ASME Code [3,4] relatingtoughness vs. temperature, (see expressions (1, 2) where the temperatures must be expressed in �C and the toughness is ob-tained in MPa m1/2): the KIc curve describes the lower envelope to a large set of KIc data while the KIR is a lower envelope to acombined set of KIc, KId (dynamical test) and KIa (crack arrest test) data, being therefore more conservative than the former.Three important features can be appreciated: first, it is assumed that the ASME curves are representative of any vessel steel;second, in both curves, LEFM is considered; finally, the large scatter is removed by taking into account lower envelopes. Con-sequently, the method provides high conservatism in most cases.

K Ic ¼ 36:45þ 22:766 � e0:036�½T�RTNDT � ð1ÞK IR ¼ 29:40þ 13:776 � e0:0261�½T�RTNDT � ð2Þ

The decrease of material toughness due to neutron irradiation in the DBT region is currently estimated through semi-empir-ical methods based on the shift experienced by the Charpy impact curves obtained from the surveillance capsule specimenswhich are retrieved periodically, according to the plant withdrawal schedule. As stated in 10CFR50 [2], the effect of neutronfluence on the behaviour of the material is predicted by Regulatory Guide 1.99 [5] which provides Eq. (3) for the evolution ofRTNDT:

RTNDT ¼ RTNDT Uð Þ þ DRTNDT þM ð3Þ

where DRTNDT represents the shift in the reference temperature due to irradiation which is assumed to be equal to theshift of the Charpy transition curve indexed at 41 J; thus, DRTNDT = DT41J. The third term, M, is the margin that is to beadded to obtain a conservative estimation. The procedure in [5] allows DRTNDT to be obtained even when no crediblesurveillance data are available by means of an equation based on the chemistry of the steel and the neutron fluencereceived.

As an alternative to this semi-empirical indirect methodology, the MC approach, originally proposed by Wallin [6–9], pro-vides a reliable tool based on a direct characterisation of the fracture toughness in the DBT region. This approach is a con-sequence of the developments in EPFM together with an increased understanding of the micro-mechanisms of cleavagefracture. Valiente et al. [10] have briefly but comprehensively reviewed the previous contributions made to understandcleavage in a ferritic matrix that leads to the MC approach. The basic MC method for analysis of brittle fracture test resultsis defined in ASTM E1921-09-a [11]. The main features and advantages of the method are hereafter summarised. The math-ematical and empirical details of the procedure are available in [12,13]:

� MC assumes that cleavage fracture in non austenitic steels is triggered by the presence of particles close to the cracktip. Therefore, fracture is mainly an initiation dependent process. As a consequence, fracture is governed by weakestlink statistics which follows a three parameter Weibull distribution. For small-scale yielding conditions, thereforeusing EPFM, the cumulative failure probability, Pf, is given by (4):

Pf ¼ 1� e� B

B0�

KJc�KminK0�Kmin

� �m

ð4Þ

where KJc is the fracture toughness for the selected failure probability, Pf, K0 is a characteristic fracture toughness cor-responding to 63.2% cumulative failure probability, B is the specimen thickness and B0 a reference specimen thickness,B0 = 25.4 mm. The experimental data allows the Weibull exponent, m = 4, to be fixed and the minimum value of frac-ture toughness for the probability density function, Kmin = 20 MPa m1/2. Therefore, only K0 must be estimated from theempirical available data.

� The dependence between K0 (in MPa m1/2) and temperature (�C) for cleavage fracture toughness is assumed to be (5):

K0 ¼ 31þ 77 � e0:019�ðT�T0Þ ð5Þ

where T0 is the so called MC reference temperature; it corresponds to the temperature where the median fracturetoughness for a 25 mm thickness specimen (1T, according to ASTM terminology) has the value 100 MPa m1/2.

� One of the main advantages of the method is that it allows data from different size specimens to be compared. As

thickness increases, the toughness is reduced, due to the higher probability of finding a critical particle for the appliedload. ASTM Standard [11] provides expressions to relate the fracture toughness for specimens of different thickness.Eq. (4) can be re-written considering the same failure cumulative probability, Pf, for two specimens of different thick-ness, namely B1 and B2, thus leading to expression (6):

K Jc;2 ¼ Kmin þ ðK Jc;1 � KminÞ �B2

B1

� �14

ð6Þ

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D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511 2499

� The distribution fitting procedure involves finding the optimum value of T0 for a particular set of data. For this task, all

data are thickness adjusted to the reference specimen thickness B0 = 25.4 mm using Eq. (6). The procedure can beapplied either to a single test temperature or to a transition curve data, Ti being the generic temperature of the dif-ferent tests. In the latter approach (the former is just a particular case) T0 is estimated from the size adjusted KJC data(KJC,1T) using a multi-temperature maximum likelihood expression (see Eq. (7)). To estimate the reference tempera-ture, T0, a previous censoring of the non size adjusted data must be applied. Fracture toughness data that are greaterthan the validity limit given by Eq. (8), as defined in [11], are reduced to the validity limit, KJc(lim) and treated as cen-sored values in the subsequent estimation stage (di = 0 in expression (7)). This condition is imposed to guarantee highconstraint conditions in the crack front during the fracture process.

