experimental study of effects of prior overload on fracture toughness of a533b steel

$: Experimental study of effects of prior overload on fracture toughness of A533B steel$
Int. J. Pres. Ves. & Piping 41 (1990) 297-331

Experimental Study of Effects of Prior Overload on Fracture Toughness of A533B Steel

D. J. Smith* & S. J. Garwood~

Engineering Department, The Welding Institute, Abington Hall, Abington, Cambridge CB1 6AL, UK

(Received 29 August 1989; accepted 8 September 1989)

A BSTRA CT

This paper presents the results of an extensive study carried out to examine the effects of prior overloading over the entire fracture transition regime for 50-ram thick A533B steel. The main variables examined are temperature, crack orientation with respect to the rolling direction, level of prior overload, the initial crack length, and the statistical variation of prior overload effects. It is found that the effect of prior overload on fracture toughness at lower temperatures is dependent on orientation, so that in the L-T orientation for short and medium cracks (0"2 and 0"5 a/W) there is a benefit throughout the transition regime of 50-ram thick A533B steel In the T-L orientation no benefit is obtained for temperatures greater than the initiation of tearing temperatures. Above these temperatures the prior overload sequence lowers the.fracture toughness. For L - T orientation long cracks (a /W= 0.7) it is found for temperatures lower than -140°C that prior overload apparently increases the toughness. At higher temperatures there is a loss of toughness even though failure is cleavage dominated up to -80°C.

* Present address: Department of Mechanical Engineering, University of Bristol, Bristol BS8 1TR, UK. :~ To whom correspondence should be addressed.

297

Int. J. Pres. Ves. & Piping 0308-0161/90/$03"50 © 1990 Elsevier Science Publishers Ltd, England. Printed in Great Britain

298 D. J. Smith, S. J. Garwood

On the lower shelf at - 170 ° C in the L - T orientation the fracture toughness variability after preloading is found (based on a sample of 14 specimens) to exhibit a birnodal distribution. This distribution is similar to that exhibited by non-preloaded material.

1 INTRODUCTION

Since the review of Nichols, 1 a considerable body of experimental and theoretical research has been conducted to quantify the advantages of overstressing in producing crack tip residual stresses. Research has also been directed to a special case of overstressing called warm prestressing and its relationship to the performance ofpressurised water reactor (PWR) pressure vessels during fault conditions. 2 In general, it is known that, when a cracked structure is loaded at a temperature at which local plastic deformation occurs, the subsequent load-bearing capacity at lower temperature is greater than that expected without the initial loading. The temperature at which the overload takes place is above the ductile-brittle transition of the fracture behaviour of the material in question. The structure is partially or completely unloaded. Experimentally, this process is found to produce an apparent increase in the material's resistance to fracture. In a recent review of warm prestressing studies, Pickles and Cowan 3 identified two principal factors which cause the apparent increase in fracture toughness: an increase in yield strength at the temperature at which subsequent loading takes place, and the formation of crack tip compressive plastic zones during unloading. Pickles and Cowan 3 pointed out that when other factors, such as irradiation, affect the yield stress, the benefit of overstressing can be lost. The experimental evidence and the theories that account for the warm prestressing effect were reviewed in detail by Pickles and Cowan 3 and Smith and Garwood (this vol., pp. 255-96). s

Much of the experimental work conducted to examine the effects of warm prestressing effects has been carried out at temperatures at which cleavage is the dominant failure mechanism. However, pressure vessels are not, or are presumed not to be, operated at temperatures where cleavage is likely to occur. Materials selection is likely to be based on fracture properties which are in the brittle to ductile transition or upper shelf regimes. Little attention has been paid to examining the subsequent fracture performance in the transition regime after warm prestressing or proof loading. Of the limited work conducted in this area, Sutcliffe e t aL 4 showed that a benefit is obtained for a pressure vessel steel throughout its transition regime. There are published experimental data and theoretical models, reviewed by Smith and Garwood, 5 which found that proof testing improved performance in the

Prior overload effects on fracture toughness: Experimental study 299

cleavage fracture regime. However, the principal objective of the present experimental work is to examine the range of applicability of the prior overloading technique throughout the transition from brittle to fully ductile behaviour in a pressure vessel steel.

