Materials Performance and Characterization

Materials Performance and Characterization P. R. Sreenivasan1

DOI: 10.1520/MPC20130079

Estimation of ASTM E1921 Master Curve of Ferritic Steels From Instrumented Impact Test of CVN Specimens Without Precracking VOL. 3 / NO. 1 / 2014

Materials Performance and Characterization

doi:10.1520/MPC20130079

/

Vol. 3

/

No. 1

/

2014

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available online at www.astm.org

P. R. Sreenivasan1

Estimation of ASTM E1921 Master Curve of Ferritic Steels From Instrumented Impact Test of CVN Specimens Without Precracking Reference Sreenivasan, P. R., “Estimation of ASTM E1921 Master Curve of Ferritic Steels From Instrumented Impact Test of CVN Specimens Without Precracking,” Materials Performance and Characterization, Vol. 3, No. 1, 2014, pp. 285–308, doi:10.1520/MPC20130079. ISSN 2165-3992

ABSTRACT Manuscript received October 21, 2013; accepted for publication May 14, 2014; published online June 23, 2014. 1

Metallurgy and Materials Group, Indira Gandhi Centre for Atomic Resaearch, Kalpakkam, Tamilnadu-603 102, India, e-mail: [email protected]

A semi-empirical cleavage fracture stress (CFS) model, mainly depending on the CFS, rf, has been derived for estimating the ASTM E1921 reference temperature (T0) and demonstrated for ferritic steels with yield strength in the range 400–750 MPa. This requires only instrumented impact test of CVN specimens without precracking and static yield stress data. The T0 estimate based on the CFS model, TQcfs, lies within a 620 C band, being conservative for most of the steels, but less conservative than TQIGC based on the IGCprocedure (see Nomenclature for definition). Applicability and acceptability of the present calibration curves for highly irradiated steels need further examination. Keywords Charpy V-notch, instrumented impact test, reference temperature, fracture toughness, cleavage fracture stress

C 2014 by ASTM International, 100 Barr Harbor Drive, P.O. Box C700, West Conshohocken, PA 19428-2959 Copyright V

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Nomenclature CFS ¼ Cleavage fracture stress, rf CV ¼ Energy absorbed by a CVN specimen during an impact test CVN ¼ Charpy V-notch DBTT ¼ Ductile-Brittle Transition Temperature. Temperature corresponding to a fixed CV, lateral expansion, or fracture appearance; for example, T28J is a DBTT d ¼ displacement experienced by the CVN specimen during IIT IGC-procedure ¼ a multi-stage correlation procedure to estimate TQIGC, ðor IGCAR procedureÞ where TQIGC is the estimate of T0 obtained using the IGCAR procedure detailed in Ref. [6]. TQIGC values are conservative to the extent of 20 C–30 C. IIT ¼ instrumented impact test KIC ¼ valid linear elastic fracture toughness as per ASTM E399 standard KJC ¼ valid linear elastic-plastic fracture toughness as per the ASTM E1921 standard [5] LTD ¼ Load Temperature Diagram; a plot of various loads from the P-d traces of several CVN specimens tested in the DBTT range plotted as a function of test temperature, with the same loads (say, PGY, PF, etc.) joined by average smooth curves, if possible. MC ¼ a standard reference fracture toughness curve for ferritic steels indexed to reference temperature, T0, as per ASTM E1921 standard [5] P ¼ a general symbol for specimen load; here, experienced by the CVN specimen during IIT PA ¼ brittle fracture arrest load on the P–d trace of a CVN IIT test record PF ¼ brittle fracture load on the P–d trace of a CVN IIT test record PGY or Pgy ¼ general yield load on the P–d trace of a CVN IIT test record Pmax or PM ¼ maximum load on the P-d trace of a CVN IIT test record T0 ¼ reference temperature determined as per ASTM E1921 standard TD ¼ the brittleness transition temperature, end of the gross elastic region in the load-temperature diagram of instrumented impact or slow-bend tests and represents almost end of 100 % cleavage fracture with PF ¼ Pmax ¼ PGY TQ ¼ Estimated T0 by a non-standard method TQIGC ¼ T0 estimated by the IGCAR-procedure Materials Performance and Characterization

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TQcfs ¼ T0 estimated by the CFS model-TQcfs is the most conservative of the four, namely, TQcfs1, TQcfs2, TQcfs3, and TQcfs4 TQcfs1 ¼ TQcfs obtained from rf/rys (as a function of temperature) and rf/rys*1 ratio TQcfs2 ¼ TQcfs obtained from rf/rys (as a function of temperature) and rf/rys*2 ratio TQcfs3 ¼ TQcfs obtained from rf/ryd (as a function of temperature) and rf/ryd*1 ratio TQcfs4 ¼ TQcfs obtained from rf/ryd (as a function of temperature) and rf/ryd*2 ratio TQBT ¼ T0 predicted from the empirical correlation of TD with T0 T28J ¼ Charpy transition temperature at which Charpy energy ¼ 28 J T41J ¼ Charpy transition temperature at which Charpy energy ¼ 41 J rf ¼ cleavage fracture stress (CFS) determined from the PF ¼ Pmax ¼ PGY loads at the temperature TD rys ¼ quasi-static yield stress, dependent on temperature ryd ¼ dynamic yield stress, dependent on temperature ryd-RT ¼ dynamic yield stress at room temperature rys-RT ¼ quasi-static yield stress at room temperature rys*1 ¼ rys at (T41J24) C rys*2 ¼ rys at (T41J50) C ryd*1 ¼ ryd at (T41J24) C rys*2 ¼ ryd at (T41J50) C

Introduction Charpy V-notch (CVN) impact test is very attractive because of its low-cost, simplicity, wide familiarity, and availability [1]. Instrumented impact test (IIT), while maintaining these advantages—in addition to the conventional Charpy energy (CV), lateral expansion (LE), and % shear fracture (PSF)—provides additional load (P)–time (t) or displacement (d) data of the CVN specimen during deformation and fracture (time can be converted to displacement). Figure 1 shows the various load parameters obtainable: general yield load-PGY (or Pgy, used interchangeably; see nomenclature), initiation load-Pinit (determined by compliance change or key-curve technique or by acoustic emission or similar sophisticated instrumentation), maximum load-PM or Pmax, brittle-fracture load-PF and arrest load-PA, and the corresponding times/displacements— say, for example, dF, the displacement to PF. The various load parameters plotted against test temperature (T) provides the load-temperature diagram (LTD) characterizing various regions of fracture [1,2]; see, for example, Fig. 2 [3]. A typical ferritic steel tested in the ductile-brittle transition temperature (DBTT) region shows characteristic P–d traces: at the lower-shelf, the fracture is purely linear-elastic with sudden brittle Materials Performance and Characterization

