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International Journal of Fracture 125: 387–403, 2004. © 2004 Kluwer Academic Publishers. Printed in the Netherlands.

Predicting reference temperature from instrumented Charpy V-notch impact tests using modified Schindler procedure for computing dynamic fracture toughness P.R. SREENIVASAN∗, A. MOITRA, S.K. RAY and S.L. MANNAN Materials Development Group, Indira Gandhi Centre for Atomic Research, Kalpakkam-603102, India; Fax: 91-04114-280301/280060/280081/280356; ∗ Corresponding author: e-mail: [email protected] Received 24 July 2003; accepted in revised form 18 December 2003 Abstract. Reactor pressure vessel (RPV) steels are increasingly being characterised in terms of the reference temperature T0 and the associated Master Curve (MC) Procedure, following the ASTM E-1921 standard. Though correlations have been proposed to predict the T0 from Charpy transition temperature T28J or instrumented impact test parameters like T4kN , none can be taken as a universal correlation. Here we are proposing a new correlation √ dy dy of T0 with T0Sch , where T0Sch is the reference temperature corresponding to a median KId = 100 MPa m evaluated by the ASTM E1921 procedure applied to KId vs T data, and KId has been calculated from instrumented CVN impact test data using modified Schindler relations. This will provide a reliable method for determining dy dy T0 from instrumented CVN tests alone. T0Sch provides a conservative alternative to T0 for application of the ASTM E 1921 MC procedure in dynamic situations. Since the above procedure depends only on instrumented CVN data, it will be less costly to apply (no precracking is necessary) and will also obviate the difficulties dy associated with determining T0 from precracked CVN testing (because of severe size limitations, associated scatter and signal oscillations from the mechanics of the test, there needs to be precise control over test temperatures and test velocity for obtaining valid data from limited number of specimens). The RTNDT (est) from the suggested procedure (or its modifications based on future work) will provide an acceptable alternative to RTNDT for application of the ASME KIR curve based on instrumented CVN tests alone. For low-uppershelf steels, the dy new reference temperature estimate T0.075 and its correlation to T0Sch will provide a methodology for application of MCs to such steels. Further comprehensive work is needed to validate the procedures and correlations suggested in this paper. Key words: ASTM E 1921, Charpy, CVN, dynamic fracture toughness, instrumented impact test, Master Curve (MC), KIC , KId , KIR , reactor pressure vessel, reference temperature T0 , RTNDT , tolerance bound, transition temperature.

1. Introduction Reactor pressure vessel (RPV) steels are increasingly being characterised in terms of the reference temperature T0 and the associated Master Curve (MC) Procedure following the ASTM E-1921 standard (1998). Though correlations have been proposed to predict the T0 from Charpy transition temperature T28J or instrumented impact test parameters like T4kN , none can be taken as a universal correlation (Wallin et al., 2002; Hubner and Pusch). It may be helpful to develop a method for evaluating the T0 using Charpy V-notch (CVN) specimen test data. One approach that has been applied is by using the correlation between CVN (or Charpy) energy (CV ) and static fracture toughness KIC (Kim et al., 2002) and then obtain T0 by applying the ASTM E 1921 procedures to the computed KIC ; however, CVN correlations

388 P.R. Sreenivasan et al. have limited applicability to the specific materials and test conditions for which they were developed and require further validation (Wallin, 1989; Schindler, 1996). Schindler (2000), by an analytical-empirical procedure, has derived expressions that can be used for obtaining conservative estimates of dynamic fracture toughness (KId ) from the data obtainable from instrumented impact test traces alone; later he gave simplified relations that correlate KId to CV (Schindler, 1996). The simplified relations have been used by various investigators for evaluating KId and have been found to correlate realistically with actual KId (Hubner and Pusch; Bohme and Schindler, 2000); Sreenivasan et al. (2003) used the simplified relations of Schindler with one modification (which makes the instrumented Charpy test self-sufficient, unlike the simplified Schindler relations which required input from tensile tests) for evaluating the KId of a service exposed 2.25Cr-1Mo steel. Here we are proposing a correlation of T0 with √ dy dy T0Sch , where T0Sch is the reference temperature corresponding to a median KId = 100 MPa m evaluated by the ASTM E1921 procedure applied to KId vs T data, and KId has been calculated from instrumented CVN impact test data using modified relations of Schindler’s (modified Schindler relations). For exploring the possibility of developing a correlation, data from the literature and our previous results (as detailed in the next Section) have been reanalysed in the above framework and the results presented and commented upon where possible. 2. Materials and data The materials and type of data used for the analysis presented in this paper are listed in Table 1 along with source references. The T0 and instrumented impact test data for three steels (JRQ, Steel A, and Steel B) presented in the JAERI Report JAERI-Research 2000-022 (Onizawa and Suzuki, 2000) form a sort of basic benchmark for this paper in addition to some valid KIC , CV and yield strength versus temperature data reported in the literature (Table 1). Other aspects like different test conditions in terms of different test velocities or (Stress Intensity Factor-SIF-loading rate, i.e., dK/dt) are also indicated in Table 1. The room temperature (RT) yield strength data for the various steels investigated are reported in Table 2 along with other reported transition properties. The data derived during the present investigation are indicated in bold. The relevant references may be consulted for complete details on the materials and welds. 3. K1d and J1d estimation by modified Schindler procedure For the power law J -R curve given by J = H · (a)p

