Fracture characterisation of a nuclear vessel steel under dynamic conditions in the transition region

Engineering Failure Analysis 17 (2010) 464–472

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Fracture characterisation of a nuclear vessel steel under dynamic conditions in the transition region D. Ferreño a,*, R. Lacalle a, I. Gorrochategui b, F. Gutiérrez-Solana a a

University of Cantabria, ETS Ingenieros de Caminos, Av/Los Castros s/n, 39005 Santander, Spain Centro Tecnológico de Componentes (CTC), ETS de Ingenieros Industriales y Telecomunicaciones, CDTUC, Universidad de Cantabria, Av. Los Castrosl s/n, 39005 Santander, Spain b

a r t i c l e

i n f o

Article history: Received 16 June 2009 Accepted 2 September 2009 Available online 8 September 2009 Keywords: Master curve Dynamic loading Instrumented Charpy test

a b s t r a c t The Master Curve (MC) methodology, originally proposed by Kim Wallin, is a standardised engineering tool for analysing the fracture toughness of ferritic steels in the ductile to brittle transition (DBT) region by means of the reference temperature T0. This temperature is normally estimated from quasi-static fracture toughness tests, nevertheless, it has been recently extended to the determination of dynamic fracture toughness. The aim of the present contribution is to characterise the fracture resistance in the DBT region under high strain rate conditions by applying the MC methodology to the steel of the Santa María de Garoña Spanish nuclear power plant (NPP). In this sense, 15 Charpy instrumented tests were performed on pre-cracked specimens from the surveillance program of the plant. The dynamic reference temperature, T0,dyn, was obtained and compared with the quasi-static reference temperature, T0,sta. The reliability of a semi-empirical formula proposed by Wallin to obtain T0,dyn from T0,sta has been analysed for this material. Ó 2009 Elsevier Ltd. All rights reserved.

1. Introduction and scope of this research The law 10CFR50 [1] establishes that fracture toughness requirements for ferritic materials must fulfil the acceptance and performance criteria of Appendix G of Section III of the ASME Boiler and Pressure Vessel Code [2]. Several normalised scenarios are considered requiring the imposition of pressure–temperature limits on the reactor coolant boundary with sufficient margin to assure that, when stressed under operating, maintenance and testing, the boundary behaves in a nonbrittle manner and the probability of rapidly propagating fracture is minimised. In this sense, the pre-service hydrostatic test, inservice leak and hydrostatic test, heatup and cooldown operations or core operation conditions must be analysed to ensure the structural integrity of the vessel. In all cases, quasi-static loading conditions are considered. Regarding this point, extreme loading situations such as accidental conditions are of special interest in design and structural integrity. The fracture mechanical safety assessment has to cover all possible loading conditions not only under normal service but also with (dynamic) accidental conditions; as a consequence, appropriate fracture mechanics material properties must be determined, in particular considering that fracture toughness can be strongly affected by loading rate, the dynamic toughness being lower than the quasi-static toughness. All nuclear utilities have a surveillance program 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 radiation dose rates are relatively high), capsules holding specimens fabricated with the same steel as that of the vessel. The objectives of a reactor vessel surveillance

* 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.2009.09.001

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program are twofold: first, to monitor changes in the fracture toughness properties and second, to make use of the data obtained to determine the conditions under which the vessel can be operated throughout its service life. For this research, 15 Cv specimens of nonirradiated LT-oriented material (see ASTM nomenclature in the standard [3]) from the surveillance program of the Spanish NPP of Santa María de Garoña were available. The main purpose of this paper consists in investigating the effect of the loading rate on the fracture behaviour in the DBT region by means of instrumented Charpy impact tests performed on pre-cracked specimens. The experimental dynamic results were analysed according to the MC method, successfully applied to the structural assessment of RPV under quasi-static loading conditions. 2. Fracture toughness and MC concept The MC approach, originally proposed by Kim 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 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. The main features and advantages of the method are hereafter summarised. The mathematical and empirical details of the procedure are available in [9,10]: 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 (1). This expression is completely determined once T0 has been calculated. 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. (2)) is determined.