SREENIVASAN ON CVN SPECIMENS WITHOUT PRECRACKING

FIG. 1 Characteristic loads marked on an IIT load-time (t) trace and energy partitioning related to fracture surface of a CVN specimen.

failure occurring at the maximum load, PM (¼PF) corresponding to 100 % cleavage, while, at higher temperatures, as the test progresses to the upper shelf, PGY precedes cleavage fracture. Then PM and PF separate, with substantial ductile crack extension preceding PF and, at some temperature interval, brittle fracture started at PF arrests at PA. Ultimately, at the upper shelf, the traces do not show any brittle fracture: no PF and PA. From the various load parameters, the % shear fracture (PSF) can be obtained as a function of temperature as also the brittleness transition temperature (TD), the temperature at which PM (¼PF) ¼ PGY [2] and at higher temperatures, brittle fracture occurs after general yielding. These features are delineated clearly in Fig. 2. From the PM (¼ PF) ¼ PGY load at TD, the microcleavage fracture stress, rf, can be calculated. Moreover, from the PGY load values at various temperatures, the dynamic yield stress, ryd, as a function of temperature is obtained [1,4]. These data are very important for the present paper, as will be shown later.

FIG. 2 Load temperature (P–T) diagram for the 9Cr–1Mo BM from instrumented CVN impact tests at V0 ¼ 5.12 ms1.

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For obtaining design relevant dynamic fracture toughness (KId), however, testing of fatigue precracked CVN (PCVN) specimens is necessary. This is costly and time consuming [4]. Moreover, testing of PCVN specimens introduces problems in data reduction due to superimposed oscillations. To overcome these difficulties, testing at reduced velocity [3] (say, at  1 ms1 instead of at the usual impact test velocity of 5 ms1, with resultant loss of strain rate) or the use of complicated dynamic analysis methods have been suggested [1]. Nowadays, reactor pressure vessel (RPV) steels are increasingly being characterized in terms of the reference temperature T0 and master curve (MC) as per the ASTM E1921 standard [5]. The present author had previously proposed a multi-stage correlation procedure to estimate TQIGC, where TQIGC is the estimate of T0 obtained using the procedure (IGCAR procedure) detailed in Ref. [6]. TQIGC values are conservative to the extent of 20 C–30 C. The present paper examines in a new empirical perspective the relation of rf to fracture toughness and, thereby, tries to derive a methodology to estimate fracture toughness and master curve from the load-temperature data and Charpy energy obtained from instrumented impact test (IIT) of CVN specimens without precracking. First, the semi-theoretical-empirical basis of the present approach will be delineated in the light of previous literature. Then the method to obtain the new empirical methodology will be given. The new methodology will be applied to the calibration steels as also to many steels presented in Ref. [6] and others. The results will be compared with actual T0 or estimated TQ (as a convention, non-standard, i.e., not following the ASTM E1921 standard for master curve determination, estimates of T0 are designated TQ [6]). In addition to being fast and less costly (as no precracking is required), the new method, being a single assessment method (compared to the multi-stage method in Ref. [6]), will simplify the evaluation and hence will be less error-prone. Moreover, it will enhance the utility and purpose of the IIT of bluntnotched CVN specimens of ferritic steels. Particularly relevant is the fact that the new procedure will help obtain more valuable and design relevant master curve from IIT of irradiation surveillance specimens.

Theoretical-Empirical Basis and Methodology LITERATURE REVIEW

Based on the concept of brittle cleavage fracture occurring ahead of a crack on the attainment of a critical cleavage fracture stress (rf) over a critical distance (X0) of Ritchie et al. [7] and using the stress analysis of Hutchinson [8], Curry [9] related the cleavage fracture toughness, KIC to rf and rys in the following way: Nþ1

(1)

ðNþ1Þ 2

KIC ¼ b

1

 X02

rf 2 N1

rys2

where: b ¼ a material dependent constant (mainly a function of Ramberg–Osgood work-hardening exponent, N, and can be evaluated based on expressions given in Ref. [8]), Materials Performance and Characterization

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X0 ¼ the critical distance (depending on the microstructure, it has been related to 2–3 ferrite grain diameters in ferrite–pearlite steels, or packet or bainite size in martensitic or bainitic steels or no microstructural feature in certain steels), and rf ¼ independent of temperature and strain rate, and in most cases, even of irradiation conditions. The accurate determination of X0 for various steels is a problem. Later, Hahn et al. [10] and Kotilianen [11] empirically put KIC to rf and rys relation as: (2)

 c rf KIC ¼a rys rys

where a and c are constants for a particular steel and also depend on the temperature range of fit. Thus, the above equations have not found universal application; although Eq 1 is theoretically more satisfying. Recently, in the mesoscopic (given as mezzo-scopic in the referred paper) analysis of fracture toughness of steels, Miyata and Tagawa [12] simplified the statistical local fracture criterion approach to fracture toughness analysis of steels by both an analogous and an empirical approach. Using a two-parameter Weibull stress distribution for predicting fracture probability, they derived the expression relating KIC to Weibull Stress, rW. Then, using analogy, they replaced the Weibull stress with the cleavage fracture stress, rc—the fracture stress defined in deterministic terms (the cleavage fracture stress, rc, is defined as the local maximum principal stress at the cleavage fracture initiation in round bar tensile specimens with 1 mm radius circumferential notch) [12]. Finally, based on both fracture toughness tests and rc tests on a large number of carbon and low-alloy steels (YS range: 250–1100 MPa), they found the following empirical relation:

(3)

 g pffiffiffiffi 2:85  103 rC KC ðMPa mÞ ¼ rC ðMPaÞ 1 r 4 ys BðmmÞ

where g is fit constant for a steel (g is represented as a in Ref. [12]). Since it needs a large number of specimens to obtain the statistical Weibull parameters compared to the determination of rc, Eq 3 is a simplification. In Ref. [12], the exponent g was given as a function of T150 temperature. Here, T150 is the temperature corresponding to a fracture toughness of 150 MPaHm. The g-T150 plot in Ref. [12] shows high scatter and wide dispersion. Hence, even Eq 3 is not amenable to practical application. This paper, while retaining the simplicity of the above methods (rf/rys and rf/ryd ratio or fracture to yield stress ratio method), tries to overcome their limitations by an empirical procedure. The features of the above theoretical, semi-empirical, or empirical formulations can be summarized as follows. Basically, at the point of brittle fracture initiation, the local crack tip tensile stress reaches a critical value, namely, the microscopic cleavage fracture stress, rf or rC (the cleavage fracture stress can be determined from either notch-tensile tests—many authors denote this as rC—or three-point bend or instrumented impact tests of Charpy V-notch (CVN) specimens, mostly denoted by rf; based on consideration of differences in sampled and stressed volume for the two types of specimens, rf; especially determined from Materials Performance and Characterization

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instrumented impact tests, is slightly larger than rC from notch-tensile tests [13]) at a critical distance, usually the distance to the weakest link. In fact, a two- or three-parameter distribution of Weibull stress distribution is the basis of the master curve based on the ASTM E1921 standard [5]. The mesoscopic analysis of fracture toughness as described in the previous section implies that fracture toughness variation with temperature can be expressed as a function of critical fracture stress to yield stress ratio.