XN

i¼1

di �e0:019�ðTi�T0Þ

11þ 77 � e0:019�ðTi�ToÞ�XN

i¼1

ðKJc;i � 20Þ4 � e0:019�ðTi�T0Þ

½11þ 77 � e0:019�ðTi�T0Þ�5¼ 0 ð7Þ

K JcðlimÞ ¼ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi

E � sY � B0

30 � ð1� m2Þ

sð8Þ

In Eq. (8) sy is the yield strength at test temperature, E is the modulus of elasticity, B0 the initial ligament and m is thePoisson’s ratio (m = 0.3 in this case). It must be stressed that the factor 30 in Eq. (8) is currently under discussion[12,13] and that, for instance, the ASTM E1820-01 Standard [14] imposes a more demanding limit with a factor 50or 100 depending on the nature of the steel.

� The standard deviation in the estimate of T0, expressed in �C, is given by:

rT0 ¼bffiffiffirp ð9Þ

where r represents the total number of valid specimens (not censored results) used to establish T0. The values of thefactor b are provided in [11].

� The statistical analysis can be reliably performed even with a small number of fracture toughness tests (usually

between 6 and 10 specimens). Moreover, as an EPFM approach is used, the specimen size requirements, Eq. (8),are much less demanding than that of the LEFM [1]. These remarks are of great relevance in nuclear reactor surveil-lance programs where the amount of material available is usually very limited and consists of small size samples (CVNspecimens).

� By rearranging Eqs. (4) and (5) it is possible to obtain expression (10) which provides an estimate of KJc for a givencumulative failure probability, Pf, once T0 has been determined. In this way, the confidence bounds of the distribution(usually taking Pf = 0.01 or 0.05 for the lower bound and 0.95 or 0.99 for the upper bound) can be obtained. As a par-ticular case the expression for the median fracture toughness (Pf = 0.5) (see Eq. (11)) is determined.

K Jc;Pf¼ Kmin þ ½� lnð1� PfÞ�0:25 � 11þ 77 � e0:019�ðT�T0 Þ

h ið10Þ

K JcðmedÞ ¼ 30þ 70 � e0:019ðT�T0Þ ð11Þ

� Finally, any test that does not fulfil the requirement for crack front straightness or that terminates in cleavage aftermore than a limit of slow-stable crack growth will also be regarded as invalid.

� To enable the use of the MC methodology without completely modifying the structure of the ASME code [3,4] a dif-ferent approach was adopted, as stated in code cases N-629 [15] and N-631 [16]. It consists of defining a new indextemperature, RTT0, for the KIc and KIR ASME curves (1, 2), as an alternative to RTNDT. The definition of RTT0 is given inEq. (12). This value of RTT0 is set, see [17], by imposing that the ASME KIc curve indexed with RTT0 in place of RTNDT willbound the majority of the actual material fracture toughness data. In this sense, RTT0 was set such that the corre-sponding ASME KIc curve falls below the MC 95% confidence bound for at least 95% of the data generated with 1Tspecimens.

RTT0 ¼ T0 þ 19:4 �C ð12Þ

3. Reconstitution of specimens

For old plants, practically no coupons of original surveillance material remain unused. However, there are usually largequantities of previously tested specimens. Therefore, it seems reasonable to use this material as efficiently as possible. Thetechnique of constructing new specimens from small quantities of material is commonly called ‘reconstitution’. The con-struction of compound specimens is achieved by attaching additional material (studs) around a material of interest (insert)to prepare a test specimen of standard dimensions. This is especially important in the context of nuclear surveillance pro-grams where the available material is limited and the evolution of fracture mechanical properties with neutron irradiation

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2500 D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511

must be monitored. The possibility of fabricating new specimens with material orientations different from the original spec-imen, previously tested, represents another important advantage of the reconstitution technique.

Methods for reconstitution of fracture toughness specimens have been developed in the past [18,19]. Currently, twoexperimental configurations are widely reconstituted from CVN specimens previously tested: PCCv, and CT specimens. Itmust be emphasised that the possibility to characterise the fracture toughness in the DBT region with these subsized con-figurations is a direct consequence of the MC methodology, which describes the fracture properties of the material by meansof EPFM, relaxing the size requirements.

The interface between stud and insert is obtained using welding techniques. Ideally, the test results on reconstitutedspecimens should be identical to those of full-size single-material samples; for this purpose, the welding technique mustguarantee that the reconstitution process has not altered the material of interest. The first condition is that the fractionof the insert that is affected by the reconstitution technique (weld and heat affected zones, WZ and HAZ, respectively) mustbe small enough to permit the correct performance of the subsequent fracture test. In addition, for irradiated material, thetemperature within the test volume must remain low enough to induce no annealing effects.

