Structural integrity assessment of a nuclear vessel with FITNET FFS and Master Curve approach

Engineering Failure Analysis 17 (2010) 259–269

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Structural integrity assessment of a nuclear vessel with FITNET FFS and Master Curve approach D. Ferreño a,*, R. Lacalle a, R. Cicero b, M. Scibetta c, I. Gorrochategui d, E. van Walle c, F. Gutiérrez-Solana a a

University of Cantabria, ETS Ingenieros de Caminos, Av. Los Castros s/n, 39005 Santander, Spain Inesco Ingenieros, ETS de Ingenieros Industriales y Telecomunicaciones, CDTUC, Universidad de Cantabria, Av. Los Castros s/n, 39005 Santander, España, Spain c SCK-CEN, Boeretang 200, 2400 Mol, Belgium d CENTRO TECNOLÓGICO DE COMPONENTES (CTC), ETS de Ingenieros Industriales y Telecomunicaciones, CDTUC, Universidad de Cantabria, Av. Los Castros s/n, 39005 Santander, España, Spain b

a r t i c l e

i n f o

Article history: Received 4 June 2009 Accepted 6 June 2009 Available online 12 June 2009 Keywords: Master Curve Structural integrity Nuclear vessel FITNET FFS

a b s t r a c t The structural integrity of a Spanish nuclear vessel was assessed by obtaining the operation pressure–temperature (P–T) limit curves relevant to the beltline region according to the ASME Code. This method requires the knowledge of the material fracture toughness in the ductile to brittle transition (DBT) region which is conventionally described by the semi-empirical reference temperature RTNDT. Recent development of the Master Curve (MC) methodology, with T0 as a reference temperature, is an attractive alternative that do not rely on empirical correlations. Thus, the P–T curves were also calculated according to the N-629 and N-631 ASME Code Cases which partially include T0 through RTT 0 . Finally, to evaluate the conservatism of the different methodologies (ASME Code and Code Cases), the analysis was performed by means of the FITNET FFS procedure that consistently incorporates the MC concepts. The study demonstrates the conservatism of the ASME code which is manly due to the definition of the semi-empirical RTNDT index. Ó 2009 Elsevier Ltd. All rights reserved.

1. Introduction and scope of the research The regulations requiring the imposition of P–T limits on the reactor coolant pressure boundary for Spanish nuclear power plants (NPPs) designed in the USA are given by the federal regulation 10CFR50 [1]. This law establishes that fracture toughness requirements for ferritic materials must fulfil the acceptance and performance criteria of Appendix G of Section 3 of the ASME Boiler and Pressure Vessel Code [2]. The fracture resistance of the vessel material in the DBT region before irradiation is described by the reference temperature RTNDT(U) obtained from Charpy impact and Pellini drop weight tests through semi-empirical and overconservative correlations [3]. As an alternative to this indirect methodology, the MC approach, originally proposed by Wallin [4–7], provides a reliable tool based on a direct characterisation of the fracture toughness in the DBT region. This approach is a consequence of the developments in elastic–plastic fracture mechanics (EPFM) together with an increased understanding of the micro-mechanisms of cleavage fracture. The basic MC method for analysis of brittle fracture test results is defined in ASTM E1921-05 [8] where a new reference temperature T0 is proposed. This is defined as the temperature at which the median fracture toughness obtained with B = 25.4 mm thickness (1T) specimens is 100 MPa m1/2. The reference temperature T0 completely characterises the fracture toughness in the DBT region of ferritic steels that experience onset of cleavage cracking at elastic or * Corresponding author. E-mail address: [email protected] (D. Ferreño). 1350-6307/$ - see front matter Ó 2009 Elsevier Ltd. All rights reserved. doi:10.1016/j.engfailanal.2009.06.007