(1)

Schindler (1996, 2000), by an analytical-empirical procedure, suggested a method for obtaining the constants H and p from the following equations: p η(a0 ) 2 1−p · (CV )p · Emp · (2) H= p {B · (b0 )1+p } Emp −1 3 · 1+ p= 4 CV

(3)

Variation of CV , Pgy , Pm vs T , RTNDT and DW/CVN KId Complete instrumented CVN data Complete instrumented CVN data and PCCVN JId at test velocity of 2.3 m.s−1

9Cr-1Mo weld

120,000 h service-exposed 2.25Cr-1Mo steel

ASTM A533 Cl 1 Steel (165 mm thick plate 650 ◦ C tempered)

533GTAW

A533-Kob

21IGC

(Xinping and Yaown, 1996) A508CLA & A508CLA

(Logsdon, 1982)

(Kobayashi, 1984)

(Sreenivasan et al., 2003)

91Wld-IGC

403SS-IGC

√ The K˙ for the CVN tests at the full velocity of ∼ 5 m/s computes to about 5.105 to 106 MPa m.s−1 ; unless otherwise stated, the test velocity for normal CVN impact tests corresponds to ∼ 5 m.s−1 . K˙ = dK/dt

A508CL3-A steel in two heat treatment/strentgth Cv − T variation and PCVN JId vs T conditions (179 mm thick plate: referred as at velocity: 2.08 m.s−1 A508CLA and A508CLA in (Xinping and Yaown, 1996))

KId from 1" C(T) at SIF rate K˙ = √ 4.88 · 104 MPa m s−1 ; Cv & YS vs T

(Sreenivasan et al., 1996)

Cv , Pgy , Pm vs T and DW/PCVN/CVN KId

AISI 403 SS

SA533B Gr A Cl 2 gas tungsten arc weldGTAW-(designated in (Logsdon, 1982) as B1035/A070)

(Koneczny et al., 1998)

KIC , Cv vs T

C-Mn steel (A52 Base Metal: 30 mm plate)

(Moitra et al., 1996)

(Logsdon and Begley, 1997) 403Ssmod; A217; A470; A471

KIC , Cv and YS vs T

AISI 403 (modified) SS; ASTM A217 (2.25Cr1Mo) cast steel; ASTM A470 (CrMoV) steel; ASTM A471 (NiCrMoV) steel

A52BM

(Onizawa and Suzuki, 2000) JRQ, Steel-A and Steel-B

T0 and MCs and variation of CV , Pgy , Pm vs T , RTNDT

Designation in this paper

ASTM A533B-1 Type steels: JRQ(ID: 5JRQ43), Steel A, and Steel B

Ref. Source

Data available

Material

Table 1. Materials and Data

Predicting reference temperature from instrumented charpy V-notch impact tests 389

−112 76.3 56 4

−137 26 17 −41 – – – –

– – – – –

−22.8 −11.6 39 – – –

– – – –

– – – – –

−98 −76 −65 – – –

– – – –

−97 −67 −66 −28 −89 −55.4 (−62.4∗ ) 96 −17 – – – −182 22 −3 −28

68.3 −13 −41 −20 –

−68 −54 −37 8 −44 −54

−167 34.5 8 −21

92 29 −36 −16 −41

−61 −42 −25 24 −32 −37

−149 60 24 −9

13 98 −27 −9 −35

−50 −23 −6 51 −16 −13

– – – –

– – – – −26

−32 −4.4 25.4 – – –

−92.5

20 26 −85

dy

690 670 560 440

611 911 540 893 753

462 469 488 675 400 362

T28J , ◦ C T41J , ◦ C T68J , ◦ C FATT, ◦ C T0 , ◦ C σys /MPa (RT YS)

Note: Values indicated in bold face were derived from the respective literature data during the present investigation. All other values were taken directly from the literature. ∗ : T Estimate from ASTM E 1921 single temperature test procedure. 0

165 109.5 −15.6 11.6 −18

65.5 72 −51 −8 −55

A470 – A471 – A508CLA – – A508CLA SA533−65 GTAW A533Kob – 403SS-IGC 32 91WldIGC −4 21IGC –

−35.6 0 14.2 98 −10 −13.6

−65.5 −52 −39 3 −62 −78

dy

RTNDT , ◦ C T0.075 , ◦ C, T T0sch (◦ C) T4kN , ◦ C T0 from T0 , ◦ C ◦ PCCVN, C (1TCT) for KId /σyd √ = 0.075 m