K Jc ;Pf ¼ K min þ ½ lnð1  Pf Þ0:25  ½11 þ 77  e0:019ðTT 0 Þ  K Jc ðmedÞ ¼ 30 þ 70  e

0:019ðTT 0 Þ

ð1Þ ð2Þ

To find the optimum value of T0 for a particular set of experimental data (toughness KJc vs. test temperatures, T), the fitting procedure described in the ASTM standard [8] must be applied. It is a well-known 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. (3), 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, before fitting the data, all data must be previously thickness adjusted to the reference specimen thickness B = 25.4 mm.

K Jc ;2 ¼ K min þ ðK Jc ;1  K min Þ 

 14 B2 B1

ð3Þ

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. To estimate the reference temperature, T0, a previous censoring of the data must be applied. Fracture toughness data that are greater than the validity limit given by Eq. (4), as defined in [8], are reduced to the validity limit, KJc(lim) and treated as censored values in the subsequent estimation stage. This condition is imposed to guarantee high constraint conditions in the crack front during the fracture process.

K JcðlimÞ

sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi E  rY  b 0 ¼ 30  ð1  m2 Þ

ð4Þ

In Eq. (4) rY is the yield strength at test temperature, E is the modulus of elasticity, b0 is the initial ligament and m is the Poisson modulus. Moreover, any test that does not fulfil the imposed requirement for crack front straightness or that terminates in cleavage after more than a limit of slow-stable-crack growth will be regarded as invalid. The standard deviation in the estimate of T0 is given by:

b r

rð CÞ ¼ pffiffiffi

ð5Þ

where r represents the total number of valid specimens (not censored results) used to establish T0. The values of the factor b are provided in [8].

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The other contribution of the method that should be pointed out is 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 [11]. These remarks are of great relevance in nuclear reactor surveillance programs where the amount of material available is usually very limited and consists of small size samples (usually Charpy-V notched, CVN, specimens). 3. Dynamic fracture toughness According to the standards [8], the value of T0 is normally estimated for quasi-static loading rates: the standard requires the specimens to be loaded at a rate such that, during the initial elastic portion, the stress intensity factor rate, K_ I ¼ dK I =dt, is between 0.1 and 2 MPa m1/2 s1. Nevertheless, in principle, the existence of nonquasi-static loading rates for the nuclear power plant applications cannot be discarded. For example, high rates of loading could be found in the pressurised thermal shock (PTS) scenario transients. The use of the MC concept has been successfully extended also to the determination of dynamic fracture toughness. Wallin [12,13] demonstrated, by analysing fracture toughness data corresponding to different loading rates, that the shape of the MC is essentially unaffected by the loading rate and that the reference temperature obtained from dynamic fracture toughness data exhibits a systematic shift in T0 from that of quasi-static tests. Dynamic tests performed by Joyce [14] reached the same conclusions. Wallin [13] has recently proposed a semi-empirical methodology to estimate the dynamic reference temperature, T0,dyn, from the quasi-static value, T0,sta. In this context, as in [8], the fracture mechanical loading rate is expressed through the stress intensity factor rate, K_ I . By applying the Zener–Holloman strain rate parameter to T0, it is possible to obtain the following expression (6):

T 0;dyn ¼

C T 0;sta C  ln K_ I

ð6Þ

or, equivalently for the loading rate induced temperature shift, DT0 (7):

DT 0 ¼ T 0;dyn  T 0;sta ¼

ln K_ I

C  ln K_ I

T 0;sta

ð7Þ

By analysing fracture toughness data published in the literature [13], it was found that the parameter C can be successfully described from the yield strength, rY, and the quasi-static transition temperature, T0,sta according to expression (8):

C ¼ 9:9  exp

" # 1:66  T 0;sta rY 1:09 þ 190 722

ð8Þ

In (6)–(9) T0,sta must be expressed in absolute scale (Kelvin units). It should be pointed out that the numerical fitting parameters in expression (8) were determined from stress intensity factor rates up to 105 MPa m1/2 s1. Other authors [15,16] have validated the general procedure even for higher energy rates. These loading conditions correspond to high impact velocities, thus avoiding the technique, present in the literature, consisting of reducing the impact energy, typically from 300 J (impact velocity 5.5 m/s) to 70 J (2.8 m/s). The measurement of dynamic fracture toughness by means of instrumented pre-cracked Charpy-V notch (PCCv) impact tests, as in this research, is complicated and has not yet been fixed by a standard, although there are code recommendations [17–19], Two methods have been adopted in this paper to estimate the dynamic fracture toughness values, as recommended in [20]. If the crack initiation occurs at peak load without general yielding as shown in Fig. 1a, linear-elastic fracture mechanics (LEFM) is employed to estimate the JId (dynamic critical J integral) values [21]. Moreover, when the crack initiation occurs

Load, P

Load, P

Deflection, Δ

Deflection, Δ

(a)

(b)

Fig. 1. Schematic description of purely linear-elastic (a) and elastic–plastic (b). Charpy instrumented test load vs. deflection curves.