2 MATERIAL AND BASIC CHARACTERISATION

2.1 Material

The material selected for the experiments was A533B Class 1 steel. This was supplied by Creusot-Loire as a 50-mm thick quenched and tempered plate 7.5 m long and 2 m wide. The plate was subjected to a chemical analysis using direct-reading optical emission spectrography on a remelted sample. The results of the analysis are given in Table l(a). Nitrogen and oxygen contents were measured by an inert gas fusion method, the results of which are given in Table l(b). Microstructural examination revealed that the structure was basically tempered bainite. There was evidence of banding of the microstructure at mid-thickness of the plate, typical of rolled C-M steels. The ASTM grain size was between 7 and 8.

TABLE 1 (a) Chemical Analyses of A533B Class 1 Steel

C S P Si Mn Ni Cr Mo V

0"18 0'005 0"006 0"24 1-41 0"56 0"18 0"48 <0"002

Cu Nb Ti AI B Sn Co CE a

0"12 <0'002 <0.002 0"018 <0'0003 0-01 0"01 0'59

C Mn (N i+Cu) a C E = IIW carbon equivalent formula = + 6 - + 15

The Welding Institute analysis Ref. No. is S/84/27.

(Cr + Mo + V)

5

(b) Nitrogen and Oxygen Analysis of A533B Class 1 Steel

Element (ppm by weight)

N 2 70, 74, 67, 74

0 2 13, 20, 10, 8


Fig. 1. Conventional terminology for fracture and tensile specimen orientations.

2.2 Tensile

A series of Hounsfield No. 14 tensile specimens to BS 18 were tested at temperatures ranging from -196 to 20°C. These samples were extracted from the mid-thickness parallel and transverse to the rolling direction of the plate, as shown in Fig. 1. All the tests were carried out at a crosshead displacement rate of 2mm/min. The 0.2% proof strength and tensile strength are shown in Fig. 2(a) and (b) respectively.

2.3 Impact

The nil-ductility transition (NDT) temperature was determined by the Pellini drop weight test. 6 Specimens measuring 50 mm wide by 130 mm long and 16mm thick were extracted 3mm from the surface of the plate. The specimen orientation was transverse to the rolling direction. The drop weight tests were conducted using an impact energy of 380J. The test results showed that the NDT temperature was -50°C.

Standard Charpy V-notch specimens were extracted from the T-L and L-T orientations (see Fig. 1) at mid-thickness, and were tested according to


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Tensile properties for A533B in directions parallel and transverse to the rolling direction: (a) yield stress; (b) tensile stress.

the ASTM E23-82 standard. 7 The absorbed energy and percentage crystallinity as a function of temperature are shown in Fig. 3(a) and (b) for the T-L and L-T orientations respectively.

3 F R A C T U R E TOUGHNESS

3.1 Selected variables

The programme of work consisted of examining the influence of a number of

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Charpy transit ion curves in A533B: (a) T - L orientation; (b) L - T orientation.

302 D.J. Smith, S. J. Garwood

variables on the fracture behaviour of A533B steel, with and without the application of a prior overload. The variables were:

(1) temperature, (2) crack orientation with respect to the rolling direction, (3) level of prior overload, (4) the initial crack length, and (5) statistical variation of prior overload effects.

Fracture toughness data were obtained for temperatures ranging from the lower shelf ( - 196°C) to the upper shelf (20°C) and for crack orientations transverse and parallel to the rolling direction of the plate material. Two levels of prior overload were applied: one where local yielding at the crack tip had just occurred, and the other where significant yielding ahead of the crack took place. The type of loading cycle used in the study is described below. Three initial crack lengths were examined, and a preliminary investigation was carried out to provide information on the statistical variation of the prior overload effect at - 170°C.