PRESENT METHODOLOGY

Based on the above considerations, the variation of the ratios, rf/rys or rf/ryd, (ryd, is the dynamic yield stress determined from instrumented Charpy V-notch specimen tests at various temperatures) are related to the relevant static MC fracture toughness data; i.e., for the same temperature range, at various temperatures, the ratio, rf/rys or rf/ryd, are evaluated along with the corresponding static MC KJC. Then the resulting, rf/rys or rf/ryd, values are plotted against the corresponding static MC KJC and a smooth curve of the following form fitted: (4)

KJC ¼ 20 þ a  expðbyÞ

where y ¼ rf =rys or rf =ryd . Equation 4 is justified because it is general practice to express variation of KJC with temperature by an equation similar to Eq 4 with rf/ryd replaced by T. Equivalently, T can be replaced by the corresponding rf/ryd (or rf/rys), which depends only on variation of ryd or rys with temperature as rf is independent of temperature. The basic methodology adopted in this paper is to fit Eq 4 based on rf/rys and rf/ryd ratios to the MC data of the 21 calibration steels (with known IIT and T0 data as described later) and determine a and b for each steel. For each steel, the 1 in. MC-KJC data are fitted to Eq 4 in the range of T0 6 50 C, as ASTM E1921 [5] MC is valid in that range. The MC equation is given by: (5)

KJC ¼ 30 þ 70  expð0:019  ðT  T0 ÞÞ

where T0 is the ASTM E1921 standard reference temperature for the material. Then an average a (aav) is determined and a second fitting done to Eq 6a or Eq 6b:   rf KJC ¼ 20 þ aavs : exp B (6a) rys   rf (6b) KJC ¼ 20 þ aavd : exp B ryd where aavs and aavd (corresponding to y ¼ rf/rys and rf/ryd, respectively, in Eq 4) are treated as constants. Then the constant B for various calibration steels based on aavs is correlated to rf/rys*1 and rf/rys*2 and B for various calibration steels based on aavd is correlated to rf/ryd*1 and rf/ryd*2, where rys*1 is the rys at (T41J – 24) C and rys*2 is the rys at (T41J – 50) C, ryd*1 is the ryd at (T41J – 24) C and ryd*2 is the ryd at (T41J – 50) C for each steel. rf/rys*1, rf/rys*2, rf/ryd*1 and rf/ryd*2 are more definitive material identifiers than the simple ratio of rf/rys-RT, as rys*1, rys*2, ryd*1 and ryd*2 lie on the steeply rising portion of the Materials Performance and Characterization

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T versus rys or T versus ryd curve. Then, there will be four B calibration curves generated: B1 based on rf/rys*1, B2 based on rf/rys*2, B3 based on rf/ryd*1 and B4 based on rf/ryd*2; application of B1 and B2 in Eq 6a generates two sets of MC KIC values and application of B3 and B4 in Eq 6b generates another two sets of MC KIC values. Thus, the present methodology ties the constant B indirectly to the Charpy transition curve (T versus CV)—DBTT-curve. APPLICATION OF THE NEW METHODOLOGY FOR ESTIMATING TQ

For a material with known IIT data but with no T0, T0 can be calculated by the following procedure: 1. Plot the load-temperature diagram (LTD) as in Fig. 2 and determine TD 2. Determine ryd at TD, using Eq 7 and rf using Eq 8 The dynamic yield stress of an impact three-point bend (TPB) specimen is given by:

(7)

ryd ¼ 2:99

PGY W BðW  aÞ2

where W ¼ B ¼ 10 mm and a ¼ 2 mm for a standard (full-size) CVN specimen, and PGY is in N. The micro-cleavage fracture stress, rf is given by [14]: (8)

rf ¼ 2:52ryd

where ryd is the value at TD. Many people have used different values for the multiplication factor (plastic stress concentration factor) on the RHS of Eq 8, which has values 2.18 or 2.52, depending on the selected yield criterion, Tresca or von Mises, respectively [15]. Some earlier studies even used a value of 2.57 [1,13]. While Chaouadi and Fabry [15] use an average value of 2.35, in this paper, the factor has been taken as 2.52 and all values of reported rf have been corrected accordingly. As such, many rf values given here will differ from those given in the source references. 3. Plot the T versus CV curve and determine (T41J – 24) C and (T41J – 50) C temperatures. In case of excessive scatter, use a lower bound (LB) curve determined by a fit to the lowest data at various temperatures. 4. Plot the T versus rys and T versus ryd (PGY at various temperatures is converted to ryd using Eq 7) and obtain rys*1, rys*2, ryd*1 and ryd*2 and the corresponding rf/rys*1, rf/rys*2, rf/ryd*1 and rf/ryd*2 values. 5. Plot the rf/rys and rf/ryd versus temperature curves. 6. By plugging-in the rf/rys*1, rf/rys*2, rf/ryd*1 and rf/ryd*2 values in the B1 versus rf/rys*1, B2 versus rf/rys*2, B3 versus rf/ryd*1 and B4 versus rf/ryd*2 calibration equations generated earlier, determine the B1, B2, B3, and B4 values corresponding to the rf/rys*1, rf/rys*2, rf/ryd*1, and rf/ryd*2 values for the particular steel. 7. For each B, for selected values from the rf/rys or rf/ryd curve, calculate KJC using Eq 6a or Eq 6b (using a spread-sheet program, this can be easily done for a column of rf/ryd values corresponding to various temperatures). Since calibration was done using MC data of calibration steels, the Materials Performance and Characterization