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with general yielding, as schematically shown in Fig. 1b, EPFM is used to evaluate the JId values [22,23]. In this sense, the crack initiation point is determined by means of the compliance changing rate method [22,23], originally proposed by Kobayashi. The compliance changing rate is defined according to:

DC C  C el ¼ C C el

ð9Þ

where DC/C is the compliance changing rate, C is the secant compliance (mm/N), and Cel is the elastic compliance (mm/N). When DC/C values are plotted against the deflection, a sudden increase in the slope of the curve can be seen as represented in Fig. 2; this point represents the crack initiation. The crack initiation energy can be obtained directly from the energy data corresponding to the deflection at the crack initiation point, obtained previously. The JId values are calculated using the equations proposed in the ASTM E1921 standard [8]. As is well-known in the case of dynamic tests, the load–deflection traces are associated with oscillations that complicate the assessment procedures of fracture toughness determination. Therefore, the load–displacement curves were smoothed before determining the plastic work to fracture (in Fig. 1 the dotted lines represent the smoothed curves). It should be pointed out that, to obtain a reliable reference temperature value, the conditions imposed in [8] were considered; thus, a KJc datum is invalid if the specimen exceeds the KJc(lim) requirement (4), or if a test has been discontinued at a value of KJ without cleavage fracture after surpassing KJc(lim). Moreover, for tests that terminate in cleavage showing slowstable-crack growth, resulting KJc value were also regarded as invalid, as mentioned above. In all cases macrographs and fractographs were taken to study the fracture surface and ensure that quasi-cleavage fracture mode was exhibited. 4. Experimental 4.1. Material 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 [24] specification. The steel presented a microstructure of ferrite with presence of bainite, therefore being suitable for characterisation in the DBT region with the MC approach (see Section 2 of this paper). For the purposes of this study it is necessary to know in advance the quasi-static reference temperature of the steel. In [10] an intense research to characterise the fracture toughness of this material with MC approach was performed. It consisted of testing up to 110 specimens under nonirradiated and irradiated conditions with standard (pre-cracked Charpy-V notch, PCCv) and reconstituted (PCCv and compact tension, CT) specimens. For the nonirradiated LT-oriented material, T0 = 98 °C was obtained, the uncertainty being measured through the standard deviation (5), r = 3 °C. 4.2. Experimental scope For this research, 15 nonirradiated LT-oriented Charpy specimens from the surveillance program of the NPP were available. Specimens were fatigue pre-cracked with an 8501 INSTRON servo hydraulic test machine at room temperature, the a/W

EL

PL

CP EL: elastic PL: plastic CP: crack propagation

Load, P

dP dΔ

Pel

Compliance changing rate

Crack initiation point

Δ el

Deflection, Δ

Fig. 2. Description of the compliance changing rate method to evaluate the crack initiation point during an instrumented Charpy impact test.