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elastic–plastic KJc instabilities. Ferritic steels are typically carbon, low-alloy, and higher alloy grades. Typical microstructures are bainite, tempered bainite, tempered martensite, and ferrite and pearlite. To enable the use of the MC technology without completely modifying the structure of the ASME Code, a different approach was proposed, as stated in Code Cases N-629 [9] and N-631 [10] (the contents of both codes are equivalent; the only difference is that the former is applied by Section 1.1 of ASME Code [11] whereas the latter is applied by Section 3 [2]). It consists of defining a new index temperature, RTT 0 as an alternative to RTNDT. RTT 0 can be used if T0 has been previously obtained (see details in Section 1.1.). The predictions of these approaches have been compared in this paper for the vessel steel of the Spanish boiling water reactor (BWR) NPP of Santa María de Garoña, currently in service. Moreover, to quantify the level of inherent conservatism, the P–T curves (which relate the maximum allowable pressure as a function of temperature in order to avoid any risk of brittle fracture of the vessel) have also been obtained following the FITNET FFS procedure [12,13] which consistently incorporates the MC approach. To achieve these goals, the fracture properties of the non irradiated LT-oriented (according to ASME nomenclature, see [14]) base metal of the vessel in the DBT region, experimentally obtained in [15], were made available together with the complete information coming from the surveillance program, provided by the plant. In accordance with the above considerations, the main goal of this study consists of applying the experimental available results [15] to the structural integrity assessment of the vessel, comparing the conventional methodologies (ASME Code and Code Cases) with the MC approach (FITNET FFS procedure). A detailed description of the theoretical and analytical tools used in this paper is presented in Sections 1.1 and 1.2. 1.1. Description of fracture toughness in the DBT region The reference temperature RTNDT(U), mentioned above, is used to index two generic curves, developed in 1973, provided by the ASME Code [2,11] relating toughness vs. temperature, see expressions (1) and (2): the KIc curve describes the lower envelope to a large set of KIc (quasi-static fracture toughness) data whereas the KIR is a lower envelope to a combined set of KIc, KId and KIa (respectively, quasi-static, dynamic and crack arrest fracture toughness) data, being, therefore, more conservative than the former. Two important features can be appreciated: first, in both cases linear-elastic fracture mechanics (LEFM) is considered, second, the typical large scatter in the DBT region is removed by considering lower envelopes. Consequently, the method provides high conservatism in most cases.

K Ic ðMPa  m1=2 Þ ¼ 36:45 þ 22:766  e0:036½Tð



CÞRTNDT ð CÞ

K IR ðMPa  m1=2 Þ ¼ 29:40 þ 13:776  e0:0261½Tð



CÞRTNDT ð CÞ

ð1Þ ð2Þ

Alternatively, the MC approach assumes a dependence between elastic–plastic fracture toughness, KJc, with temperature in the DBT region for a given cumulative failure probability, Pf, which is given by formula (3). This expression is completely determined once T0 has been determined. With this tool, 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 particular case the expression for the median fracture toughness (Pf = 0.5), see Eq. (4), is determined.

  K Jc ;Pf ¼ K min þ ½ lnð1  Pf Þ0:25  11 þ 77  e0:019ðTT 0 Þ

ð3Þ

K Jc ðmedÞ ¼ 30 þ 70  e0:019ðTT 0 Þ

ð4Þ

To find the optimum value of T0 for a particular set of experimental data (toughness KJc vs. test temperatures, T) the maximum likelihood (MML) method described in the ASTM standard [8] should be used. It is a matter of fact that, as specimen thickness increases, the toughness is reduced, due to the higher probability of finding a critical particle for the applied load. Eq. (5), provided by the ASTM standard [8], represents one of the main contributions of the method, allowing data from different size specimens to be compared. For this reason, data must be thickness adjusted to the reference specimen thickness B = 25.4 mm before using the MML method.