Steel-B −45 Steel-A −35 JRQ −15 403SSmod – A217 – A52BM –

Material

Table 2. Reference and transition temperatures for various steels

390 P.R. Sreenivasan et al.

Predicting reference temperature from instrumented charpy V-notch impact tests 391 where Emp is the plastic energy up to maximum load and CV is the total CVN energy, i.e., the impact energy. Though Equations (2) and (3) were originally proposed for precracked CVN (PCVN) specimens, present authors have found that for a 308 weld and 316 stainless steels Equations (2) and (3) applied to unprecracked CVN specimens give JR curves that are conservative with respect to those obtained from PCVN specimens by normal procedures (Sreenivasan et al., 2002, 2001). Schindler (1996), by extending his analysis to unprecracked CVN specimens, has suggested later that minimum JId can be obtained as the minimum obtained from the following J0 =

7.33 · m · CV · 10−3 CV 1 − 1.47 · σfd

(4)

J0.2 = 3.92 · m1/3 · CV · 10−3 J0.2t = 11.44 · m1/3 · CV ·

(5) 3.05 · CV 103 · σfd

3

2/3 · m + 0.2

· 10−3 ,

(6)

where J is in J mm−2 , m is the power-law exponent and σfd is the dynamic flow stress. m is obtained by first identifying the maximum load or crack initiation point (by key-curve or similar procedure (Sreenivasan et al., 2001, 2002)) on instrumented impact test traces. Then the P − dpl data up to maximum load or crack initiation point are fitted to a power law, P = kdplm , and is equivalent to dynamic uniform plastic strain in a dynamic tensile test. Original Schindler relations make use of uniform strain from a tensile test and the procedure suggested here makes the instrumented Charpy test self-sufficient. σfd is estimated by using Pfd (in N) = (Pgy + Pm )/2, i.e., mean of general yield and maximum loads, in the usual yield stress relation given below: σfd =

2.99 · Pfd · W , B · (W − a0 )2

(7)

where σfd is in MPa, W , B and a0 are the specimen width, thickness and crack length in mm respectively. Equation (4) gives J for a ∼ 0; Equation (5) gives J for a = 0.2 mm and Equation (6) gives J for a = 0.2 mm from the intersection of blunting line with the JR curve (see, Schindler, 1996, for further details). As, from the lower-shelf to the upper-shelf regions of the transition region, Equation (4) gives lower estimates, it is used for estimating JId . JId is converted to KId using Equation (8): (8) KId = E · JId , where E is the Young’s modulus. When load-time data are not available for estimating m, the following relations (usually used for static tensile results, but for dynamic tests replaced by the corresponding dynamic quantities) may be used (Dahl and Hesse, 1986; Hubner and Pusch): m = 10−(log(σyd )−log(60))

(9)

m can also be estimated from the ratio of σmax d /σyd (or equivalently from a ratio of Pmax /Pgy ) by an iterative procedure using Equation (10) given below (Kirk and Wang, 1995):

392 P.R. Sreenivasan et al. σmax d = (500m/2.178)m σyd

(10)

For ease of estimation, using the data from an iterative solution of Equation (10), the authors have fitted the resultant m and load-ratio data to polynomials to give: m = −1.1402 + 2.1962p − 13388p 2 + 0.292p 3

(11)

p = 0.9867 + 1.6647m + 19.1370m2 − 6.6306m3 ,

(12)

where the load ratio p = Pm /Pgy (or σmax d /σyd) usually is in the range of 0.8 to 2.0. For variation with temperature of the Young’s modulus, E, the following relation has been used (with E in GPa and temperature T is in ◦ C) (Hong and Lee, 2001): E = 207.2 − 0.0571T

(13)

Generally, it was found that Equation (9) underpredicts m when compared with data estimated from actual load-time traces and Equation (10); hence Equations (11) and (12) seem to be more realistic. When yield load data is not available, the variation of static yield stress, static flow stress, dynamic yield stress and dynamic flow stress can be estimated using the following equations, respectively (Server et al., 1977): σys = 6.895 · (73.62 − 0.0603T + 1.32 · 10−4 · T 2 − 1.16 · 10−7 · T 3 )

(14a)

σfs = 6.895 · (85.9 − 0.00671T + 5.09 · 10−4 · T 2 − 4.26 · 10−8 · T 3 )

(14b)

σyd = 6.895 · (112.95 − 0.0158T + 3.78 · 10−4 · T 2 − 3.96 · 10−7 · T 3 )

(14c)

σfd = 6.895 · (131.53 − 0.0132T + 2.68 · 10−4 · T 2 − 2.68 · 10−7 · T 3 ),

(14d)

where T is in ◦ F (◦ F =◦ C∗ (9/5) + 32) and σys is in MPa. The σys obtained from, say, Equation (14a) at room temperature (RT) is compared with the RT yield stress of the material and the ratio (σys (material)/σys (Equation (14a)) at RT) is used to scale all the yield stress values at various temperatures from Equation (14a) and the new values are taken as the values for the particular material. In the absence of instrumented CVN data, dynamic yield stress, σyd , can also be estimated from the following relation (ASTM E399-83, 1990): σyd = σys (RT ) +

666500 − 190, (T + 273)∗ log(2∗ 1010∗ t)