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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ratio being maintained between 0.4 and 0.5 following the recommendations in [25] for instrumented Charpy tests. The fractured surfaces were examined in all cases with micrographs and scanning electron microscopy (SEM) to analyse the mode of fracture. The instrumented impact tests were carried out on an instrumented AMSLER RKP 300 J impact pendulum. The device can record up to 2000 points during a typical test time, including the applied load, the displacement and the fracture energy. The test temperatures were selected in the range 80 to 100 °C to capture the whole DBT region. 4.3. Analysis of experimental results 4.3.1. Comparison with the Charpy impact curves from the surveillance program Fig. 3 shows the fracture energy, E, obtained from PCCv specimens vs. the test temperature; in order to properly compare with the conventional Charpy tests (performed on Charpy-V, 2 mm notched specimens, CVN) from the surveillance program, the fracture energy measured on each of the PCCv tests was corrected to take into account the differences in the resistant ligament (thus, in any test the fracture energy per unit surface was calculated and then multiplied by the surface of a CVN specimen, 80 mm2). The experimental data were fitted with a hyperbolic tangent curve and the parameter T41J was obtained, as indicated in the figure, obtaining T41J = 18 °C. Three macrographs of the fracture surface are superimposed on the figure, showing the evolution of fracture mechanisms from brittle to ductile fracture with temperature. The macrographs and the fitting show that the entire transition region was covered with the selected temperature range (80 to 100 °C), as intended. The fracture surfaces were examined by SEM to distinguish which specimens manifested cleavage as the physical fracture mechanism and to locate the initiation sites. The general aspect of the fracture surface in a valid specimen (for the subsequent MC analysis) is illustrated by an example in Fig. 4 showing the multifaceted surface with river patterns typical of cleavage fracture.

160 140

139 J 120

E (J)

100

-80 ºC

80

18ºC

60 40

0 -100

100 ºC

25 ºC

20

-50

0

50

100

150

200

250

300

T (ºC) Fig. 3. Fracture energy vs. test temperature obtained through instrumented impact tests on PCCv specimens.

Fig. 4. SEM images of the fracture surface showing the cleavage area.

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These results are compared in Fig. 5 with those from the surveillance program, consisting of CVN impact tests. In a conventional CVN test, the fracture energy is the result of a first stage to initiate the crack, followed by a fast propagation of the crack and the final fracture. The initiation is absent when testing a PCCv specimen; therefore, the resilience curves show lower energies when obtained from PCCV specimens. This difference is reduced in the lower shelf because the embrittled response of the material masks this crack effect. All these features are shown in Fig. 5. The CVN curve was also fitted with a hyperbolic tangent curve and T41J was obtained yielding T41J = 25 °C. As can be appreciated, the presence of the crack represents a T41J shift of DT41J = 43 °C and a reduction of 60 J in the upper shelf energy. 4.3.2. Obtaining the dynamic reference temperature For structural assessment, it is necessary to know the fracture toughness of the material. By using the information included in the load–displacement curves of the instrumented Charpy test, it is possible to obtain the dynamic fracture toughness and, therefore, to estimate the dynamic reference temperature, T0,dyn. In Fig. 6 a comparison between two experimental curves obtained at 40 °C and 80 °C, respectively, is presented. The former shows no plastic response prior to fracture whereas the latter is a clear example of elastic–plastic fracture. The methodology employed to analyse each of the curves to obtain the dynamic toughness JId was explained in Section 3. The results are represented in Fig. 7 as a function of test temperature. To obtain a reliable value of T0,dyn, the representative experimental results must be previously selected, discarding those showing an excessive stable-crack growth prior to fracture; the analysis of fracture surfaces revealed this feature for test temperatures above 40 °C, therefore, only those tests performed with lower temperatures were used to estimate the reference temperature. The experimental values used for the estimation (six noncensored and two censored results) are shown in Fig. 8. The curves for KJc(med) (median fracture toughness) the lower bounds KJc(0.01)–KJc(0.05), and the upper bounds KJc(0.95)– KJc(0.99) have also been plotted. With these values the result T0,dyn = 25 °C was obtained and, thus DT0 = 73 °C. The experimental result of T0,dyn was compared with the predictions of the semi-empirical model proposed by Wallin [13]. For this purpose, the stress intensity factor rate was calculated from the load vs. time (P–t) test record. First, the loading

260 220

199 J

E (J)

180

Δ (USE)= 60 J 139 J

140 100 60

-25ºC 18ºC

20

Δ T41J = 43ºC -20 -100

-50

0

50

100

150

200

250

300

T (ºC) Fig. 5. Comparison between the fracture energy curves obtained from CVN and PCCv specimens.

10

6

(a) P (kN)

P (kN)

4 3

(b)

8

5

6 4

2 2

1

0

0 0

5

10

15

Δ (mm)

20

25

0

5

10

15

20

25

Δ (mm)

Fig. 6. Comparison between experimental instrumented Charpy impact curves obtained at 40 °C (a) (linear-elastic fracture) and 80 °C (b) (elastic–plastic fracture).