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generated KJC data are treated as MC curve data. Then select the KJC data in the 80–120 MPaHm and the corresponding temperatures and apply to the multi-temperature equation due to Wallin to obtain the corresponding TQ estimate. 0¼

i¼n X i¼1

(9)



di expf0:019ðTi  T0 Þ ½31  Kmin þ 77 expf0:019ðTi  T0 Þg

n X ðKJCi  Kmin Þ4 expf0:019ðTi  T0 Þg 5 i¼1 ½31  Kmin þ 77 expf0:019ðTi  T0 Þg

where: Kronecker di ¼ 1 for valid data and 0 for non-cleavage or censored data (in the present case, take di ¼ 1 always), Kmin ¼ 20 MPaHm, and Ti ¼ the test temperature (temperature corresponding to a particular KJC value with the corresponding rf/rys or rf/ryd ratio). 8. Because the new methodology is mainly based on rf, the micro-cleavage stress, T0 estimate, TQ, based on the new methodology will be designated TQcfs, to imply the cleavage fracture stress (CFS) method. Since there will be four estimates of TQcfs values corresponding to the four values, B1, B2, B3, and B4, they will be designated as TQcfs1, TQcfs2, TQcfs3, and TQcfs4, respectively. The criterion for selection of the final estimate, TQcfs, will be given later. Calculated values of TQcfs for the calibration steels and also for other steels will be compared with actual T0 or other estimates like TQIGC (where T0 is not available).

Material Data The 21 steels listed in Tables 1 and 2 along with source references (listed in brackets appropriately) were used for generating the calibration curves as described in the previous section. All the steels, except the five Said steels, have rf values determined from either IIT or 4-point bend tests (only for the Lambert steels). All the rf values have been adjusted as described after Eq 8. The static yield stress data and its variation with temperature for the 21 calibration steels are experimentally determined. For the five Said steels, rf values have been determined in the following two independent ways. Based on Chaouadi’s data, Sreenivasan [6] gave the following fit: (10)

TQBT ¼ 1:5TD þ 40

where TQBT is the estimate of T0 obtained from TD, with BT representing brittleness-transition (as TD is called the brittleness-transition temperature). Thus, putting the actual T0 in Eq 10, an estimate of TD can be obtained. The TD values listed in Table 2 for the five Said steels were so determined; hence the exact agreement between actual T0 and TQBT (see, Table 2). As mentioned in Ref. [6], although Eq 10 has a tendency to accuracy, due to various reasons, including the robustness of the TD measurement from experimental data (especially for steels exhibiting high scatter), the estimated TQBT can be highly non-conservative as is demonstrated by the values for the two steels, 16 MND5 and HT9, listed in Table 2. After estimating Materials Performance and Characterization

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TABLE 1 (Micro) CFS and other strength properties of calibration steels. Steel

rys-RT (MPa)

rf (MPa)

T41J ( C)

rys*1 (MPa)

rys*2 (MPa)

ryd*1 (MPa)

ryd*2 (MPa)

JAERI Steels [16] JRQ

488

1873

25

542

569

718

746

Steel-A Steel-B

469 462

2089 2089

42 61

536 560

568 607

757 802

813 852

2262 2488

66 72

620 648

683 719

968 934

1082 1037

SCK–CEN (Report R-4122) Steels [17] T91 E97

544 557

EM10

495

2310

96

650

748

937

1024

F82H

562

2293

65

633

697

888

1005

433 586

2065 2211

84 16

542 626

598 645

832 775

933 817

Lambert-Perlade Steels [18] BM CGHAZ-100s ICCGHAZ-100s

534

1755

3

529

537

713

751

CGHAZ-500s

481

1483

29

470

477

648

680

SAID Steels [19] Steel-1

591

2148

86

703

766

792

849

Steel-2 Steel-3

493 266

1856 1280

60.5 8.5

588 315

633 339

775 516

836 547

Steel-4

339

1297

61.5

411

454

481

543

Steel-5

387

1651

40

505

556

664

713

DuplxSS [20]

450

2290

83

676

732

904

944

HT9 [21] JSPS [15]

604 461

2381 1701

18.5 36

651 475

672 479

802 568

844 609

20MnMoNi55 [15]

430

2129

70

496

549

793

892

16MND5 [22]

491

2331

88

612

675

856

945

Other Steels

the TD, the dynamic yield stress at TD is determined and, then, applying Eq 8, the corresponding rf value is determined. It must be mentioned that for the five Said steels and also for the four Lambert steels in Tables 1 and 2, as IIT data are not available, dynamic yield stress variation with temperature has been determined using Eq 11 [23]. (11)

ryd ¼ rys-RT þ

666 500  190 ðT þ 273Þ  logð2  1010  tÞ

where: rys-RT ¼ the room-temperature static yield stress, T ¼ the test temperature in  C, and t ¼ the time in ms (usually 0.1 ms is the time to general yielding in IIT). As such, t can be taken as 0.1 for obtaining the ryd corresponding to standard impact tests at a strain rate of 103 [1]. Additionally, a scaling can be applied, if the actual ryd at room or other temperature is known. If, for example, the actual ryd at RT exceeds that from Eq 11, the absolute difference (of the RT values) can be added to Materials Performance and Characterization

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TABLE 2 TQcfs estimates for the calibration steels compared with other TQ estimates

Steel

T0 ( C) (ASTM E1921)

TD ( C)

TQ-BT ( C)

TQ-IGC ( C)

TQcfs1 ( C)

TQcfs2 ( C)

TQcfs3 ( C)

TQcfs4 ( C)

TQcfs ( C)

JAERI Steels JRQ

65

60

50.0

47.2

70

64

70

59

59

Steel-A

76

80

80.0

60.1

83

78

85

87

78

Steel-B

97

85

87.5

84.3

96

101

106

106

96

SCK-CEN (Report R-4122) Steels T91

118

110

125

87.4

100

109

106

116

100

E97 EM10

115 138

110 115

125 132.5

101.8 101.7

103 125

113 140

110 144

116 142

103 125

F82H

118

95

102.5

95.8

98

108

101

108

98 116

Lambert-Perlade Steels 132

–

–

103

116

124

123

128

CGHAZ-100s

45

–

–

45

81

52

65

66

52

ICCGHAZ-100s CGHAZ-500s

12 6

– –

– –

30 18

56 23

27 5

50 29

37 19

27 5

SAID Steels Steel-1

119

106

119

101

114

108

85

86

85

Steel-2

103

103

73

95

94

110

105

94

BM

95.3

Steel-3 Steel-4

48.5 83

59 82

48.5 83

38 92

31 92

33 92

48 96

34 105

31 92

Steel-5

78

78.7

78

77

73

78

90

82

73

90.3

Other Steels DuplxSS HT9

120 34.8

90

95

120

126

150

133

120

112

128

33

71

50

64

58

50

JSPS 20MnMoNi55

5 126

40 105

20 117.5

18 82

55 122

13 125

3 106

8 114

13 106

16MND5

95a

133

160

94

120

128

126

133

120

a

Values vary from 85 to 102 C; the mid-value is reported in the table.