4. Experimental

4.1. Materials

The steel examined in this research was supplied by the Spanish NPP of Santa María de Garoña. It consists of a set of 26non tested LT oriented CVN specimens. 15 of them were under unirradiated condition whereas the rest were irradiated to afluence up to 3.55 � 1017 n cm�2 (E > 1 MeV). The irradiated material comes from an experimental capsule exclusively dedi-cated to this research project. Also, the experimental results obtained from the capsules 1 and 2 of the surveillance programof the plant (with fluences up to 5.70 � 1017 and 1.26 � 1018 n cm�2, E > 1 MeV, respectively) were available for analysis andcomparison. The chemical composition, given in Table 1, corresponds to a SA-336 steel, according to the ASME [20] speci-fication. The steel presented a microstructure of ferrite with presence of bainite. This result is coherent with the Vickersmicrohardness tests performed, with a result of 210 ± 15 HV.

4.2. Description of the reconstitution of specimens

The experimental configurations adopted in this study consist of standard and reconstituted PCCv and 0.4T CT reconsti-tuted specimens. PCCv specimens were reconstituted by SCK-CEN with arc stud-welding whereas CTs were reconstructedwith electron-beam welding by FRAMATOME ANP. To optimise the use of available representative material, the reconstitu-tion process has been as exhaustive as possible. From each half broken Charpy specimen, two PCCv implants, 10 mm long, oralternatively, one CT implant were obtained. One of the major advantages of using 10 mm implant PCCv specimens is that itallows material reorientation to be performed; the process to obtain LT and TL oriented implants from a (previously tested)CVN or PCCv LT specimen is shown in Fig. 1.

The welding conditions in both cases were selected in order to minimise the volume of affected material and followingthe recommendations given in [19]. Furthermore, during and after welding, the fraction of the insert that is affected by thereconstitution process was determined by several techniques. A brief summary is here included for both techniques [21]:

� To ensure that the welding process does not introduce material property changes in the insert, it is important to guaranteethat the temperature in the volume of the insert remains less than the specimen irradiation temperature. Therefore, mea-surements of temperature in the insert were performed by brazing into the dummy insert calibrated insulated thermo-couples. For ASW reconstituted PCCv specimens, the maximum temperature at a distance of 5 mm from the weldingdid not exceed 240 �C for all measurements, well below the reactor operating temperature (280 �C). For EBW reconsti-tuted CTs, temperatures occasionally exceeded 280 �C during a time less than 2.4 s, which is considered to be insufficientto promote any annealing in the irradiated material.

� On selected ASW PCCv specimens, hardness Vickers tests was performed. Distance between first and second weld was inall cases in the order of 5 mm, see Fig. 2, which is sufficient for fracture mechanics tests.

� Non-irradiated unnotched PCCv reconstituted specimens were tested under quasi-static loading conditions (displacementcontrol) up to a maximum load of 30 kN at upper shelf temperature (+200 �C) and at lower shelf temperatures (�70 �C).The results from the three point bend tests show that all arc stud welds are resistant to fracture well above loads that willbe applied during 3 PB tests of PCCv specimens.

Table 1Chemical composition of the steel of this research (wt.%).

C Mn P S Si Ni Cr Mo Cu

0.181 0.580 0.012 0.013 0.350 0.720 0.320 0.610 0.100

Page 7: Validation and application of the Master Curve and reconstitution techniques to a Spanish nuclear vessel

Fig. 1. Schematic description of material reorientation to obtain 10 mm implant PCCv, LT or TL, reconstituted specimens.

0 2 4 6 8 10 12 14 16 18

200

250

300

350

400

450

500

3.5 mm4 mm 5 mm

weld 2 weld 1

HV 05

Distance (mm)

Fig. 2. Microhardness tests performed on ASW PCCv reconstituted specimens.

D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511 2501

4.3. Experimental scope

Table 2 summarises the experimental test matrix. In all cases, KJc tests have been performed to subsequently obtain T0.After testing the (15 + 11) standard PCCv LT oriented specimens, 60 PCCv were reconstituted with ASW, half of them LT ori-ented and the rest TL oriented, together with 24 0.4T-CT LT specimens with EBW (as specified in Tables 2 and 4 of the CTirradiated specimens were available in the beginning of the study and were not obtained from the 26 PCCv available for thisresearch).

The reconstitution strategy was as follows: after testing 15 PCCv non-irradiated standard specimens (see Table 2), 15halves were used to reconstitute 15 LT plus 15 TL specimens with 10 mm insert, 13 halves were reused to reconstitute13 CT and the additional two halves were used for fabrication of eight miniature tensile specimens. The same rule was

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Table 2Experimental test matrix.

Standard Reconstituted

PCCv PCCv 0.4T CT

Non-irradiated 15 LT 15 LT 13 LT15 TL

Irradiated (3.55 � 1017 n cm2 E > 1 MeV) 11 LT 15 LT 11 LT5 TL (7 + 4 avail.)