(15)

where t is the fracture time in ms (usually taken as 0.1 ms for impact tests at full velocity) and T is in ◦ C. Equation (10) can then be used to compute m and then Equation (12) used for converting m to load ratio and the corresponding flow stress. In the present paper, where necessary, Equations (14a) and (15) have been employed for computing the static and dynamic yield stresses and mostly Equation (11) for computing m. After estimating the KId from the Cv values by the above procedures, the resulting KId data are examined for validity by using the ASTM E 1921 equation (1997): Eb0 σys 0.5 (16) KI C ≤ 30


Figure 1. Variation of KId with temperature for four steels

where, in the case of dynamic tests, the corresponding dynamic quantities are used (for an unprecracked CVN specimen, b0 = 8 mm). Usually for the KId estimated from √ Cv values, than 100 MPa m. Then the validity extends up to the uppershelf or to much larger KId values √ KId values from the lower-shelf (only those above 50 MPa m) to the upper cut-off limit are size corrected (to 1T equivalence) using the following ASTM E 1921 equation (1997): Bx 0.25 (17) KI C(1T) = 20 + [KJ C(x) − 20] B1T As before, dynamic quantities are used for dynamic tests, and for a CVN specimen Bx = 10 mm (and B1T = 25.4 mm). Some valid KIC data from the literature for specimens larger than 1" thickness were size corrected using the appropriate Bx . Then an exponetial (for data not reaching up to the uppershelf) or Tanh (for data valid and size corrected up to the uppershelf) function is fitted to the size corrected data. The Tanh function is of the form: f = A + B ∗ (tanh((x − T )/C)),

(18)

where A, B and C are constants. √ Then the T −KIC (orKId ) pairs between 80 to 120 MPa m are selected and analysed for the dy reference temperature (T0 or T0Sch , as the case may be) using the multi-temperature equation given below (Wallin, 1997; Kim et al., 2002): 0=

i=n i=1

n δi exp{0.019(Ti −T0 )} (KJ Ci −Kmin )4 exp{0.019(Ti −T0 )} − , (19) [31−Kmin +77 exp{0.019(Ti −T0 )}] i=1 [31−Kmin +77 exp{0.019(Ti −T0 )}]5

where the √ Kronecker δi = 1 for valid data and 0 for non-cleavage or censored data and Kmin √= 20 MPa m. In case, 8 to 10 raw data points were available within the 80 to 120 MPa m, they were also directly evaluated using Equation (19).

394 P.R. Sreenivasan et al. Table 3. Various transition temperature correlations Eq. No.

R2 of the correlation

Equation Correlations developed during this investigation (T in ◦ C)

T2-1 T2-2 T2-3 T2-4 T2-5 T2-6 T2-7 T2-8 T2-9 T2-10

dy

T0 = 76.18 + 0.534T0Sch

dy T0.075 = −44.16 + 0.781T0Sch dy RTNDT = −33.34 + 0.732T0Sch dy T28J = −39.66 + 0.643T0Sch dy T41J = −37 + 0.8T0Sch dy T68J = −18.9 + 0.974T0Sch

T0 = −8.728 + 1.309T28J T0 = −39.92 + 1.155T41J T0 − T68J Correlation poor, No linear Correlation RTNDT = −0.3153 + 0.765T41J

0.97073 0.905519 0.85595 0.93870 0.92444 0.95386 0.90477 0.90988 – 0.95539

Correlations from the Literature (T in ◦ C) T2-11 T2-12 T2-13

RTNDT = 1.01T41J + 12 (Ref. (Wallin et al., 2001)) TNDT = T4kN + 3.5 (Ref. (Hubner and Pusch) T0 = 1.0873T28J − 29.677 (Ref. (Brumovsky and Wallin, 2001))

T2-14

T0

T2-15

dy

= 1.1T0st + 99(±20) (Ref. (Viehrig et al.))

dy T0 = T41J + 35(±25) (Ref. (Viehrig et al.))

– 0.8662 0.84 0.73

4. Results and discussion For illustration, Figure 1 shows the KId results obtained by the procedures detailed above for the 403SSmod, A52BM, A470 and A471 steels (see Table 1 for identification). The notable feature is that in the case√of 403SSmod, A470 and A471 steels the uppershelf is reached just at or below the 100 MPa m level. The implications of this are discussed later. Along with the √ various transition temperatures, Table 2 also lists a temperature at which KId /σyd = 0.075 m, estimated from the 1T equivalent KId curve and designated as T0.075 √. In the literature (Sreenivasan et al., 1992), the temperature for a ratio of KId /σyd = 0.075 m, where an ASTM E399 valid KId is used, has been found to coincide with the drop-weight nilductility transition temperature (NDTT or TNDT ). The various correlations obtained between dy T0Sch with other transition temperatures or T0 given in Table 2 are presented in Table 3 along with the respective correlation coefficients. Other correlations obtained, for example, T28J with other transition temperatures, are also shown Table 3. The correlations with the highest cordy relation coefficients are those between T0Sch and T0 or T0.075, i.e., Equations (T2-1) and (T2-2) dy in Table 3. For the T0-T0Sch correlation given by Equation (T2-1), 7 data points were available (see Figure 2 and Table 2), and two, those of A52BM and A471 steels, are treated as outliers, as they, if included in the fitting, brings down the correlation coefficient from ∼ 0.97 to much below 0.87 (see also the discussion in the Appendix). ASTM A471 steel has an uppershelf √ KId of less than 100 MPa m. In the case of the ASTM A52BM, except for Charpy energy (CV ) and RT YS, all properties were estimated; evaluation based on complete instrumented