those computed using Eq 11 at various temperatures to yield the scaled ryd values. If the actual ryd at RT is less than that from Eq 11, the absolute difference (of the RT values) must be subtracted from those computed using Eq 11 at various temperatures to yield the scaled ryd values. If the RT ryd (ryd-RT) is not known, a good way to estimate ryd-RT is to apply the empirical relation—Eq 12—to obtain the RT dy dynamic general yield load (Pgy-RT , as will be obtained from instrumented Charpy V-notch tests) from the easily available RT static yield stress (rys-RT). The empirical equation for estimating Pgy-RT due to Mathy and Greday [24] is as follows: (12)

dy

Pgy-RT ðNÞ ¼ 6300 þ 14:8rys-RT dy

where rys-RT is in MPa and ryd-RT is estimated from the Pgy-RT using Eq 7. Combining Eqs 7 and 12 and applying the full-size standard Charpy specimen dimensions results in the following direct relation: (13)

ryd-RT ¼ 294:33 þ 0:691rys-RT

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Equation 12 is based on the results of low and medium strength steels (mainly of the ferrite-pearlite type) and, hence, its applicability (as also of Eq 13) to higher strength steels and to irradiated steels is to be done with caution. However, it can be used, provided the results are consistent with each other as will be shown when discussing the results. Similarly, for estimation of the rys as a function of temperature, the following equation due to Server [25] can be used:

(14a)

(14b)

T ¼

   9 T þ 32 5

  rys ¼ 6:895 73:62  0:0603T  þ ð1:32  104 ÞðT  Þ2 ð1:16  107 ÞðT  Þ3

where T is in  C and rys is in MPa. Equation 14a actually converts  C to  F for use in Eq 14b. The above equation can also be scaled using known values at one or two temperatures as discussed after Eq 11 in the case of ryd estimation. As far as possible, test results are preferred. One method of estimating rf for the Said steels in Tables 1 and 2 was described above. The other method is based on the experimental tensile fracture stress results presented in Said et al. [19]. The fracture stress of a smooth tensile test specimen at zero-ductility temperature (designated Tsp in Ref. [19]) is related to rf. For the ratio rf/rf-tension (rf-tension is the fracture stress in a smooth specimen tensile test where tensile and yield stress coincide at Tsp and is designated Skop in Ref. [19]), Saario et al. [26] give a value of 1.66 while Said and Talas [27] give a value of 1.71. Taking the value as 1.71, rf can be computed using the rf-tension values given in Ref. [19]. The results are reported in Table 3 for the five Said steels listed in Table 1. They show excellent agreement within experimental error with those determined from ryd at TD (also listed in Table 3) and, hence, a mean value is taken as the rf value for the steels in Table 3 and the same has been reported in Table 2. Thus, the agreement between rf values from two independent methods lends confidence to our estimates. Another point to be mentioned about the Said steels (Tables 1 and 2) is that they range from low-strength ferrite-pearlite to medium to intermediate strength quenched and tempered steels, too. The Lambert steels in Tables 1 and 2 pertain to base metal (BM) and weld heat affected zones including coarse grained heat affected zone (CGHAZ) material in two cooling conditions and one intercritical CGHAZ material (ICCGHAZ) in one cooling condition simulated in a Gleeble type test machine. The other steels are mostly of the reactor pressure vessel (RPV) type ASTM A533B or similar steels or the 9Cr1Mo type and reduced activation martensitic (RAFM) steels and a Q&T 12Cr martensitic steel, HT9. The steels used for prediction based on the present CFS model, apart from the calibration steels in Tables 1 and 2, are listed in Table 4 with source references. The IGCAR steels consist of a quenched and tempered 9Cr-1Mo martensitic steel (91BM-IGC) [3], a normalized C-Mn steel A48P2 steel (A48P2-IGC) [28], a serviceexposed 2.25Cr-1Mo steel (21IGC) [2], a post-weld heat treated 9Cr-1Mo weld (91Wld-IGC) [29], a quenched and tempered ASTM A403 stainless steel (12Cr martensitic SS) (403SS-IGC) [30] and a normalized and tempered ASTM A203D 3.5 %Ni steel (A203D-IGC) [31]. The ASTM STP 870 steels are old RPV steels of Materials Performance and Characterization

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297

TABLE 3 Cleavage fracture stress for the five Said steels [19] of Table 1 by two independent methods.

Steel Steel-1 Steel-2

TD from T0 ( C)

ryd at TD (MPa)

rf from ryd at TD (MPa)

TSP ( C)

Skop (MPa)

rf from Skop ¼ 1.71 Skop (MPa)

Mean rf (MPa)

106 95.3

892 798

2248 2011

219 199

1198 995

2049 1702

2148 1856

Steel-3

59

572

1441

180

654

1118

1280

Steel-4 Steel-5

82 78.6

474 690

1194 1739

201 197

819 914

1401 1563

1297 1651

types similar to ASTM A302B or A212B and their welds and some old A533B steels also. The ASTM STP 1046 steels are modern RPV steels of the types ASTM A533B or A508 and their welds tested at BARC, India under an IAEA program. Among the other steels, HSST-02 is an ASTM A533, Grade B, Class 1 steel well-characterised under the Heavy Section Steel Technology (HSST) program in both unirradiated and irradiated (I) conditions [34,35] while the 403SS-DQT steel is an ASTM ASTM 403 SS (12Cr martensitic SS) [36,37] subjected to double quench and tempering heat treatment to improve toughness. All the Table 4 steels have IIT data available, but the static yield stress variation with temperature has been estimated mostly using the relations discussed before.