2502 D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511

followed in the case of irradiated material. It was considered that 10 mm insert length is the lowest limit for PCCv reconsti-tution by means of ASW. The overall number of fracture tests rises up to 110. Moreover, eight mini-tensile specimens (with adiameter of 2.4 mm) have been manufactured from one PCCv specimen of non-irradiated material and four from other PCCvspecimen of irradiated material.

This test matrix allows some interesting comparisons to be performed. In particular, the effect of irradiation, the influenceof the specimen configuration (CT vs. PCCv), the material orientation (LT vs. TL) and the reliability of reconstitution as a suit-able tool to obtain T0 can be analysed, as stated in the summary of goals of this paper.

4.4. Tensile tests

The experimental scope of the project included some tensile tests to be performed both under non-irradiated conditionand the research dose level, 3.55 � 1017 n cm�2 (E > 1 MeV). The non-irradiated steel was subjected to tensile testing at tem-peratures ranging from �150 to 300 �C. Eight tests were performed to cover this temperature interval. Four specimens ofirradiated material (3.55 � 1017 n cm�2) were tested in the range from �150 �C to 260 �C. The load–elongation curves wereregistered up to fracture by using extensometers designed to operate at low and high temperatures. Fig. 3 summarisesthe results of yield stress, sY, and ultimate strength, sU. The temperature dependence of these results in both material con-ditions was fitted using suitable expressions of the form sY, sU = A + B�exp(�C � T). The fittings can also be appreciated inFig. 3. The knowledge of the tensile properties is necessary to evaluate the maximum measurement capacity within theASTM E1921 Standard [11], expression (8). Moreover, Fig. 4 shows the results of uniform and total strain as a function oftemperature whereas the data of reduction of area are included in Fig. 5.

4.5. Fracture characterisation in the DBT region

4.5.1. Previous results from the surveillance capsules 1 and 2The main results from the surveillance program consist of CVN impact tests performed on LT specimens both under non-

irradiated condition and two levels of fluence corresponding to capsules 1 and 2, respectively. The experimental values of

0

200

400

600

800

1000

1200

1400

-200 -150 -100 -50 0 50 100 150 200 250 300Temperature (ºC)

Stre

ss (M

Pa)

sY fitting (non irradiated) sU fitting (non irradiated)

sY fitting (3.55 E+17 n·cm-2) sU fitting (3.55 E17 n·cm-2)

sY (non irradiated) sU (non irradiated)

sY (3.55 E17 n·cm-2) sU (3.55 E17 n·cm-2)

Fig. 3. Yield stress (sY) and ultimate strength (sU) as a function of temperature (including fittings) for non-irradiated and irradiated (3.55 � 1017 n cm�2)material.

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0

5

10

15

20

25

30

35

-200 -150 -100 -50 0 50 100 150 200 250 300 350

Temperature (ºC)

Stra

in (%

)

Non irradiated - uniform strainNon irradiated - total strainIrradiated (3.55 E+17 n/cm2) - uniform strainIrradiated (3.55 E+17 n/cm2) - total strain

Fig. 4. Uniform and total strain as a function of temperature for non-irradiated and irradiated (3.55 � 1017 n cm�2) material.

62

64

66

68

70

72

74

76

-200 -150 -100 -50 0 50 100 150 200 250 300 350

Temperature (ºC)

RA

(%)

Non irradiated - Reduction of area

Irradiated - Reduction of area

Fig. 5. Reduction of area as a function of temperature for non-irradiated and irradiated (3.55 � 1017 n cm�2) material.

D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511 2503

absorbed energy vs. temperature are shown in Fig. 6. The hyperbolic tangent fitting of the data are included in the figure and,from them, the values of T41J and the upper shelf energy for the three material conditions were obtained. Moreover, addi-tional CVN tests were performed on TL oriented non-irradiated specimens. A comparison between the LT and TL responsesof the non-irradiated material is shown in Fig. 7. Again, the fitting of the data were included together with T28J and T41J andUSE values. Finally, the reference temperature RTNDT = 16 �C was provided by the NPP.

4.5.2. Fracture toughness testsThe fracture KJc tests were carried out according to the ASTM 1921 Standard [11]. The fatigue precrack was produced at

room temperature using the tensile properties of the steel to choose the fatigue loads. The Standard requires the relationbetween the initial fatigue crack length and the width of the specimen, a0/W, to be within the interval (0.45–0.55). Thisrequirement has been fully satisfied for PCCv specimens. Nevertheless, with regards to CT reconstituted specimens, this isnot possible because, under those conditions, the fracture process zone lies within the WZ or HAZ, at the interface betweenthe insert and the implant. To ensure the representativeness of the test, it is absolutely essential that the crack tip is well

Page 10: Validation and application of the Master Curve and reconstitution techniques to a Spanish nuclear vessel

-20

20

60

100

140

180

220

260

-100 -50 0 50 100 150 200 250 300

T (ºC)

E (J

)

LT non irradiated

Fitting LT Non Irradiated

LT Capsule 1 (5.70 E+17 n/cm2)

Fitting LT Capsule 1 (5.70 E+17 n/cm2)

LT Capsule 2 (1.26 E+18 n/cm2)

Fitting LT Capsule 2 (1.26 E+18 n/cm2)

199 J190 J187 J

-25ºC-29ºC

-54ºC 41J

Fig. 6. CVN curves (energy vs. temperature) for non-irradiated, capsules 1 and 2, LT oriented material.