dy

Figure 2. Correlation between reference temperature, T0 and T0Sch

dy

Figure 3. Correlation between T0Sch and T0.075 dy

test data might have improved the correlation. The T0.075 -T0Sch correlation (see Figure 3 and Equation (T2-2) in Table 3) should be expected because both are derived from the same data. However, importance of this correlation is that for those steels whose uppershelf KId is much √ dy below 100 MPa m, this correlation may help estimate a fictitious T0Sch which when put in the dy T0 -T0Sch correlation will enable the estimation of T0 of those steels from instrumented CVN tests alone. dy The T0 estimated for all the steels in Table 2 from the new T0 -T0Sch correlation are shown in Table 4 along with predictions from other correlations obtained now and from the literature. Other correlations, on comparison with the T0 values in Table 2, do not seem to fare much better. Equations (T2-1) and (T2-7) seem to give conservative and realistic estimates in most

dy

dy

−49.6 −47.0 −59.1 −83 −30.4 −32.5 −33.3 −56 −13.3 −19.4 −23.1 −39 36.2 18.0 37.6 18 −20.3 −24.8 −40.6 −49 −25.4 −28.6 −43.2 −46 104.9 70.1 86.0 −20 41.3 21.9 45.9 65 −24.4 −27.9 −44.6 −60 −4.2 −12.6 −25.0 −42 −29.4 −31.7 −46.4 −68 −156.7 −128.1 −115.4 −182 46.9 26.1 21.9 27 20.1 5.8 7.2 −9 −9.2 −16.4 −30.5 −42

−25 −6 11 60 4 −1 128 65 0 20 −5 −131 70.5 44 15

T0 (T2-7)

T0 (T2-8)

T0 (T2-13)

−17.4 −95.2 −97.7 −110.4 −103.6 12.6 −76.2 −79.4 −88.4 −88.4 21 −68.6 −57.2 −68.8 −69.9 65.4 −23.9 1.7 −12.2 −21.0 −7.8 −81.5 −66.3 −76.9 −77.5 32.5 −83.4 −79.4 −82.7 −88.4 214.2 11.9 80.7 66.3 44.6 78.6 −17.7 −25.7 −6.4 −43.8 – −84.5 −62.4 −81.5 −74.3 – −70.0 −34.9 −58.4 −51.4 – −85.8 – −87.3 – – −136.0 −247.0 −232.8 −227.6 – −35.4 20.1 −0.1 −5.8 – −46.3 −12.7 −30.7 −32.9 – −74.0 −45.4 −64.2 −60.1

RTNDT RTNDT RTNDT RTNDT (T68J -33) T0 T0 T0 (T2-12) (T2-11) (T2-10) (T2-3) (T2-15) (T2-14) (T2-1)

Steel-B −19.3 Steel-A −8.1 JRQ 42.5 403SSmod – A217 – A52BM – A470 – A471 – A508CLA – – A508CLA SA533-GTAW – A533Kob – 403SS-IGC – 91WldIGC – 21IGC –

Material

Table 4. Various transition temperature predictions using Eqs. in Table 3 (Eq. No. indicated in parentheses: all temperatures are in ◦ C)


Predicting reference temperature from instrumented charpy V-notch impact tests 397 of the cases: so the suggestion is to take the most conservative, that is, the highest temperature, estimate from the two (see also the discussion in the Appendix). Another correlation reported in the literature is between T41J and RTNDT temperatures, Equation (T2-11) (Wallin et al., 2001); Equation (T2-11) predictions given in Table 4 are reasonable and conservative when compared with the available RTNDT temperature results in Table 2. The RTNDT predictions from Equations (T2-3) and (T2-10) and (T68J – 33 C) are also shown in Table 4. In the absence of DW test data, the maximum of the four estimates can be taken as the conservative estimate of RTNDT (referred as RTNDT(est)) for the purpose of using the ASME KIR curve equation. Recent ASME Code Cases N-629 and N631 (applicable to ASME Sections IX and III respectively) (Code Case N-631, 1999) authorise the use of reference temperature RTT0(= T0 + 19.4◦ C) as an alternative to RTNDT for applying the ASME KIR (and also the lower bound KIC ) equations. Equation (T2-12) predictions in Table 4 are very conservative. The following ASTM E 1921 MC equations were used for computing the dynamic MCs dy dy and the 5% and 95% tolerance bounds (TBs) from T0Sch and T0 (where available): KId (median) = 30 + 70 exp(0.019(T − T0 ))

(20)