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4

JId (MPa·m)

3

2

1

0 -100

-80

-60

-40

-20

0

20

40

60

80

100

120

Temperature (ºC) Fig. 7. Representation of the dynamic toughness JId as a function of test temperature.

200

T0, dyn = -25 ºC

180

KJc (0.99) KJc (0.95)

160

KJc (med)

1/2

KJc (MPa·m )

140 120

KJc,lim,1T

100 80

KJc (0.05)

60

KJc (0.01)

40

Valid datum

20

Censored datum

0 -60

-50

-40

-30

-20

-10

0

10

20

30

40

Temperature (ºC) Fig. 8. Representation of KJc vs. temperature including the most representative confidence bounds.

rate P_ (which is the slope of the linear-elastic part of the P–t curve) was obtained. Then, considering the expression of the stress intensity factor given by [11] for a PCCv specimen and applying the procedure described in formula (10), the stress intensity factor rate was calculated, yielding a rate of 6  105 MPa m1/2 s1.

KI ¼

PS BW

3=2

 f ða=WÞ

)

K_ I ¼

P_  S B  W 3=2

 f ða=WÞ

ð10Þ

where B is the specimen thickness, S the span, W the specimen width and a the crack length. Expression (8) (T0,sta = 98 °C = 175 K and rY = 397 MPa) yields C = 39.9; then, introducing C in expression (7) and the average stress intensity rate above mentioned, the reference temperature shift DT0 = 88 ± 18 °C is obtained. As can be appreciated, there is good agreement between the model and the experimental data, thus demonstrating the reliability of the model proposed by Wallin [13]. 5. Summary and conclusions In this research, 15 Charpy instrumented tests (300 J) were performed on PCCv specimens (nonirradiated LT-oriented material) from the surveillance program of the Spanish NPP of Santa María de Garoña. The comparison between the CVN curves from the surveillance program (notched specimens) and the PCCv curves (pre-cracked specimens) revealed a shift in the reference temperature DT41J = 43 °C and a reduction in the upper shelf of 60 J.