-20

20

60

100

140

180

220

260

-100 -50 0 50 100 150 200 250 300

T (ºC)

E (J

)

LT non irradiated

Fitting LT non irradiated

TL non irradiated

Fitting TL non irradiated

41J

28J

199 J

115 J

LT

TL

-25ºC -11ºC

-24ºC-29ºC

Δ (USE) = 84 ºC

Fig. 7. Comparison between CVN curves of LT and TL oriented non-irradiated material.

W + HAZ

Plastic zone at crack tip

Fig. 8. Schematic description of the possible contact between plastic zone at crack tip and WZ + HAZ.

2504 D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511

Page 11: Validation and application of the Master Curve and reconstitution techniques to a Spanish nuclear vessel

Fig. 9. Scheme of microhardness profiles performed on reconstituted CT specimens.

D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511 2505

within the insert (material of interest) even avoiding the contact between the plastic zone developed during the test and theWZ + HAZ (see Fig. 8). Therefore, CT reconstituted configuration requires initial crack lengths slightly longer than the limitprovided in the current Standard. To limit the crack length, the EBW technique for the reconstitution of CT specimens wasselected because, with appropriate welding parameters, it provides the narrowest possible WZ + HAZ. Previous works havedemonstrated the reliability of this solution even for a fracture process under elastic–plastic conditions [22,23]. Taking intoaccount the fact that deeper cracks exhibit more constraint than shallow cracks, no loss of constraint effect due to the vio-lation of this requirement is expected.

To determine the crack length necessary, the thickness of WZ + HAZ was evaluated prior to precracking any CT specimen.Microhardness measurements (200 g, 20 s) were performed to evaluate the width of this region. In Fig. 9, a scheme showingthe position for the microhardness profiles is presented. As shown in Fig. 10, the most adverse conditions imply a minimuminitial crack length longer than 11.8 mm (a0/W = 0.59). This length allows the crack to go beyond the region of hard embrit-tled material melted during the welding process and quenched during cooling, and the soft annealed material close to thatregion. Moreover, to conservatively guarantee the representativeness of the analysis, only those reconstituted CT specimensshowing cracks longer than 12.19 mm (a0/W � 0.61) were used to establish T0.

The fracture test temperatures were selected using the empirical correlations provided by the Standard [11], based on thevalues of T28J and T41J (see Figs. 6 and 7). It is worth noting that these correlations can be far from accurate; nevertheless, the

150

200

250

300

350

400

450

500

550

6 7 8 9 10 11 12 13 14

Distance to load line (mm)

Mic

ro H

ardn

ess

Vick

ers

(200

g. 2

0s. ) Line A

LIne B

Line C

Fig. 10. Example of microhardness profiles performed to estimate the width of the WZ + HAZ on a reconstituted CT specimen.

Page 12: Validation and application of the Master Curve and reconstitution techniques to a Spanish nuclear vessel

Fig. 11. Experimental set-up for fracture tests of CT specimens.

2506 D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511

multi-temperature procedure permits different specimens to be tested at different temperatures, getting closer to T0 as thenumber of tests increases. The test temperatures ranged from �130 to �60 �C.

The loading method consisted of applying a controlled rate of crack opening displacement as measured at the load line(1 mm/min). The J integral was calculated from the load-crack opening displacement curve, in accordance with the proce-dure of the ASTM Standard [11], at crack instability, which presumably corresponds to a cleavage event. The arrangementshown in Fig. 11 has been designed to obtain a direct measurement of load line displacement in CT specimens in spite ofthe possibility, present in the Standard [11], to obtain load line displacements (necessary to properly evaluate J integral)from face displacement measurements with a proportionality factor.

All specimens failed due to unstable crack extension, not preceded by stable tearing, at loading levels ranging from linear-elasticity to general yielding, as shown by the load–displacement curves recorded prior to fracture. Most of the failures oc-curred before the end of J dominance, namely, below the limit adopted by the Standard [11] to ensure high constraint con-ditions (expression (8)). The results that did not fulfil this requirement were censored for T0 estimation.

After performing the tests, evaluating the KJc values, applying the censoring process (8) and calculating the size adjustedKJc data (KJc,1T), according to Eq. (6), the reference temperatures T0 were estimated for every family (by applying Eq. (7))shown in Table 2. The final results are summarised in Table 3, including 2r as the uncertainty. The results obtained by jointlyanalysing all the LT unirradiated and irradiated specimens (which will be relevant for later analysis) are shown in Table 4.

Fig. 12 represents the KJc,1T experimental results vs. (T � T0). This normalised graph permits all the points to be includedtogether, with independence of the material orientation or state of irradiation. The curves for KJc(med), the lower bandsKJc(0.01) � KJc(0.05), and the upper bands KJc(0.95) � KJc(0.99) have also been plotted, showing the great coherence between thetheoretical predictions and the experimental results.