3.2 Specimens

A total of 172 three-point single-edge bend specimens (SENB) were extracted from the plate, 118 specimens with their axes parallel (L-T) and 54 specimens transverse (T-L) to the rolling direction, as shown in Fig. 1. The specimens were notched perpendicular to the plate surfaces to notch depths suitable for crack lengths equivalent to a/W of 0.2, 0.5 and 0.7. Although the a~ W values of 0.2 and 0-7 are non-standard, in other respects the specimens complied with the geometry requirements of the 'preferred' testpiece to BS 57628 (i.e. B=fu l l th ickness=50mm and W = 2 B = 1 0 0 m m ) . In all specimens the notches were extended by fatigue cracking to the appropriate crack length, using the recommended procedures, s

3.3 Tests without prior overload

Tests were conducted over a range of temperatures from -196 to 20°C, following the procedure recommended for crack tip opening displacement (CTOD) testing. 8 Each specimen was instrumented with a clip gauge mounted across the notch mouth. Loading to fracture was carried out in a 1.8MN universal testing machine, again following recommended procedures, a Tests were conducted on specimens with initial crack lengths a/W equal to 0.2, 0.5 and 0.7 in the L-T orientation. For the T-L orientation, tests were conducted at a/W= 0.5.


To examine the variability of the fracture toughness on the lower shelf, a further 13 tests were conducted at -170°C on L-T orientation specimens with a / W = 0.5.

On completion of the tests all the specimens were broken open where necessary and both the initial fatigue crack length (ao) and the extent of the crack growth (Aa) were measured.

From each test a number of parameters were evaluated, namely the CTOD, J-integral and a reference stress (oR) or an equivalent net section stress. The clip gauge opening displacement at fracture with and without ductile tearing or at the attainment of maximum load was used to calculate

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orientation.


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K 2 0 " 4 ( W - ao)V p (1) 60 = 6~, 6u, 6m -- E'2av + 0"4W+ 0"6ao + z

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K 2 2 U v W Jo = ~ - r + 0 .4W+ 0"6ao + z [ B ( W - ao)] (2)

The reference stress was evaluated following the techniques used in creep analysis, x° An upper bound reference stress is given by

aR <- (P/PL)aV (3)

where the limit load PL for a beam containing a crack is given by

PL = fl av (W-- a°)2B 2 W (4)

fl is termed a constraint factor. For deep notched specimens using a Von Mises yield criterion, slip line theory gives fl = 0.727 for plane conditions and fl = 0.536 for plane stress. For convenience the deep notch plane strain

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Fig. 5. Distribution of results for non-preloaded 0.5a/W L-T orientation A533B SENB specimens at -170°C in terms of: (a) CTOD; (b) K~c.


value of fl is used throughout and the reference stress at failure or maximum load using eqns (3) and (4) is given by

2.75P (re = B[1 -- (ao/W)] W 2 (5)

The fracture toughness test results are given in Appendix A, together with later fracture toughness results. CTOD results are shown in Fig. 4(a)-(d) as a function of temperature for the various initial crack lengths and specimen orientations. The distributions of the CTOD and K~c results from the fracture toughness tests conducted at - 170°C are shown in Fig. 5(a) and (b).

3.4 Tests with prior overload

To examine the effects of prior overload, tests were conducted using the same SENB specimen geometries and orientations as for the tests without prior overload. The test cycle consisted of loading each specimen to a fixed displacement and unloading to zero load at room temperature (20°C). This was followed by cooling the specimen to the temperature of interest and subsequently loading it to fracture. This is commonly termed the load- unload-cool-fracture (LUCF) cycle. The displacement to which selected specimens were initially subjected was, in the case of a/W= 0"5, sufficient to promote local yielding at the crack tip. The total crack mouth opening displacement, V=, was chosen to be 0.5 mm. A further series of tests were conducted on specimens with a/W= 0"5 which were subjected to an initial displacement, V=, of 1.8 mm.

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Tempera lure, °C

Fig. 6. CTOD transition curve for pro- loaded (V==0.2mm) at 20°C 0"2a/W L-T

orientation A533B SENB specimens.