KId (5% TB) = 25.4 + 37.8 exp(0.019(T − T0 ))

(21)

KId (95% TB) = 34.6 + 102.2 exp(0.019(T − T0 ))

(22)

KIR curves were computed using the following ASME equation: KIR = 29.4 + 13.675 exp(0.026(T − RT )), dy

dy

(23)

where RT denotes RTT0 (from T0Sch or T0 ) or RTNDT or RTNDT(est). Comparisons of master curves (MCs) and KIR curves for six steels (for which independent KId values were available), namely, 533GTAW, A533Kob, A508CLA, A508CLA , 403SS-IGC and 91IGC, are presented in Figures 4 to 9 respectively along with the respective KId data. The RTNDT(est) values obtained for these steels are shown in bold in Table 4. The first thing that is evident from the MC figures is the very conservative nature of the dy dy KIR curves, whether based on RTT0 (from T0Sch or T0 ), RTNDT or RTNDT(est), the first one being the most conservative. All figures (except Figure 4) show similar behaviour. Predictions dy dy from T0Sch are more, but acceptably, conservative with reference to T0 (where available). The dy 5% TB MC based on T0Sch seem to lower-bound the KId data, and hence it can be used for dy safety assessment applications; in all cases, this 5% TB MC (based on T0Sch ) closely follows dy or lies above the KIR curve based on RTNDT (est). Hence, in the absence of DW or T0 data, the RTNDT(est) based on our procedure can be taken as a conservative alternative for application dy of the ASME KIR curve. It may be noted that estimation of T0Sch using complete instrumented impact test data (without making any assumptions for m and dynamic yield and flow stresses) in all cases would have resulted in more reliable predictions. This is a task for further research. dy Now, in Figure 4, there is a wide disparity between predictions from T0 (computed from dy mainly a strain rate effect: the actual KId data actual KId data) and T0Sch . This is likely to be √ correspond to a √ SIF loading rate of ∼ 104 MPa m/s while the CVN tests represent a SIF rate the uppershelf data for A508CLA steel seem to saturate at of 105 -106 MPa m/s. In Figure 4, √ a level above a KId of ∼ 150 MPa m, despite the fact that ASTM E 1921 validity limit lies much above for this steel as shown in the upper part of the Figure 4. Whether an uppershelf


dy

Figure 4. MC(T0Sch ) and KIR comparison for SA533-GTAW

dy

Figure 5. Comparison of MC (T0Sch ) and KIR curves for Steel A533Kob

√ cut-off (for KId values, say at 150 MPa m) should be placed (as is the case with the RCCM KId curve) or some constraint correction should be applied for acceptance of such values, dy need further investigation. Also, the T0Sch -T0 correlation equation (Equation T2-1) needs to be investigated further based on complete data from instrumented CVN tests and T0 tests on a wide variety of steels. 5. Concluding remarks √ dy • T0Sch (corresponding to a median KId = 100 MPa m evaluated by the ASTM E1921 procedure applied to KId vs T data, where KId has been calculated from instrumented


dy

Figure 6. Comparison of MC (T0Sch ) and KIR curves for Steel A508CLA

Figure 7. Comparison of MC (T0Sch ) and KIR curves for Steel A508CLA dy

CVN impact test data using modified Schindler relations) provides a conservative alterndy ative to T0 for application of the ASTM E 1921 MC procedure. • Since the above procedure depends only on instrumented CVN data, it will be less costly to apply (no precracking is necessary) and will also obviate the difficulties associated dy with determining T0 from precracked CVN testing (because of severe size limitations, associated scatter and signal oscillations from the mechanics of the test, there needs to be precise control over test temperatures and test velocity for obtaining valid data from limited number of specimens). dy • T0Sch -T0 correlation proposed here (or its modifications based on future work) provides a reliable method for determining T0 from instrumented CVN tests alone.


dy

Figure 8. Comparison of MC (T0Sch ) and KIR curves for Steel 403 SS-IGC

dy

Figure 9. Comparison of MC (T0Sch ) and KIR curves for 91WIdIGC

• The RTNDT(est) from the suggested procedure (or its modifications based on future work) provides an acceptable alternative to RTNDT for application of the ASME KIR curve based on instrumented CVN tests alone. • For low-uppershelf steels, the new reference temperature estimate T0.075 and its correlady tion to T0Sch will provide a methodology for application of MCs to such steels. • Further comprehensive work is needed to validate the procedures and correlations suggested in this paper.


dy

Figure A1. Comparison of T0 (Test) with T0 (Predicted) from T0Sch using the various fit equations.