From a macromechanical point of view, the failure of all the fracture specimens occurred by unstable propagation of thefatigue crack. The fracture surfaces were examined by SEM to ascertain the existence of cleavage as the physical fracturemechanism and to localize the initiation sites. The SEM images were in agreement with the macromechanical cleavageevent. All the specimens showed the multifaceted surface with river patterns typical of cleavage fracture. The general aspectof the fracture surfaces is illustrated by the photograph of a reconstituted CT specimen shown as an example in Fig. 13. Thetwo SEM images of the fracture surface, also included in the figure, show (b), a detail of the cleavages and (c), the borderbetween the fatigue and propagation regions.

Table 3Reference temperature results, T0 (�C) ± 2r.

Standard Reconstituted

PCCv PCCv 0.4 CT

Unirradiated �105 ± 11 (15 LT) �95 ± 10 (13 LT) �89 ± 15 (7 LT)�104 ± 11 (15 TL)

Irradiated (3.55 � 1017 n cm�2) �91 ± 11 (11 LT) �93 ± 10 (14 LT) �84 ± 11 (11 LT)�93 ± 10 (15 TL)

Table 4Reference temperature results, T0 (�C) ± 2r, by jointly analysing LT results.

LT

Unirradiated �98 ± 6 (35 LT)Irradiated (3.55 � 1017 n cm2) �89 ± 6 (36)

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0

25

50

75

100

125

150

175

200

225

250

-60 -40 -20 0 20 40 60

T-T0 (ºC)

KJc

,1T (

MPa

·m1/

2 )

PCCv standard, LT, non irradiated PCCv reconstituted, LT, non irradiated

PCCv reconstituted, TL, non irradiated CT reconstituted, LT, non irradiated

PCCv standard, LT, irradiated (3.55 E+17 n/cm2) PCCv reconstituted, LT, irradiated (3.55 E+17 n/cm2)

PCCv reconstituted, TL, irradiated (3.55 E+17 n/cm2) CT reconstituted, LT, irradited (3.55 E+17 n/cm2)

KJc (0.01)

KJc (0.05)

KJc (0.95)KJc (0.99)

KJc (med)

Fig. 12. Representation of KJc,1T results on a uniform temperature scale, (T � T0). Several confidence bounds are included.

Fig. 13. Photograph of the fracture surface of a reconstituted CT specimen (a) and SEM images showing the cleavage area (b) and the fatigue precrack (c).

D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511 2507

For the purpose of comparing the different approaches to describe the fracture toughness in the DBT region, Fig. 14 showsthe experimental results concerning non-irradiated PCCv standard specimens (LT orientation). For this family of specimens,see Table 3, T0 = �105 �C. Several confidence MC bounds together with the KJc(lim) condition (8) were also represented. TheASME curves (1, 2) indexed at RTNDT = 16 �C and RTT0 = �105 + 19.4 = �85.6 �C have been included in the figure. As can beappreciated, both the KIc and KIR, indexed at RTNDT hardly follow the pattern of experimental points with temperatureand are, in general, extremely conservative. This is a consequence of the great difference (121 �C) between RTNDT and T0.

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0

50

100

150

200

-170 -150 - 130 -110 -90 -70 -50 -30

T (ºC)

KJc

(MPa

·m1/

2 )

KJc (med)

KJc (0.05)

KJc (0.95)

KJc (0.01)

KJc (0.99)Non irradiated material Standard PCCv - LT specimensT0 = -105 ºC

KJc(lim)

KIc (RTNDT)KIR (RTNDT)

KIc (RTT0)

KIR (RTT0)

Fig. 14. Comparison between the different available curves (ASME curves indexed at RTNDT or RTT0 and MC) to describe the fracture toughness in the DBTregion for the non-irradiated LT material.

2508 D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511

Moreover, the curves indexed at RTT0 seem to follow fairly closely the MC lower bounds; this is particularly evident in thecase of the KIc curve and the KJc(0.01) MC confidence bound.

4.6. Discussion of experimental results

The previous results can provide some answers to the open issues mentioned above concerning the fracture response ofthe material in the DBT region, as summarised in the following sections.

4.6.1. Material embrittlementThree independent sources are available to evaluate the potential damage of fracture properties due to neutron irradia-

tion: tensile, CVN and fracture toughness tests. Concerning the tensile results, see Fig. 3, it is evident that irradiation(3.55 � 1017 n cm�2) has scarcely affected the tensile properties of the material. The comparison between the fittings ofthe non-irradiated and irradiated material shows a slight increment of yield stress and ultimate strength, up to 60 and30 MPa for high temperatures, respectively. Concerning the effect of fluence on deformation characteristics, a slight reduc-tion of uniform or total strain can be appreciated in Fig. 4. The same pattern is evident in Fig. 5, where the reduction of area isrepresented.