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Temperafure, °C

Fig. 7. CTOD transition curve for preloaded (V==0-Smm) at 20°C O.5a/W L-T


306 Ol J. Smith, S. J. Garwood

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100 -

~161 -

I

A 6m

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-200 -80 40

Temp era ture, °C

Fig. 11. CTOD transition curve for preloaded (Vg= 1-8mm) at 20°C 0"5a/W T-L


In the case of specimens with short (a/W=0.2) and long (a/W=0.7) cracks, the prior overload displacement was determined by making the mouth opening angle the same as that for a/W=0.5, that is V~/(a/W)= 1"00 mm.

To examine the variability of the fracture toughness on the lower shelf after a prior overload of Vg -- 0"5 mm, a further 13 tests were conducted on specimens with crack depths of a/W= 0.5, following the loading procedure described above. As in the earlier fracture toughness tests, the specimens


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Distribution of results for preloaded (~ = 0"5 mm) at 20°C 0-5 a~ W L-T orientation A533B SENB specimens in terms of: (a) CTOD; (b)K~¢.

were broken open and the crack lengths were measured on completion of testing.

For each test the parameters CTOD, J and aR, as defined by eqns (1), (2) and (5), were evaluated for the prior overload conditions at room temperature, and for the subsequent fracture conditions at low temperature. The test results are tabulated in Appendix A. The CTOD results are shown in Figs 6-11 for the various conditions. The distribution of results at -170°C after prior overload at room temperature are shown in Fig. 12.

4 OBSERVATIONS AND DISCUSSION

4.1 Basic material data

4.1.1 Tensile and impact properties In terms of the yield and tensile properties, there appears to be no significant difference between the results for specimens extracted parallel to and transverse to the rolling direction. However, higher failure elongations are exhibited in the parallel than in the transverse orientation. In Appendix B a comparison is made between the tensile data obtained in this study and an earlier study. With regard to the impact properties, the L-T orientation exhibits higher Charpy values than the T -L orientation at the nil-ductility transition (NDT) temperature of -50°C.

3 0 8 D. J. Smith, S. J. Garwood

4.1.2 As-received fracture properties In Appendix B the condition for valid Ktc fracture toughness has been examined. For T-L and L-T orientations and crack length-to-width ratios equal to 0"2, 0"5 and 0"7, valid Ktc fracture toughness was obtained only for temperatures below - 140°C.

On the lower shelf, at - 170°C, the fracture toughness variability, based on a sample of 14 specimens, exhibited a bimodal distribution (see Fig. 5(a) and (b)). This remains unexplained. The results of fracture toughness tests carried out throughout the range of temperatures from - 1 9 6 to + 20°C are shown in Fig. 4(a)-(d). These figures indicate the temperatures above which some ductile tearing preceded final failure. For specimens with crack lengths a/W=0"5 in the L-T orientation this was about -60°C, and for similar specimens in the T-L orientation it was about -80°C. The initiation of tearing as well as the maximum load toughness was found to be dependent on the crack length for the L-T orientation. At temperatures close to and below the initiation of tearing temperature, the transition curves for each crack length were found to be practically independent of crack length. However, there was a tendency for higher CTOD values to be obtained in the transition regime for short and long cracks.

T A B L E 2 (a) Polynomial Coefficients b i and Temperature Limits to C T O D 60 Transition Curves for

Non-preloaded SENB Specimens

Orientation Coefficients Temperature and limits (°C)

('rack

length, a /W h 3 b 2 b 1 b o T L T v

L T0"7 - 0 " 5 1 6 4 5 × 1 0 - " - 0 ' 1 7 8 7 3 x 1 0 -3

L - T 0 " 5 - 0 " 4 6 6 3 5 × 1 0 ° - 0 ' 1 4 1 6 1 x 1 0 3

T - L 0 ' 5 - 0 " 3 7 0 5 5 x 1 0 6 - 0 " 1 4 2 7 5 x 1 0 3

L - T 0 " 2 - 0 " 4 2 9 4 8 × 10 -~ - 0 " 1 3 2 7 3 × 10 -3

- 0 ' 1 2 4 62 × 10 -2 0"165 32 - 1 9 6 - 4

- 0 ' 4 9 5 69 × 10 -2 0"312 21 - 196 16

- 0 " 2 3 6 70 × 10 -2 - -0"252 84 - 1 9 6 - - 1 0

0"571 25 × 10 -2 0"495 81 - 1 9 6 20

(b) Polynomial Coefficients b~ and Temperature Limits to Jo Transition Curves for Non-preloaded SENB Specimens