Appendix It may be noted that for the materials considered in this paper, only for steels JRQ, Steel A and Steel B the T0 data strictly conform to the ASTM E 1921 procedure. For other steels, T0 has been obtained from actual KIC data determined following ASTM E 399 procedure using even 4T to 6T specimens. These data may have much more than 0.2 mm crack extension, for ASTM E 399 permits upto 2% crack extension based on initial ligament depth (usually ∼ W/2 for specimens with a/W ratio of 0.5). So from the point of view of certainty of data, one should strictly correlate only data for the above three steels. Based on trial and error, it was found that excluding the data for A 52 BM and A470 steels yields much improved correlation and also the data (after excluding the A 52 BM and A470 steels) tend to lie in the scatterband for the above three steels. More importantly, the work in this paper is only of an exploratory nature, and to fully examine the validity and acceptability of the suggested procedure, evaluation must be done on a variety of steels using dynamic and static tests on PCVN and CVN specimens availing of complete instrumented test data so that no assumptions are made regarding variation of yield stress and work-hardening with temperature. For the sake of improved clarity, predictions from Eqs. (T2-1), (T2-7), (T2-8) and (T213) have been plotted against T0st (test) in Figure A1 (Appendix). Excepting the case of A470 steel, in all other cases, prediction from either Equation (T2-1) or Equation (T2-7) is the most conservative, i.e., gives the highest value. Hence, it was suggested that the most conservative value from Equations (T2-1) and (T2-7) be taken. In Figure A1, predictions from two more equations, Equations N1 and Equation N2 have been presented. Equation N1 is a fit to all the data in Figure 2, while Equation N2 is a fit obtained by excluding only A470 steel. The respective equations are as given below: dy

T0st = −72.62 + 0.772T0Sch (R2 = 0.8542)

(N1)

402 P.R. Sreenivasan et al. dy

T0st = −72.89 + 0.5T0Sch (R2 = 0.8677)

(N2)

It may be noted that Equation N2 predicts values that are mostly coincident with those from Equation (T2-1), while those from Equation N1 are not more conservative than those from Equations (T2-1) and Equation (T2-7). Recently the authors have analysed T0st (test) results from PCVN specimens for five different A533 or A508 Type steels and found that the above observations with respect to Equation (T2-1) or Equation (T2-7) are valid: excepting in one case, there is an underestimation which is less than 15 ◦ C; the exceptional case showed an overestimation of less than 5 ◦ C. In Figure A1 also, the ±15 ◦ C lines are shown along with the 1:1 line. At least for the low alloy pressure vessel steels, the present method seems to be promising. References ASTM E1921-97. (1998). Standard test method for determination of reference temperature, T0 , for ferritic steels in the transition range. Annual Book of ASTM Standards, Vol. 03.01 pp. 1060–1076. ASTM, Philadelphia, USA. ASTM E399-83. (1990). Annual Book of ASTM Standards. Vol. 03.01, ASTM pp. 488–512. Bohme, W. and Schindler, H.J. (2000). Application of single-specimen methods on instrumented Charpy tests: results of DVM round-robin exercises. Ibid, as in (Schindler, H., -J., 2000), pp. 327–336. Brumovsky, M. and Wallin, K. (2001). Status of CRP-IV (‘Assuring Structural Integrity of Reactor Pressure Vessels’: 1995–1999) on Master Curve: Brief Summary and Results. Presented at the IAEA Specialists Meeting on ‘Master Curve Testing and Application’, Prague, Czech. Republic, Sept. 17-19. Code Case N-631 (1999). (For ASME Section III Applications) and Code Case N-629 (For ASME Section XI Applications). ASME. Dahl, W. and Hesse, W. (1986). Stahl u. Eisen. 12, 695. Hong, J.-H. and Lee, G.-S. (2001). The Korean contribution to the IAEA CRP on ‘Surveillance Programme Results Application to RPV Integrity Assessment’. In IAEA Co-Ordinated Research Programme: CRP V‘Surveillance Programme Results Application to RPV Integrity Assessment’, 2nd RCM, 12-14 Sept. 2001. Prague. Hubner, P. and Pusch, G. Correlations between Charpy energy and crack initiation parameters of the J-integral concept. Ibid as in (Wallin et al., 2002), pp. 289–295. Kim, S.H., Park, Y.W., Kang, S.S. and Chung, H.D. (2002). Estimation of fracture toughness transition curves of RPV steels from Charpy impact test data. Nucl. Engng. & Design 212, 49–57. Kirk, M.T. and Wang, Y.-Y. (1995). Wide range CTOD estimation formulae for SE(B) specimens. Fracture Mechanics, Vol. 26, ASTM STP 1256. (Eds.) W.G. Reuther et al. ASTM, Philadelphia. Kobayashi, T. (1984). On the information about fracture characteristics obtained from instrumented impact test of A533 steel for reactor pressure vessel. Engng. Fracture Mechanics 19(1), 67–79. Koneczny, H., jjanosch, J.J. and Debiez, S. (1998). Definition of a ‘materials’ database for heavy structures by experimental characterisation of toughness-impact strength. Nucl. Engng. & Design 185, 203–219. Logsdon, W.A. (1982). Dynamic fracture toughness of heavy section, narrow gap, gas tungsten arc weldments. Engng. Fracture Mechanics 16(6), 757–767. Logsdon, W.A. and Begley, J.A. (1997). Upper shelf temperature dependence of fracture toughness for four low to intermediate strength ferritic steels. Engng. Fracture Mechanics 9, 461–470. Moitra, A., Ray, S.K., Mannan, S.L. and Chandramohan, R. (1996). Dynamic fracture toughness properties of a 9Cr-1Mo weld from instrumented impact and drop-weight tests. P.R. Sreenivasan, Int. J. Pres. Ves. and Piping 69, 149–159. Onizawa, K. and Suzuki, M. (May 2000). JAERI’s Contribution to the IAEA Coordinated Research Programme on ‘Assuring Structural Integrity of Reactor Pressure Vessels’ (CRP-IV), Final Report.JAERI-Research: 2000022. Japan Atomic Energy Research Institute. Schindler, H.-J. (1996). Estimation of the Dynamic J-R curve from a Single Impact Bending Test, Proceedings of the 11th European Conference on Fracture, Poitiers, EMAS, London, pp. 2007–2012.