Results and Discussion CALIBRATION CURVES AND TQcfs ESTIMATIONS FOR THE CALIBRATION STEELS

The values of T41J, rys*1, rys*2, ryd*1, and ryd*2 for the calibration steels are listed in Table 1 along with rf. As discussed in the Present Methodology, the a and b values determined based on free fit of rf/rys and rf/ryd in Eq 4 are given in Table 5. as and bs are values obtained from the fit to the static ratio rf/rys and ad and bd are values obtained from the fit to the dynamic ratio rf/ryd, respectively. Table 5 also gives the average values of a and b for both the static and dynamic cases, i.e., aavs and bavs, aavd and bavd, respectively. Then, keeping the a values as constants at aavs (0.366) and aavd (1.858), respectively, a second fit was done as described in the Present Methodology section. The resulting b values (designated B) are shown in Figs. 3–6; Figs. 3 and 4 display the B values based on fit to Eq 6a against rf/rys*1 and rf/rys*2, respectively (Figs. 3 and 4 have the same B values as ordinates but abscissae are different), while Figs. 5 and 6 display the B values based on fit to Eq 6b against rf/ryd*1 and rf/ryd*2, respectively (Figs. 5 and 6 have the same B values as ordinates but abscissae are different). The B values based on best fit to the data in Figs. 3–6 are designated B1, B2, B3, and B4, respectively. It was found that a cubic fit gives the best fit in all cases. The resulting equations or calibration curves are given below.

(15a)

B1 ¼  11:0557 þ 11:8355

Materials Performance and Characterization

    rf rf 2 rf 3  3:5165 þ 0:3344 rys1 rys1 rys1

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298

TABLE 4 (Micro) CFS and other strength properties of test steels. T41J ( C)

rys-RT (MPa)

rf (MPa)

91BM-IGC [3]

512

2425

79

3.9113

3.6631

2.558

2.1032

A48P2-IGC [28] 21IGC [2]

556 280

2019 1613

76.6 20.5

2.7695a 4.7949a

2.5428a 4.3128a

2.4682 2.5767

2.141 2.3209

Steel

rf/rys*1 (MPa)

rf/rys*2 (MPa)

rf/ryd*1 (MPa)

rf/ryd*2 (MPa)

IGCAR Steels [6]

91Wld-IGC [29]

560

2140

8.5

3.5762

3.435

2.5659

2.4318

403SS-IGC [30] A203D-IGC [31]

656 390

2143 1367

34.5 57.5

3.2226 2.7672a

3.1376 2.5744a

2.7369 2.0252a

2.6424 1.8374a

M–Y–Wld M–Y–Wld–I

453 703

1907 1926

36 182

3.5922 2.9051

3.4133 2.8877

2.4799 2.7086

2.3241 2.6316

M–Y–TL–P

436

1735

29

3.4909

3.3111

2.4506

2.295

M–Y–TL–P–I EPRI–EP–24–Wld

558 350

1723 1890

77 29.5

3.1442 4.621a

3.0823 4.3349a

2.6549 2.9905

2.5451 2.8125

EPRI–EP–24–Wld–I

541

1943

72.5

3.5783

2.9938a

2.9a

ASTM STP 870 Steels [6,32]

3.673 a

a

EPRI–EP–23–Wld EPRI–EP–23–Wld–I

367 552

1824 1911

11 72

4.4706 3.5323

4.2125 3.4432

2.9372 2.922

2.7511 2.8103

A302BRCM–P

432

1695

12

3.5759

3.3968

2.5148

2.3673

A302BRCM–P–I

599

1801

58

3.0269

2.9573

2.5729

2.4841

AP AP-I

440 504

1778 1884

3.5 28

3.7669 3.6092

3.5992 3.575

2.6498 2.6761

2.3519 2.3788

FH

467

2343

88

3.9378

3.6898

2.6355

2.277

FH-I GW

559 478

2143 1895

78 60

3.1842 3.3246

3.0141 3.1426

2.3971 2.3927

2.1957 2.8931a

GW-I

568

1949

34

3.0887

2.953

2.351

2.1801

JH JH-I

459 595

2190 2285

80.5 63

3.7955 3.3116

3.561 3.1561

2.6165 2.5474

2.3224 2.3777

HSST Plate02 [34,35] HSST Plate02–I [34]

489 580

1801 2054

6 34

3.5876 3.4932

3.4971 3.395

2.6761 2.5934

2.4272 2.4136

403SS–DQT [36,37]

615

2550

25

3.8231

3.701

2.9651

2.6927

ASTM STP 1046 Steels [33]

Other Steels

a

The bold underlined values indicate values outside the permitted range.

with correlation coefficient R ¼ 0.9153 and validity range for (rf/rys*1) ¼ 3.05 to 4.3. MC KJC data from B1 is given by:   rf KJC ¼ 20 þ 0:366  exp B1 (15b) rys

(16a)

    rf rf 2 rf 3 B2 ¼  25:8537 þ 25:7755  7:8807 þ 0:7858 rys2 rys2 rys2

with correlation coefficient R ¼ 0.8525 and validity range for (rf/rys*2) ¼ 2.8 to 3.89. MC KJC data from B2 is given by: Materials Performance and Characterization

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299

TABLE 5 aavs and aavd estimates for the calibration steels. Steel/Statistics

as

bs

ad

bd

4.11  103

2.9453

4.4571

9.55  104

3

9.00  10 0.1556

2.3494 1.7351

2.4095 2.7429

0.1092 0.077

0.8643 0.9783

1.3824 1.2427

1.9 1.3052

1.8084 1.6686

EM10

1.29  1068

34.2963

0.922

1.9038

F82H

1.1233

1.3223

3.1518

1.4515

BM CGHAZ-100s

1.1427 3.10  105

1.2325 4.513

2.0854 7.93  103

1.6468 3.2716

ICCGHAZ-100s

3.00  1021

15.5422

2.47  103

4.1667

15

1.46  10

12.2952

3.44  103

4.4864

Steel-1

0.2736

1.9249

2.3646

0.3507

Steel-2 Steel-3

0.1775 0.4476

2.0533 1.4149

2.5768 2.9937

0.2379 0.0844

Steel-4

0.2687

1.8056

1.5252

1.2825

Steel-5

0.8321

1.4731

2.6979

0.129

3.47  105

3.9743

4.22  103

3.9731

27

18.0926 1.1781

3.1917 0.1183

5.51  103 2.252

JAERI Steels JRQ Steel-A Steel-B SCK–CEN (Report R-4122) Steels T91 E97

Lambert-Perlade Steels

CGHAZ-500s SAID Steels

Other Steels DuplxSS HT9 JSPS

8.46  10 0.9407

20MnMoNi55

0.1134

1.5854

2.3438

1.5211

16MND5

0.1134

1.5854

2.2129

0.115

STATISTICS aavs 5 0.3659

bavs ¼ 5.4297

aavd 5 1.8579

bavd ¼ 1.4544

Median SD

0.1775 0.423

1.8056 8.2441

2.2129 1.2755

1.4515 1.4717

Standard Error

0.0923

1.799

0.2783

0.3212

95 % Conf 99 % Conf

0.1925 0.2626

3.7527 5.1192

0.5806 0.792

0.6699 0.9139

Mean

(16b)

(17a)

KJC

  rf ¼ 20 þ 0:366  exp B2 rys

    rf rf 2 rf 3 B3 ¼ 2:1566 þ 1:1167  0:8441 þ 0:1263 ryd1 ryd1 ryd1

with correlation coefficient R ¼ 0.8695 and validity range for (rf/ryd*1) ¼ 2.28 to 2.995. MC KJC data from B3 is given by: Materials Performance and Characterization

SREENIVASAN ON CVN SPECIMENS WITHOUT PRECRACKING

FIG. 3 B values obtained from fit to Eq 6a—KJC ¼ 20 þ aavsexp (B*rf/rys) with aavs ¼ 0.366 plotted against rf/rys*1 for each steel in Table 1.