K Jc ;2 ¼ K min þ ðK Jc ;1  K min Þ 

 14 B2 B1

ð5Þ

The procedure can be applied either to a single test temperature or to a transition curve data, Ti being the generic temperature of the different tests. In the latter approach (the former is a particular case) T0 is estimated from the size adjusted KJc data using a multi-temperature randomly censored maximum likelihood expression, MML. It should be mentioned that 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 are much less demanding than those of the LEFM [16]. These remarks are of great relevance in nuclear reactor surveillance programs where the amount of material available is usually very limited and consisting of small size samples (usually Charpy-V notched, CVN, specimens). In the cleavage regime for ferritic steels, the MC approach has been further developed to provide improved estimates of lower bound fracture toughness [17]. FITNET FFS [12] includes a modified MML estimation procedure [17] to obtain a

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reasonable conservative estimate of T0 and therefore determine the fracture toughness Kmat in the DBT region. The procedure involves a series of three steps to ensure that the Weibull toughness distribution fitted to experimental data is conservative. It is recommended to follow the three steps when the number of available tests is between 3 and 9 (therefore, even with a small sample it is possible to obtain a reliable estimation of T0) or when material inhomogeneities are present. In the rest of cases, the two first steps are sufficient for a conservative description. The first one, normal MML estimation, is essentially equal to that described above, proposed in the ASTM Standard [8]. The second, referred to as lower tail MML estimation, ensures that the estimate is biased towards the lower tail of the toughness distribution. The final step, step 3 referred to as minimum value estimation, requires an estimate using the minimum fracture toughness value in the data set. ASME code cases N-629 [9] and N-631 [10] define a new index temperature, RTT 0 , for the KIc and KIR ASME curves (1, 2), as an alternative to RTNDT, given in Eq. (6). RTT 0 is set (see [18]), by imposing that the ASME KIc curve indexed with RTT 0 in place of RTNDT will bound the majority of the actual material fracture toughness data. In this sense, RTT 0 was set such that the corresponding ASME KIc curve falls below the MC 95% confidence bound for at least 95% of the data generated with 1T specimens.

RTT 0 ¼ T 0 þ 19:4 C

ð6Þ

1.2. Description of structural integrity analytical tools The structural integrity assessment of a NPP vessel according to the ASME Code [2,11] requires the operation limit P–T curves to be calculated. The code provides expressions for the stress intensity factors (SIFs) considering a very conservative postulated defect in the vessel, schematically represented in Fig. 1 (the depth of the crack is 1=4 of the thickness of the vessel, t), which can be either axially or circumferentially oriented, at both the inside or outside surfaces of the vessel. The worst condition must be selected for the analysis. In shell and head regions, remote from discontinuities, the only significant loadings are general primary membrane stresses due to pressure, and thermal stresses due to thermal gradient through the thickness during start up and shutdown operations in the vessel. ASME Code [2] recommends a safety factor of 2 to be applied to the stress intensity factor produced by primary stresses. This leads to a purely LEFM requirement (see Eq. (7), to be satisfied and from which the allowable pressure for any assumed rate of temperature change can be determined.

2  K Im þ K It < K IR ðTÞ

ð7Þ

Code Case N-588 [19] provides expressions for calculating both membrane tension and radial thermal gradient SIFs. In the latter case, the formulation also covers the inside or outside surface defect by means of suitable analytical expressions; these require the thermal stress distribution at any specified time during the heat up or cooldown to be obtained as a polynomial function of the third degree, (see Eq. (8)):

t/4 3t/2

Ri

t

t t/4

3t/2 Fig. 1. Postulated defect by the ASME Code [2,11].