Predicting reference temperature from instrumented charpy V-notch impact tests 403 Schindler, H.-J. (2000). Relation Between Fracture Toughness and Charpy Fracture Energy: An analytical Approach, Pendulum Impact Testing: A Century of Progress, ASTM STP 1380, ASTM, West Conshohocken, PA, pp. 337–353. Server, W.L., Oldfield, W. and Wullaert, R.A. (Principal Investigators) (May 1977). Experimental and statistical requirements for developing a well-defined KIR curve. Final Report-EPRI NP-372 (Research Project 696-1) (Prepared by Fracture Control Corporation). Electric Power Research Institute, Palo Alto, California, USA. Sreenivasan, P.R., Shastry, C.G., Mathew, M.D., Bhanu Sankara Rao, K., Mannan, S.L. and Bandyopadhyay, G. (April, 2003). Dynamic Fracture Toughness and Charpy Transition Properties of a Service Exposed 2.25Cr1Mo Reheater Header Pipe. J. Engng. Materials Tech. (Trans. ASME) 125, 221–233. Sreenivasan, P.R., Ray, S.K., Mannan, S.L. and Rodriguez, P. (1996). Dynamic fracture toughness and Charpy impact properties of an AISI 403 martensitic stainless steel. J. Nucl. Materials 228, 338–345. Sreenivasan, P.R., Ray, S.K. and Mannan, S.L. Dynamic JR Curves of 308 Stainless Steel Weld from Instrumented Impact Test of Unprecracked Charpy V-Notch Specimens. Ibid as in (Wallin et al., 2002), pp. 315–323. Sreenivasan, P.R., Ray, S.K. and Mannan Dynamic, S.L. (2001). JR Curves from Instrumented Impact Test of Unprecracked Charpy V-Notch Specimens of Austenitic Stainless Steel. (REFERENCE: ICF100359PR). Poster Paper Int. Conf. On Fracture:ICF-10. Honolulu, Hawai, 2-6 December (2001). Proc. ICF-10 (Eds. K. Ravi-Chandar, B.L. Karihaloo, T. Kishi, R.O. Ritchie, A.T. Yokobori, Jr), Elsevier Science Ltd., Oxford (UK). Sreenivasan, P.R., Ray, S.K., Mannan, S.L. and Rodriguez, P. (1992). Determination of KId at and below NDTT using instrumented drop-weight testing. Int. J. Fracture 55, 273–283. Viehrig, H.-W., Boehmert, J. and Dzugan, J. Use of instrumented Charpy impact tests for the determination of fracture toughness values. Ibid as in (Wallin et al., 2002), pp. 245–252. Wallin, K. (1997). Small specimen fracture-toughness characterisation-state of the art and beyond. In: Proc. Advances in Fracture Research, ICF-9, (Eds. B.L. Karihaloo, Y.W. Mai, M.I. Ripley and R.O. Ritchie), Amsterdam, pp. 2333–2344. Wallin, K. (1989). A simple theoretical Charpy V-KIC correlation for irradiation embrittlement. In: Innovative Approaches to Irradiation Damage and Fracture Analysis. D.L. Marriott et al. (Eds.). (ASME Pres. Ves. & Piping Conference, Honolulu, Hawai, (July 1989)). PVP-Vol.170, ASME: New York pp. 93–100. Wallin, K., Rintamaa, R. and Nagel, G. (2001). Conservatism of the ASME KIR -reference curve with respect to crack arrest. Nucl. Engng. & Design 206, 185–199. Wallin, K., Nevasmaa, P., Planman, T. and Valo, M. (2002). Evolution of the Charpy-V test from a quality control test to a materials evaluation tool for structural integrity assessment. From Charpy to Present Impact Testing (Selected papers from Charpy Centenary Conference-2001, Poitiers, France, Oct. 2–5, 2001) (Eds.) D. Francois and A. Pineau. ESIS Publication No. 30. Elsevier Sci. Ltd. And ESIS, Oxford, UK pp. 57–68. Xinping, Z. and Yaown, S. (1996). How to obtain a complete dependence curve of impact toughness or fracture toughness vs temperature on nuclear pressure vessel steels by using only one Charpy specimen (Technical Note). Int. J. Pres. Ves. and Piping 65, 187–192.