Orientation Coefficients Temperature and crack limits (°C)

length, a~ W

b3 b2 bl bo TL Tu

L T0"7 - 0 " 4 3 8 74 × 10 -6 - 0 " 1 0 5 9 0 × 10 - 3 0"39651 × 10 -2 0 ' 1 6 4 4 6 178 17

L - T 0 ' 5 - 0 " 4 1 5 8 1 × 1 0 -6 - 0 ' 1 1 5 5 5 × 1 0 3 0 . 6 8 2 7 1 × 1 0 - 2 0"32262 - 1 9 6 16

T - L 0 " 5 - 0 " 3 2 1 38 × 10 -6 - 0 " 1 0 6 8 9 × 10 -3 0"19341 × 10 -2 0"11270 - 196 - 10 L - T 0 " 2 - 0 " 7 1 3 2 0 × 1 0 - 6 - 0 " 1 7 6 4 9 X 1 0 - 3 0 ' 6 4 8 4 9 × 10 -2 0"59066 - 8 1 17


In Appendix B a comparison has been made between the fracture properties obtained in this study and other A533B data.

To compare the effect of initial crack length and orientation on the fracture toughness, for as-received and prior loading conditions (see below), curves were fitted to the experimental data. It was found that to obtain the best fit to the range of experimental data a least-squares third-order polynomial function was the most appropriate, where the relationship between the CTOD (or J) and temperature is given by

3

,6, \ g . ~ /

i = l

The solid lines in Fig. 4(a)-(d) are the fitted curves using eqn (6). The coefficients for each fit to the data given in terms of CTOD and J are given in Tables 2(a) and 2(b) respectively, together with the temperature range, T L to T U, for which eqn (6) is valid. Curves were not always fitted to the reference stress data using eqn (6), as the goodness of fit was sometimes unacceptable.

4.2 Fracture properties after prior overload

4.2.1 The extent of prior overload The amount of prior overload was intended to cause a certain level of plasticity ahead of the crack tip, as described above. The controlled variable was the crack mouth opening displacement, Vg. Because the initial fatigue crack lengths measured on the fracture surface after test were not constant

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there was an inevitable variation in the applied load for a given value of V~. This variation in terms of the prior overload reference stress is shown in Fig. 13. At a particular initial crack length the variation is _+6% about a mean reference stress. However, by retaining the same crack mouth opening angle for each initial crack length the reference stress is, as shown in Fig. 13, lower for a/W=0.2 than that for a/W=0.5, whereas a slightly higher prior overload reference stress is obtained at the longer crack length of a/W= 0-7. At the higher mouth opening displacement of V~ = 1-8 mm for specimens with a/W= 0"5, the prior overload reference stress is approximately at the yield stress level, as shown in Fig. 13.

The variability in the prior overload J is of the same order as that for the reference stress. Although the level of crack mouth opening displacement was kept constant for any crack length, the amount of plastic displacement Vp varied, and thus a wide range of values of prior overload CTOD was obtained. For all crack lengths and specimen orientations this variation was _+30%.

The extent of crack growth for the various prior overload sequences has been obtained from the measurements of crack length after the cleavage fracture which occurred during the final loading. For low levels of prior overload there was no indication of crack growth during overloading. At the higher overload level (Vg=l.8mm), applied to 0.5a/W specimens, the average value of crack growth in the L-T orientations was about 0.15 mm (see Appendix A, Table A2(iii), for temperatures below - 120°C), whereas in the T -L direction the crack growth was about 0-5 mm (see Appendix A) for temperatures below -120°C. Thus, for each orientation, the applied prior overload gave a reference stress close to or greater than yield and was sufficiently high to promote stable tearing.