Fig. 6 shows the CVN curves obtained from LT oriented specimens. The effect of embrittlement is very difficult to estimateas the reference temperature for non-irradiated material is T41J = �25 �C, for capsule 1 (5.70 � 1017 n cm�2), T41J = �54 �C andfor capsule 2 (1.26 � 1018 n cm�2) T41J = �29 �C. This negative shift in the DBT region can be regarded as an artefact motivatedby a non adequate selection of test temperatures: as can be appreciated, most of the points lie in the DBT region, where alarge scatter is present; this fact is particularly evident in the case of the specimens from capsule 1. Therefore, it seems dif-ficult to make accurate statements concerning material embrittlement from this source unless, if this exists, it is not verysevere.

This assertion is supported by the almost unchanged results of USE with irradiation. Furthermore, it is worth noting thatthe embrittlement predictions proposed by the Regulatory Guide [5], the ASTM Standard E900-02 [24] or the correlationsgiven in the EPRI report [25] are consistent with the experimental results here reported (see Table 5). For example, ASTME900-02 [24] implies an expected transition temperature shift of 15 �C at a fluence level corresponding to capsule 2

Table 5Theoretical predictions for material embrittlement.

Fluence (n cm�2) RG 1.99 [5] ASTM E900–02 [22] EPRI [23]

SMD CRP Total SMD CRP Total

3.55 � 1017 9 ± 5 3 5 8 ± 12 3 3 65.70 � 1017 12 ± 6 3 6 9 ± 12 4 5 91.26 � 1018 17 ± 9 5 10 15 ± 12 5 9 14

Page 15: Validation and application of the Master Curve and reconstitution techniques to a Spanish nuclear vessel

-125

-115

-105

-95

-85

-75

25,11

T 0 (º

C)

PCCv standard LT (non irradiated) PCCv reconstituted LT (non irradiated)

CT reconstituted LT (non irradiated) PCCv standard LT (3.55 E+17 n/cm2)

PCCv reconstituted LT (3.55 E+17 n/cm2) CT reconstituted LT (3.55 E+17 n/cm2)

UNIRRADIATED MATERIAL

IRRADIATED MATERIAL

(3.55 E+17 n·cm-2)

ΔT0 = 9ºC

Fig. 15. Effect of irradiation, (3.55 � 1017 n cm�2) on T0 for the different experimental configurations (PCCv and CT) of LT oriented specimens.

D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511 2509

(1.26 � 1018 n cm�2) which is the consequence of a combination of stable matrix damage, SMD, of 5 �C and copper rich pre-cipitation, CRP, of 10 �C. However, the uncertainty related to this estimation is in the order of 12 �C.

Finally, Tables 3 and 4 present the results for the MC reference temperature T0. No important shift in T0 is observed. Theaverage shift in LT orientation is 9 �C (obtained after jointly analysing all the data corresponding to LT material, that is, PCCvnormalised and reconstituted together with reconstituted CT specimens, see Table 4) whereas in TL material, this is 11 �C.Figs. 15 and 16 present a summary of reference temperature results for the different configurations including the standarddeviation.

In summary, for this level of fluence, a moderate embrittlement is appreciated in the steel from the tensile tests, CVN andT0 tests. This result is coherent to what the theoretical models predict.

4.6.2. Effect of material orientation: LT vs. TLTwo different material orientations, LT and TL, have been analysed. A large difference in behaviour is suggested by the

Charpy impact tests, see Fig. 7. It is well known that the fracture toughness of many materials is highly sensitive to the ori-entation of the crack plane with respect to the principal axis used during fabrication. In particular, the TL orientation is usu-ally considered to be weaker than the LT orientation. Nevertheless, this effect does not seem to be reflected in the T0 results(see Table 3). Only small differences, compared to the uncertainties, are appreciated between the reference temperature for

-120

-115

-110

-105

-100

-95

-90

-85

-80

-75

-70

T 0 (º

C)

PCCv reconstituted TL (non irradiated)

PCCv reconstituted TL (3.55 E+17 n/cm2)

UNIRRADIATED MATERIAL

IRRADIATED MATERIAL (3.55 E+17 n·cm-2)

ΔT0=11ºC

Fig. 16. Effect of irradiation (3.55 � 1017 n cm�2) on T0 for TL oriented reconstituted PCCv specimens.

Page 16: Validation and application of the Master Curve and reconstitution techniques to a Spanish nuclear vessel

-110

-105

-100

-95

-90

-85

-80

-110 -105 -100 -95 -90 -85 -80

T0 (CT)

T 0 (P

CC

v)

LT-Unirradiated / PCCv standard vs. CT reconstitutedLT-Unirradiated / PCCv reconstituted vs. CT reconstitutedLT-Irradiated (3.55 E+17 n/cm2) / PCCv standard vs. CT reconstitutedLT-Irradiated (3.55 E+17 n/cm2) / PCCv reconstituted vs. CT reconstituted

T0 (PCCv) = T0(CT) + 10ºC

T0 (PCCv) = T0(CT) - 10ºC

T0 (PCCv) = T0(CT)

16ºC

6ºC

9ºC7ºC

Fig. 17. Comparison between T0 results obtained from PCCv and CT specimens.