FIG. 4 B values obtained from fit to Eq 6a—KJC ¼ 20 þ aavsexp (B*rf/rys) with aavs ¼ 0.366 plotted against rf/rys*2 for each steel in Table 1.

FIG. 5 B values obtained from fit to Eq 6b—KJC ¼ 20 þ aavdexp (B*rf/ryd) with aavd ¼ 1.858 plotted against rf/ryd*1 for each steel in Table 1.

Materials Performance and Characterization

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FIG. 6 B values obtained from Eq 6b— KJC ¼ 20 þ aavdexp (B*rf/ryd) with aavd ¼ 1.858 plotted against rf/ryd*2 for each steel in Table 1.

(17b)

(18a)

  rf KJC ¼ 20 þ 1:858  exp B3 ryd     rf rf 2 rf 3 B4 ¼ 33:7517  37:5066 þ 14:713  1:9472 ryd2 ryd2 ryd2

with correlation coefficient R ¼ 0.8334 and validity range for (rf/ryd*2) ¼ 2.09 to 2.82. MC KJC data from B4 is given by:   rf (18b) KJC ¼ 20 þ 1:858  exp B4 ryd Estmates of T0, based on MC KJC data from Eqs 15b, 16b, 17b, and 18b are designated TQcfs1, TQcfs2, TQcfs3, and TQcfs4, respectively. Following the procedures described above, the estimated TQcfs values for the calibration steels are tabulated in Table 2 along with actual T0 and other estimates like TQIGC and TQBT. The four TQcfs values show excellent agreement with each other within the error that can be expected for this. The criterion for choosing the estimate based on the present CFS model, namely, TQcfs, is the most conservative of the four: TQcfs1, TQcfs2, TQcfs3, and TQcfs4. The TQcfs values so determined are given in the last column of Table 2. Figure 7 compares the T0 of calibration steels with the TQcfs estimates tabulated in Table 2. TQcfs estimates are conservative for most of the calibration steels. This becomes more apparent in Fig. 8, which plots the residuals (i.e., (T0 – TQcfs) values) for the 21 calibration steels; only for two steels, the values lie outside the 620 C band, one being conservative (Steel 1) and the other being non-conservative (16MND5 steel). For the 16MND5 steel, there is large scatter in the basic T0 data (see footnote to Table 2) and other estimate TQBT also shows this tendency for nonconservatism. It must be noted that for the Steel 1 (also, for all the Said Steels in Tables 1 and 2), as discussed before, the dynamic yield stress has been estimated using Eqs 11 to 13. There may be some error in this leading to large over conservatism as TQcfs for the Steel 1 is determined by TQcfs3, the value based on ryd and ryd*1. Materials Performance and Characterization

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FIG. 7 ASTM E1921 reference temperature (T0) of calibration steels compared with TQcfs estimated using the CFS model.

Otherwise, based on the results for all other steels, the premise for choosing the value of TQcfs, namely, the most conservative of the four estimates TQcfs1, TQcfs2, TQcfs3, and TQcfs4, seems to be sound. Figure 9 compares the TQcfs estimates for the calibration steels with their TQIGC estimates. TQcfs is less conservative than TQIGC and, hence, closer to the actual T0. This is to be expected from Fig. 8, which shows 620 C error band for TQcfs whereas TQIGC is mostly conservative to the extent of 20 C–30 C (see the Introduction).

TQcfs ESTIMATIONS FOR THE TEST STEELS

For all the IGCAR steels in Table 6, except for the A48P2-IGC, static yield stress estimates have been made by scaling yield stress data given in literature for similar steels, instead of Eq 14. For example, for 9Cr-1Mo steels, the data given by Chaouadi

FIG. 8 Residuals for the calibration steels based on TQcfs estimated using the CFS model.

Materials Performance and Characterization

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FIG. 9 TQIGC compared with TQcfs for the calibration and test steels.

in Ref. [17] has been used with suitable scaling. This makes the static yield stress estimates closer to actual values. Estimates of TQcfs obtained for the test steels are tabulated in Table 6 while the strength properties are given in Table 4. In Table 6, the ratios rf/rys*1, rf/rys*2, rf/ryd*1, and rf/ryd*2 have been given and the values that fall outside the range specified by Eqs 15a–18a have been highlighted in bold fonts with underlining. Comparison of highlighted values in Table 4 with the corresponding TQcfsx (where x stands for 1, 2, 3, or 4, as the case may be) in Table 6 indicates that when the rf/rys*x or rf/ryd*x value falls outside the specified range, the corresponding TQcfsx prediction is unacceptably large or small and such values have not been considered for evaluation. In such cases, the largest of the valid values has been reported as TQcfs in the last column of Table 6. Such behavior occurs mostly for steels with low-strength, say, rys-RT < 400 MPa. For the low-strength IGCAR steel, A203D-IGC, none of the TQcfsx values are acceptable (vide Tables 4 and 6). Hence, the present CFS model based on the calibration curves derived in this paper seems to be applicable to steels with rys-RT in the range of 400–700–750 MPa. The requirement for strict concurrence of the rf/rys*x or rf/ryd*x values to the validity range arises from the fact that the calibration curves are cubic equations which will behave wildly outside their range of fit. For the GW steel, though rf/ryd*2 has been shown outside the validity range in Table 4, the corresponding prediction in Table 6, namely, TQcfs4 does not show much difference from other values. This may be due the fact that rf/ryd*2 value for the GW steel exceeds the upper limit only marginally and hence does not have any effect on B4 value. For the HSST-02 plate, Ref. [35] gives a T0 of 28 C which excellently compares with the TQcfs value of 18 C reported in Table 6. For the M–Y Wld, Ref. [38] gives a T0 of 106 C in the unirradiated condition and a T0 of 106 C in the irradiated condition (fluence: 6.11  1011 n cm2; E > 1 MeV) which also compare excellently with the TQcfs value of 76 C and 118 C (fluence: 1.3  1011 n cm2; E > 1 MeV) reported in Table 6 for the two conditions, respectively. However, TQIGC values are very conservative. Finally, Fig. 9 also shows the TQIGC values for the Materials Performance and Characterization