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To analyse the fracture behaviour from the instrumented dynamic tests performed on PCCv specimens, the crack initiation was obtained from the load vs. time curves, provided by the instrumented Charpy device, with two complementary methods: if the crack initiation occurred at peak load without general yielding, LEFM was applied, whereas when the crack initiation occurred with general yielding, the crack initiation point was determined by means of the compliance changing rate method proposed by Kobayashi [22,23]. A dynamic reference temperature, T0,dyn = 25 °C was obtained; taking into account the quasi-static reference temperature, T0,sta = 98 °C, reported in [10], the dynamic test conditions imply a shift in the reference temperature DT0 = 73 °C, the average applied loading rate being 6  105 MPa m1/2 s1. The following overall conclusions were obtained: s High loading rate fracture toughness tests can be performed using PCCv specimens in the transition temperature range in accordance with the method proposed in the ASTM Standard E1921-05 [8]. s The MC method is applicable under dynamic conditions data and T0 can be calculated. s Good agreement between the experimental results and the predictions of a recent semi-empirical model proposed by Wallin [13] has been found. According to this model, a shift DT0 = 88 ± 18 °C is expected for the loading conditions and material properties reported in this paper. The codes for design and assessment in nuclear vessels assume the existence of quasi-static conditions; nevertheless, the fracture mechanical safety assessment has to cover all possible loading conditions not only under normal service but also with (dynamic) accidental conditions. As demonstrated in this paper, high strain rates can undergo an important shift in T0, thus leading to a more brittle material response in the DBT region. For this reason, it is recommended to perform a detailed analysis of the existence of nonquasi-static loadings in nuclear reactors. This issue will be the subject of future research. Acknowledgments This investigation was performed within a research project (CUPRIVA) sponsored by the Spanish Nuclear Regulatory Body (CSN) and the company UNESA. 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. New York: American Society of Mechanical Engineers; 1986. [3] ASTM E1823-96 (reapproved 2002). Standard terminology relating to fatigue and fracture testing. Annual book of ASTM standards, section 3, vol. 03.02; 1996. [4] Wallin K, Saario T, Törrönen K. Statistical model for carbide induced brittle fracture in steel. Metal Sci 1984;18:13–6. [5] Wallin K. The scatter in KIC results. Eng Fract Mech 1984;19(6):1085–93. [6] Wallin K. The size effect in KIC results. Eng Fract Mech 1985;22(1):149–63. [7] Wallin K. A simple theoretical Charpy V–KIc correlation for irradiation embrittlement. Innovative approaches to irradiation damage and fracture analysis, PVP 170. ASME 1989:93–100. [8] ASTM E1921-05: Test method for the determination of reference temperature T0 for ferritic steels in the transition range. Annual book of ASTM standards, section 3, vol. 03.02; 2005. [9] Merkle J, Wallin K, McCabe D. Technical basis for an ASTM standard on determining the reference temperature, T0, for ferritic steels in the transition range. Washington, DC: Oak Ridge National Laboratory; 1998 [NUREG/CR-5504, ORNL/TM-13631]. [10] Ferreño, D. Integridad estructural de vasijas nucleares en base a la curva patrón obtenida mediante probetas reconstruidas. PhD thesis, University of Cantabria, Santander, Spain; 2008. [11] ASTM E 399-90 (reapproved 1997). Standard test method for plane-strain fracture toughness testing of metallic materials, Annual book of ASTM standards, vol. 03.01. Philadelphia: American Society for Testing and Materials; 2005. [12] Wallin K. Loading rate effect on the master curve T0. In: VTT manufacturing technology. Finland: Espoo; 1997. [13] Wallin K. Effect of the strain rate on the fracture toughness of ferritic steels. IAEA specialists meeting on master curve testing and results application, Prague, Czech Republic; 17–19 September 2001. [14] Joyce JA. On the utilization of high rate Charpy test results and the master curve to obtain accurate lower bound toughness predictions in the ductileto-brittle transition, small specimen test techniques. In: Corwin WR, Rosinski ST, Van Walle E, editors. Small specimens test technique, ASTM STP 1329. American Society for Testing and Materials; 1997. [15] Dlouhy´ I, Lenkey G, Holzmann M. Master curve evaluation at static and dynamic conditions of loading for cast ferritic steel. Transactions SMiRT 16. Washington (DC); August 2001. [16] Dlouhy´ I, Kohout J, Jurasek V, Holzmann M. Evaluation of strain rate effects on transition behaviour applying the master curve methodology. Transactions SMiRT 17, Prague, Czech Republic; August 2003. [17] Cornell P, Server WI. EPRI instrumented impact test procedures. In: Wullaert RA, editors. Proc CSNI specialist’s meeting on instrumented pre-cracked Charpy testing. EPRI-NP-2102-LD, Palo Alto, USA; 1980. p. 1. [18] Ireland DR. A review of the proposed standard method of test for impact testing pre-cracked Charpy specimen of metallic materials. In: Wullaert RA, editors. Proc CSNI specialist’s meeting on instrumented pre-cracked Charpy testing. EPRI-NP-2102-LD, Palo Alto, USA; 1980. p. 1–23. [19] Proposed standard methods for instrumented pre-cracked Charpy impact testing of steels – combined KId, JId, and CTOD tests methods. Draft international standard: ESIS, European Structure Integrity Society. Draft 8; October 1997. [20] Angamuthu K, Guha B, Achar DRG. Investigation of dynamic fracture toughness (JId) behaviour of strength mis-matched Q & T steel weldments using instrumented Charpy impact testing. Eng Fract Mech 1999;4(64):417–32. [21] Rajanna K, Bhambri K, Achar DRG. An assessment of a CrMoV cast steel weld joint for dynamic fracture behaviour. Eng Fract Mech 1988;29(4): 387–99. [22] Kobayashi T. Evaluation of dynamic fracture toughness parameters by instrumented Charpy impact test. Eng Fract Mech 1986;24(5):773–82.

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D. Ferreño et al. / Engineering Failure Analysis 17 (2010) 464–472

[23] Kobayashi T, Yamamoto I, Niinomi. Introduction of a new dynamic fracture toughness evaluation system. J Test Eval, JTEVA 1993;21(3):145–53. [24] American Society of Mechanical Engineers (ASME). Boiler and pressure vessel code, section II: materials. New York: American Society of Mechanical Engineers; 1986. [25] Instrumented impact testing. American society for testing and materials. In: Symposium of the ASTM 66th annual meeting, STP 563; 1974.