2510 D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511

LT and TL orientations. This contradiction is only apparent because, as can be seen in Fig. 7, the LT and TL CVN curves overlapin the lower part of the DBT region, where the fracture process is mainly an initiation controlled process and T0 is represen-tative. Therefore, the CVN and T0 results were coherent. In addition, the low Charpy upper shelf energy of the TL orientationis partly due to the reconstitution with 10 mm insert reorientation.

4.6.3. Effect of specimen geometryA systematic difference between the results of MC reference temperature obtained through PCCv and CT specimens is re-

ported in the literature [11,12,26,27] and acknowledge since version 2003 of the ASTM E1921 Standard. It is assumed that asystematic bias exists leading to higher T0 values, between 10 and 15 �C, obtained from CTs than from PCCv specimens. Inthis sense, the results coming from CTs are systematically more conservative than those coming from PCCv. This phenom-enon has been explained as a consequence of the different constraint conditions for these two geometries.

Table 3 confirms the existence of this effect for the specimens tested in this research. For example, for non-irradiatedmaterial, the difference between PCCv standard specimens and CT reconstituted specimens is 16 �C whereas for irradiatedmaterial, it is 7 �C. Even though the bias is in the order of the uncertainties reported in Table 3, the relevant fact is that itappears in a systematic fashion: in all cases T0 values obtained through CT specimens were higher than those coming fromPCCv. This difference can hardly be considered as an artefact derived from the constraint rising due to the use of slightlylonger initial crack lengths in CT specimens, mentioned above. This increase in constraint is presumably negligible in com-parison with the effect associated with the specimen geometry (CT vs. PCCv). In fact, the initial precrack lengths adoptedwere in the range a0/W = 0.61–0.68 whereas the interval imposed in the ASTM Standard [11] is 0.45–0.55. This small increasehardly influences the fracture resistance of the material. Empirical information supporting this argument is present in[22,28]. In Fig. 17, a comparison between the T0 results from PCCv and those from CTs is presented, emphasising the conser-vative response given by CT specimens.

5. Summary and conclusions

An exhaustive research has been carried out to characterise the base metal of the vessel of the Spanish NPP of Santa Maríade Garoña in the DBT region. The scope of the research included 110 fracture toughness (KJc) tests performed on standard(PCCv) and reconstituted (PCCv and 0.4T-CT) specimens under unirradiated and irradiated (3.55 � 1017 n cm�2) conditions.The reconstitution of 10 mm insert-PCCv specimens allowed material reorientation and therefore to characterise the behav-iour of the TL oriented material and to compare with the LT response (obtained through standard and reconstituted speci-mens). The analysis of the experimental results allows the following conclusions to be established:

� The two configurations of reconstituted specimens here analysed, 0.4T-CT and PCCv (10 mm implant), were validated. ForCT specimens EBW technique was used whereas the PCCv were reconstituted by means of ASW. The small differences in T0

detected between standard and reconstituted specimens are in agreement with the expected uncertainty in the determi-nation of T0.

� It was proved that MC accurately describes the response of the base material analysed in this research, LT and TL orien-tations, both under non-irradiated and irradiated conditions.

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D. Ferreño et al. / Engineering Fracture Mechanics 76 (2009) 2495–2511 2511

� PCCv specimens were reconstituted using implants of 10 mm, which is considered to be the technical limit to perform areliable fracture toughness test. With this implant size, it is possible to obtain four reconstituted specimens from each pre-viously tested CVN specimen, thus allowing the most exhaustive use of the available material.

� Specimen reorientation in PCCv reconstituted specimens allowed comparison between LT and TL orientations to be per-formed. No significant differences between LT and TL reference temperatures (T0) were detected; therefore, it can be con-cluded that the material orientation does not play a role in the fracture toughness at cleavage initiation for this material,which is consistent with the CVN curves (surveillance program) in the lower part of the DBT region.

� A systematic bias has been observed between T0 values from CT and PCCv. This bias is not unexpected and is a conse-quence of the different constraint conditions between these two configurations, leading to a more conservative character-isation by means of CT, as compared to PCCv.

� The ASME curves KIR and KIc indexed at RTNDT and RTT0 were compared with the MC description in the DBT region. Thiscomparison allowed highlighting the great conservatism implicit in the conventional approach to describe the fracture inthe DBT region. The differences between RTNDT and T0 in all the families here analysed were as high as 100 �C, whichimplies a large underestimation of safety margins.

Acknowledgments

This investigation was performed within a research project (CUPRIVA) sponsored by the Spanish Nuclear Regulatory Bodyand the electric company UNESA, represented by J.M. Figueras and L. Francia, respectively. The authors wish to express par-ticular gratitude to their colleagues Ph.D. A. Ballesteros and Eng. X. Jardí for their contribution to the dosimetry measure-ments and analysis, and to Eng. Javier Martín for the tasks of coordination and analysis.

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