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304

TABLE 6 TQcfs estimates for the test steels compared with other TQ estimates. Steel

TQIGC ( C)

TD ( C)

TQ-BT ( C)

TQcfs1 ( C)

TQcfs2 ( C)

TQcfs3 ( C)

TQcfs4 ( C)

TQcfs ( C)

112

110

122

157

110

277a 2322a

232a 2640a

91 60

69 63

69 60

IGCAR Steels 91BM-IGC

86.6

105

A48P2-IGC 21IGC

104.8 56.6

69 49

118

91Wld-IGC

26.9

34

11

36

26

32

25

25

403SS-IGC A203D-IGC

22.3 70

39 86

18.5 89

20 257a

8 215a

9 296a

4 2142a

4 ––

M-Y–Wld M–Y–Wld–I

52.5 150

73 71

69.5 146.5

81 113

76 113

86 100

78 118

76 118

M–Y–TL–P

52.4

53

39.5

69

65

73

65

65

M–Y–TL–P–I EPRI–EP–24–Wld

77.8 66

10 109

55.0 123.5

28 2145a

38 211a

23 56

33 56

38 56

70.8

49

33.5

39

20

52a

233a

39

a

a

53 9

58 24

53 34

53

51

64 33.5

ASTM STP 870 Steels

EPRI–EP–24–Wld–I EPRI–EP–23–Wld EPRI–EP–23–Wld–I

39.6 72

93 39.5

99.5 19.3

298 5

2138 34

A302BRCM–P

37.3

34

11.0

57

51

61

17

65.5

34

16

A302BRCM–P–I

67

6.4

2.4

34

ASTM STP 1046 Steels AP AP–I

33 24

FH

95

FH-I GW

89 91

33.3 4 120 82.9 64.9

GW-I

58

32

JH JH-I

92 64

112.8 91.8

HSST-02 HSST-02-I

25 37

33.5 2

403SS-DQT

32.3

69

9 34

46 32

33 6

35 2

42 6

33 6

140

127

125

120

132

120

84 57

119 86

115 84

124 89

123 285a

115 84

8

46

40

63

60

40

129 97

119 111

117 106

117 115

123 108

117 106

10.3 37

36 27

18 5

31 11

36 8

18 5

63.5

45

33

46

60

33

Other Steels

a

Values highlighted and underlined are invalid.

Table 6 test steels plotted against the respective TQcfs values which are in excellent accord with the trend shown by the calibration steels displayed in the same figure and discussed in the previous section. Figure 9 also suggests the conclusion that TQIGC values for steels with high reference temperature (like the highly irradiated steels) have a tendency to be much more conservative than the corresponding TQcfs values as compared to steels having lower reference tempaeratures (i. e., steels near the left axis of Fig. 9). The implication is that, in such cases, TQcfs values may even be unacceptably non-conservative. However, the data for HSST-02 and M–Y Wld do not warrant such a conclusion. Nevertheless, the present data are not sufficient to draw a definite conclusion. This aspect needs further verification.

Materials Performance and Characterization

SREENIVASAN ON CVN SPECIMENS WITHOUT PRECRACKING

Conclusions Thus, a semi-empirical formulation of the CFS (cleavage fracture stress) model, based on the microscopic cleavage fracture stress, rf, for estimating the ASTM E1921 reference temperature of ferritic steels from instrumented impact test (IIT) of unprecracked CVN specimens has been established. The relevant calibration equations necessary for applying the model have been derived and demonstrated for steels with room temperature yield strength in the range 400–750 MPa, including irradiated steels. However, applicability and acceptability of the present calibration curves for highly irradiated steels need further examination. The estimate of T0, based on the present CFS model, TQcfs, lies within a 620 C band, being conservative for most of the steels, but less conservative than TQIGC based on the IGC-procedure. Moreover, the CFS model is a single step assessment procedure as compared to the multi-stage IGCAR-procedure and, hence, less error-prone due to calculation errors. However, the parameters must strictly meet the validity conditions for the calibration equations. CFS model enhances the validity and utility of the CVN IIT by enabling estimation of design-relevant master curve from unprecracked CVN specimens. Although some researchers have called an approach based on the CFS mesoscopic (i.e., lying between microscopic and macroscopic scales), at least in steels, rf operates over microstructural distances. As such, the CFS model or approach should be the preferred Nomenclature. ACKNOWLEDGMENTS

The writer acknowledges with thanks the excellent support and encouragement received from Director, MMG and Director, IGCAR, Kalpakkam, India.

References [1] Sreenivasan, P. R., “Instrumented Impact Testing—Accuracy, Reliability, and Predictability of Data,” Trans. Indian Inst. Metals, Vol. 49, No. 5, 1996, pp. 677–696. [2] Sreenivasan, P. R., Shastry, C. G., Mathew, M. D., Bhanu Sankara Rao, K., Mannan, S. L., and Bandyopadhyay, G., “Dynamic Fracture Toughness and Charpy Transition Properties of a Service Exposed 2.25Cr–1Mo Reheater Header Pipe,” Trans. ASME J. Eng. Mater. Technol. Vol. 125, No. 2, 2003, pp. 221–233. [3] Moitra, A., Sreenivasan, P. R., Mannan, S. L., and Singh, V., “Ductile–Brittle Transition Temperatures and Dynamic Fracture Toughness of 9Cr–1Mo Steel,” Metall. Mater. Trans., Vol. 36, No. 11, 2005, pp. 2957–2965. [4] Server, W. L., “Impact Three–Point Bend Testing for Notched and Precracked Specimens,” J. Test. Eval., Vol. 6, No. 1, 1978, pp. 29–34. [5] ASTM E1921-10: Standard Test Method for Determination of Reference Temperature, T0, for Ferritic Steels in the Transition Range, Annual Book of ASTM Standards, ASTM International, West Conshohocken, PA, 2010. [6] Sreenivasan, P. R., “Inverse of Wallin’s Relation for the Effect of Strain Rate on the ASTM E1921 Reference Temperature and its Application to Reference Temperature Estimation From Charpy Tests,” Nucl. Eng. Des., Vol. 241, No. 1, 2011, pp. 67–81. Materials Performance and Characterization

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