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rðxÞ ¼

3 X

Cj 

j¼0

xj a

¼ C0 þ C1 

x a

þ C2 

x2 a

þ C3 

x3

ð8Þ

a

where x is a dummy variable that represents the radial distance from the appropriate (inside or outside) surface, and a is the maximum crack depth. The coefficients C0 C1, C2 and C3 allow the SIF to be determined with the expressions provided in [19]. FITNET FFS approach [12] is based on quite different basis. As stated in [20] the underlying principles of the FITNET method concerning structural integrity assessments, which are of relevance for this study, are:  A hierarchical structure based on the quality of available data input which implies decreasing conservatism with increasing data quality.  Detailed guidance on determination of characteristic input values such as fracture toughness.  Possibility to perform the assessment in terms of a failure assessment diagram (FAD) or crack driving force diagrams (CDFD), thus leading to an EPFM analysis.  Compendia of solutions for stress intensity factors, and limit load solutions. Several points must be made concerning the principles summarised above. The procedure provides guidance on selection of the level of analysis, depending on the user’s available information (see [21] for details), and includes compendia of SIF solutions and limit load solutions for many geometries (some of them used in this study). Two equivalent analysis techniques are available in FITNET, namely the FAD or CDFD, the former being chosen in this research. The basis of both approaches [20] is that failure is avoided as long as the structure is not loaded beyond its maximum load bearing capacity defined using both fracture mechanics criteria and plastic-limit analysis. While both the FAD and CDFD approaches are based on elastic–plastic concepts, their application is simplified by the use of only elastic parameters. A FAD gives a 2-parameter approach to assessing a defect. It accounts for the possibility of fracture and plastic collapse simultaneously by plotting a point with coordinates (Lr, Kr) defined as follows:

Lr ¼

F FY

Kr ¼

KI K mat

ð9Þ

ð10Þ

where F is the primary applied load, FY is the limit load defined from the yield strength, KI is the SIF and Kmat the value of fracture toughness which characterises the initiation of cracking, whether by ductile or brittle mechanisms (FITNET also considers crack growth by ductile tearing, not relevant for the purposes of this paper). Depending on whether the point is inside or outside the failure assessment line (FAL), provided by FITNET, the conditions in the component are those of safety or failure, respectively, as schematically shown in Fig. 2. KI in (10) includes both the primary, KIP (due to internal pressure in the vessel) and secondary, KIS (thermal stresses) contributions:

K I ¼ K IP þ K IS

ð11Þ

Nevertheless, to obtain a reliable comparison between FITNET and the conventional assessment methodologies (ASME Code and Code Cases), the expression (11) has been slightly modified in this research. In this sense, the alternative definition (12) has been considered, which includes a safety factor of 2 in the primary SIF, in agreement with condition (7):

K I ¼ 2  K IP þ K IS

ð12Þ

FRACTURE Kr CRITICAL

SAFE

Lr Fig. 2. Appearance and significance of FAD.

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2. Material 2.1. Chemistry and microstructure The chemical composition of the steel of Santa María de Garoña NPP is given in Table 1 corresponding to a SA-336 steel, according to the ASME [22] specification. The steel presented a microstructure of ferrite with presence of bainite, being therefore suitable to be characterised in the DBT region with MC approach, (see Section 1 of this paper). 2.2. Fracture properties in the DBT region For the purposes of this study, it is necessary to know in advance the reference temperatures that define the behaviour of the steel in the DBT region. The owners of the NPP have reported a value of RTNDT = 16 °C. Moreover, in [15] an intense research in characterising the fracture toughness of this material with MC approach was performed. It consisted of testing up to 110 specimens under non irradiated and irradiated conditions with standard (pre-cracked Charpy-V notch, PCCv) and reconstituted (PCCv and compact tension, CT) specimens. For the non irradiated LT-oriented material, T0 = 98 °C was obtained, the uncertainty being measured through the standard deviation, r = 3 °C. Although the comparative structural assessment of the vessel was performed considering the non irradiated material, several remarks concerning the neutron irradiation embrittlement in reactor pressure vessel (RPV) steels must be included; this is necessary in order to properly understand the rationale in some of the arguments presented in the analysis and conclusions section of this paper. Nuclear RPV steels are degraded in the beltline region 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 increase in strength, a decrease in toughness and, as a consequence, a shift in the ductile to brittle transition temperature (DBTT). Therefore, it is necessary to know in advance the evolution of material properties with irradiation to avoid the in service failure of the vessel. The damage of the material toughness due to neutron irradiation in the DBT region is currently estimated through semiempirical methods based on the shift experienced by the CVN impact curves obtained from the surveillance capsule specimens which are retrieved periodically, according to the plant withdrawal schedule. As stated in 10CFR50 [1], the effect of neutron fluence on the behaviour of the material is predicted by Regulatory Guide 1.99 [23] which provides Eq. (13) for the evolution of RTNDT:

RTNDT ¼ RTNDT ðUÞ þ DRTNDT þ M where DRTNDT represents the shift in the reference temperature due to irradiation, which is assumed to be equal to the shift of the Charpy transition curve indexed at 41J; thus, DRTNDT = DT41J. [23]. The third term, M, is the margin that is to be added to obtain a conservative estimation. Furthermore, it is worth noting that other embrittlement predictions are available to estimate the shift in the reference temperature; among others, RG 1.99 [23] the ASTM Standard E900-02 [24] or the correlations given in the EPRI report [25] should be mentioned. The three procedures [23–25] allow DRTNDT to be obtained even when no credible surveillance data are available by means of equations based on the chemistry of the steel and on the characteristics of the neutron irradiation. In order to determine the operation limit P–T curves, the reference temperature postulated by each of the procedures was calculated, as summarised in Table 2. The two first values, RTNDT and T0, were previously presented. The third one, RTT 0 , can be directly obtained by applying expression (6). The last one, TR, was calculated with the modified MML estimation procedure [12,17] described in Section 1.1. 3. Analysis of structural integrity The P–T curves relate the maximum allowable pressure as a function of temperature so that the risk of brittle fracture in the vessel is avoided. The calculation of these curves is especially relevant in the beltline zone of the vessel, where the neutron fluence is highest. Indeed, as a consequence of the severe irradiation conditions of this region, the P–T curves can undergo an important shift with time, therefore reducing the operational window of the plant. For the vessel of Santa María de Garoña NPP, analysed in this paper, the calculation of the P–T curves for heatup and cooldown operations must be carried out according to the 10CFR50 Code [1], which refers to the ASME Code [2,11] as the procedure to be applied. The development of MC methodology has highlighted the high degree of conservatism of the conventional ASME procedure [3]. For this reason, the above described alternatives which take into account MC rudiments (Code Cases N-629 [9] and N-631 [10]), were also applied for the calculation of the P–T curves. Finally, in addition to these

Table 1 Chemical 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

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Table 2 Reference temperatures used for the calculation of the P–T curves. Methodology

Reference temperature (°C)

ASME Code MC Code Case N-629/N-631 FITNET FFS

RTNDT = 16 T0 (LT) = 98 RTT 0 = 78 TR = 92

conventional methods, the curves were also obtained according to the European structural integrity procedure FITNET FFS [12]. The relevant features of these calculations are hereafter summarised:  As the purpose of the calculations is purely comparative, the unirradiated LT material has been considered since it is the best characterised, see [15]; nevertheless, in Section 3.3 the expected irradiation embrittlement of the material is discussed for 40 and 60 year of operations of the plant.  The vessel of the Santa María de Garoña NPP was modelled in the beltline region to a cylinder with an inside radius Ri = 2.38 m and a wall thickness t = 0.12 m.  The ASME Code [2,11] suggests two representative temperature rates in RPV, 50 or 100 °F/h. The P–T curves were obtained in this research considering a cooling rate of 100 °F/h, which represents the worst condition of the two.  The ASME postulated defect (Fig. 1) was axially oriented (thus leading to the maximum membrane circumferential stress), at the inside surface of the vessel. In [15] it was demonstrated that this position corresponds to the worst operative condition.  The membrane stresses, rm, produced by the nominal pressure P = 7 MPa in the vessel for that orientation of the defect were calculated according to expression (14) which is a good approximation as the thickness is small compared to the radius where t represents the thickness of the vessel:

rm ¼

PRi t

ð13Þ

 The thermal stresses, see Section 3.1, were obtained through Finite Elements (FE) simulation.  When applying each of the procedures, their own expressions for SIF calculations were used. For consistently comparing the three alternatives a safety factor of 2 was applied to primary stresses contribution to the SIF, as stated in expression (12).  The MC methodology allows a failure probability, Pf, to be selected (see formula (3)). In this paper, based on judicious selection, a probability of Pf = 0.01 was selected.

3.1. Obtaining the thermal stresses in the vessel The thermal stress distribution at any specified time during the heatup or cooldown operations of the vessel must be obtained in order to use Eq. (8) for the calculation of the SIF, according to Code Case N-588 [19]. In this research, this task was tackled through two different methodologies, namely, an analytical and a numerical (FE) approach. The complete consistency between these two techniques was demonstrated in [15]. For the sake of simplicity, only the numerical results obtained with ANSYS FE software [26], are considered in this paper. The physical, thermal and mechanical properties of the steel of the vessel of Santa María de Garoña NPP, used for the simulations, are included in Table 3. Taking into consideration the symmetry of the problem, an axisymmetric one dimensional (1-D) model would be sufficient to perform the analysis; however, for facilitating the obtaining of the circumferential stresses a three dimensional (3-D) FE model of one quarter of the vessel was built; the details of the meshing are shown in Fig. 3. A sequential thermo-mechan-

Table 3 Physical, thermal and mechanical properties used for the FE calculations. Density, q (kg m3) Thermal conductivity (J m1 s1 K1) Film coefficient, h (J m2 s1 K1) Specific heat, c (J kg1 K1) Coefficient of thermal expansion, a (K1) Elastic modulus, E (MPa) Poisson ratio, m

7858 43.6 1106 536 1.31  105 186 0.3

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265

Fig. 3. Perspective of the meshing of the FE model.

ical simulation was performed: first, the temperature fields were obtained and from them, in a second stage, the thermal stresses were calculated. Sequential analysis requires non coupled conditions to be considered, thus assuming that the stresses do not affect the thermal distribution in the part. SOLID70 element [26] was chosen for the thermal stage of the simulation. This element has eight nodes with a single degree of freedom, temperature, at each node. The element is applicable to a 3-D, steady-state or transient thermal analysis, as in this case. SOLID45 was used for the modelling of the mechanical process, to obtain the stress fields as a function of time. This element is defined by eight nodes having three degrees of freedom at each node: translations in the nodal x, y, and z directions. Full integration method was chosen in all cases. The thermal boundary conditions consist of a convection process in the inside surface of the vessel, in contact with the reactor coolant (controlled by the film coefficient, h, in Table 3), and adiabatic conditions in the rest of surfaces. In this way, the heat flux is exclusively radial. The coolant initial temperature is 288 °C, the final temperature is 20 °C and the cooldown rate, mentioned above was chosen as 100 °F/h (56 °C/h). In Fig. 4, a plot showing the temperature profiles in the thickness of the vessel with time can be appreciated. For the mechanical boundary conditions, the axial displacements in the vessel were blocked and symmetry conditions were imposed in the vertical sections. As an example, in Fig. 5 the circumferential stresses (which are relevant for the crack

Fig. 4. Temperature profiles in the thickness of the vessel from t = 0 min to t = 300 min.

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Fig. 5. Circumferential stresses in the thickness of the vessel wall during cooling down obtained through FE.

orientation selected, see Section 3) in the thickness of the vessel wall during cooling down at four different instants (60, 120, 180 and 240 min) are represented. The position of the crack front in the thickness is superimposed in the figure. The values in Fig. 5 can be compared with the circumferential stresses for the nominal pressure in the vessel, P = 7 MPa: applying formula (14) rm = 139 MPa is obtained. In Fig. 6 the thermal SIF is represented as a function of time. As can be appreciated a plateau is reached approximately 90 min after starting the cooling down operation. For the sake of simplicity and conservativeness, the dependence of the thermal SIF with time shown in Fig. 6 was obviated in the structural assessment by taking KIt = 5.3 MPa m1/2, corresponding to the plateau. 3.2. FITNET FFS assessment FITNET FFS [12] procedure resorts to the FAD as a tool to evaluate the structural integrity. Fig. 7 shows a scheme to indicate how, from any value of pressure, P, the corresponding limit temperature, T, is evaluated to obtain the P–T curve. Next, the procedure is briefly explained. First, from the pressure, P, the applied load, F, and (using the compendia provided by FITNET) the limit load FY must be calculated; therefore, the parameter Lr can be obtained according to expression (9). Next, by using the FAL (provided by the procedure), the critical Kr is evaluated (see Fig. 5) and, from this, once KIP and KIS are known (as explained in Section 3.1, it was assumed KIS = 5.3 MPa m1/2), see formula (12), the material toughness, Kmat, is deduced. As the relation between toughness and temperature is given by (3), after selecting a cumulative failure probability (Pf = 0.01, in this research), the minimum temperature, T, necessary to avoid the failure is calculated. By repeating this process with different values of pressure, the complete P–T curve can be obtained. It must be stressed that for this analysis the Standard Option 1 of the FITNET FFS procedure [12] was chosen. 3.3. Comparison between P–T curves In Fig. 8, the P–T curves obtained through the three mentioned procedures are represented. The results clearly exhibit the excessive degree of conservatism of the ASME conventional procedure in comparison with any other methodology based on

Fig. 6. Thermal SIF as a function of time for a 100 °F/h cooling down rate.

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T

K mat

Kr

Lr

P

Fig. 7. Scheme of the calculation of P–T curves using a FAD.

the MC. FITNET FFS approach is the less conservative of the three; nevertheless, it must be stressed that this description is necessarily reliable, as the following conditions were considered: first, a safety factor of 2 in the primary SIF was imposed (thus allowing a consistent comparison with ASME procedure and Code Cases to be performed); second, based on engineering judgement, a demanding failure probability, Pf = 0.01, was selected for the assessment. Another aspect, related to the ASME postulated defect, (Fig. 1), should be mentioned to underline the reliability of the analysis. As stated in [27] ‘‘With regard to flaw indications in RPVs, there have been no indications found at the inside surface of any operating reactor in the core region, which exceed the acceptance standards of ASME Code Section 1.1, in the entire history of Section 11. This is a particularly impressive conclusion considering that core region inspections have been required to concentrate on the inner surface and near inner surface region since the implementation of U.S. Nuclear Regulatory Commission (NRC) Regulatory Guide 1.150 in 1983. Flaws have been found, but all have been qualified as buried or embedded”. Finally, for the sake of simplicity and conservativeness, the differences between mechanical constraint in the specimens used to measure fracture toughness (thus, to obtain T0) and on a flaw in the vessel were obviated. It is well known that lower constraint will increase the crack resistance curve, and vice versa. The use of fracture mechanics with highly constrained specimens, as in this paper (following the requirements in [8] to obtain T0, see [15]), leads to conservative results when the crack resistance curves are used for failure assessment analyses of low constraint geometries. Taking all these facts into consideration, it must therefore be admitted that the FITNET FFS limit curve in Fig. 8 is necessarily conservative under real operational conditions. As indicated in the figure, for the nominal pressure in the vessel, P = 7 MPa, the difference between the ASME and the Code Case N-629/N-631 solutions is about 95 °C whereas with FITNET FFS curve is up to 114 °C. To properly understand the magnitude of this difference, this can be compared with the expected shift in T41J according to RG 1.99 [23], ASTM standard E900-02 [24] or the correlations given in the EPRI report [25]. The expected fluence (E > 1 MeV) in the inside surface of the beltline region of the vessel here studied after 40 or 60 years operating is, (see [15]), respectively, 2.61  1018 n cm2 and 3.72  1018 n cm2 (corresponding to 32 and 54 EFPY – effective full power years). The foreseen shifts in T41J for these levels of fluence are included in Table 4. To perform a reliable comparison, in Fig. 9 the ASME P–T curves of the vessel at 40 and 60 years were calculated. For the nominal pressure, P = 7 MPa, the differences between the minimum allowable temperatures are, respectively, of 29 and 33 °C. According to these results, the distances between the ASME and the other P–T curves (Fig. 8), is between three and

Fig. 8. Comparison among P–T curves calculated by different methodologies.

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Table 4 Predicted T41J shift in the inside surface (beltline region) according to [23–25].

DT41J (°C) Operating time

RG 1.99 [23]

ASTM E900 [24]

EPRI [25]

40 years (32 EFPY) 60 years (54 EFPY)

23 27

20 23

20 22

Fig. 9. Effect of neutron irradiation in the P–T curves calculated with ASME Code.

four times the expected value of DT41J in the inside surface of the vessel after 60 years operating. This result clearly clarifies the high degree of inherent conservatism in the ASME procedure. The reason for such a huge difference, leading to very overconservative structural assessments, lies in the definition of RTNDT which attempts to describe the quasi-static fracture toughness in the DBT region from Pellini and Charpy impact tests. Finally, as can be observed, N-629/N-631 Code Cases provide a reliable assessment of the structural integrity of the vessel; as emphasised in Section 1.1, this fact is a consequence of the ad hoc definition of RTT 0 which is conveniently based on T0. 4. Summary and conclusions The operation limit P–T curves of the vessel of the Spanish NPP of Santa María de Garoña were obtained for a 100 °F/h cooldown operation. The study was carried out by comparing the ASME Code (based on RTNDT as reference temperature) and Code Cases N-629/631 (based on RTT 0 ) with FITNET FFS Procedure which consistently incorporates the MC approach (based on T0). To obtain a reliable comparison the structural assessment was performed by imposing in all cases the safety factors proposed by the ASME Code to the primary and secondary SIFs. For obtaining the thermal stress fields during cooldown operation, a FE model was built. The analysis of the obtained P–T curves revealed the high degree of inherent conservatism displayed by the conventional ASME approach in comparison with the other methodologies analysed. It was demonstrated that the conservativeness due to the methodology was up to between three and four times the expected material embrittlement after 60 years (extended life) operating the plant. This fact is a direct consequence of the definition of the ASME reference temperature RTNDT which is obtained from Charpy and Pellini tests through very conservative correlations to cover the uncertainties in the estimation. The prediction obtained from FITNET FFS procedure proved to be the least conservative methodology of the three analysed, in spite of demonstrating the reliability of this technique. N-629/N-631 Code Cases provide also a reliable assessment of the structural integrity of the vessel which is as a consequence of the ad hoc definition of RTT 0 which is conveniently based on T0. Acknowledgments This investigation was performed within a research project (CUPRIVA) sponsored by the Spanish Nuclear Regulatory Body (CSN) and the company UNESA. The authors wish to express particular gratitude to their colleagues Ph.D. Antonio Ballesteros and Eng. Xavier Jardí for their contribution to the dosimetry measurements and analysis. References [1] Rules and Regulations Title 10 Code of Federal Regulations Part 50.61, Appendix G. Fracture toughness requirements for protection against pressurized thermal shock events. Washington, DC: US Government Printing Office, US Nuclear Regulatory Commission; 1986. [2] ASME Boiler and Pressure Vessel Code, Section III. American Society of Mechanical Engineers, New York.

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