4.2.2 The influence of orientation The fracture toughness transition curves for specimens which had been preloaded were derived by fitting the polynomial function given by eqn (6) to the data.

The coefficients for each fit to the data are given in terms of CTOD and J (6r and J0 in Table 3 and the curves for each data set are shown in Figs 6-11. In this and the following sections the influence of the variables investigated will be compared. For clarity this will be done using the curves given in eqn (6) rather than the individual data points.

For the initial crack length, a/W, of 0.5 the transition curves after preloading to V~ = 0.5 mm are shown in Fig. 14(a). On the lower shelf the effect of preloading is approximately the same for each orientation, but as the temperature increases the T - L transition curve drops below the L-T transition curve. This behaviour is similar to that observed for the transition

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curves (Fig. 4(a)-(d)) without preloading. F rom the measured crack lengths the initiation CTODs after preloading to Vg=0-5mm and without preloading appear to be similar.

After a higher preload level of Vg = 1.8 mm, the transition curves for each orientation, shown in Fig. 14(b), deviate further from each other as the temperature is increased, with the prior overload transition curve for the T - L orientation exhibiting lower CTOD levels.

4.3 Comparison of fracture properties with and without prior overload

4.3.1 Variability at - 1 7 0 ° C The log normal distributions in terms of CTOD (60, 60 and Kj (Kit, Kf) are shown in Fig. 15(a) and (b) respectively for the non-preloaded and preloaded conditions. It can be seen that the trends for both parameters are similar. Furthermore, the two fracture toughness distributions exhibit, within each set, two distinct linear portions. The steepest distribution curves for both data sets have approximately the same slope, i.e. they have the same scatter (and standard deviations). On the one hand, the steepest distribution curve for the non-preloaded material lies above the 70% confidence limit, and on the other hand, the steepest curve for the preloaded material commences at about 30% confidence level. The shallower curves for the two data sets do not, however, have the same slope. On the basis of the data displayed, extrapolation of the shallower preloaded fracture toughness distribution suggests, albeit at an extremely low probability, that little or no benefit will be obtained from the prior overload sequence. Clearly, such a postulate requires validation through further testing.


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for non-preloaded and preloaded (V==0.5mm) at 20°C for A533B SENB specimens at -170°C in terms of:

(a) CTOD; (b) K~c.

4.3.2 The amount o f prior overload Earlier, the amount of prior overload was discussed in terms of the ratio of the reference stress to the yield strength. In particular, for a crack length of a / W = 0"5 and an applied crack mouth opening displacement /I= = 0.5 mm, the stress ratio aR/av was about 0-57, and similarly, for V= = 1.8 mm, a, /av was about 1.0. The effect of these two overload levels on the CTOD temperature transition curves for the L-T orientation can be gauged by inspection of Fig. 16(a), where for clarity only the curves fitted through the data points are shown. It can be seen that there is a benefit throughout the temperature range of - 1 9 6 to -20°C, for both amounts of prior overload, even beyond the point of initiation of tearing. At temperatures below about -140°C, which is about the upper temperature for valid K~c results (see Appendix B), there appears to be little difference in benefit obtained between prior overloads of aR/aV of 0"57 and 1"0. However, with increasing temperature, there is an increase in benefit from the higher preload level.

The data corresponding to the T -L orientation are shown in Fig. 16(b). Again, at low temperatures, where cleavage fracture is the dominant mechanism of failure, there is a benefit obtained from preloading at room temperature. However, at temperatures above - 100°C, prior overloading is detrimental. The loss in toughness was greater as the amount of prior overload increased. The changeover from a beneficial effect on the toughness to a loss of toughness appears to occur at about the temperature at which ductile tears are first observed. Thus, for the T -L orientation, loss of

314 D.J. Smith, S. J. Garwood

101

1o 0

| 101

~bz!

(a)

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-80

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(b) Tempera ture,*C Fig. 16. Comparison of curve-fitted data for non-preloaded and preloaded 0"5a/W A533B

SENB specimens for: (a)L-T orientation; (b)T-L orientation.

toughness was obtained from a prior overload sequence when cleavage was no longer the dominant failure mechanism.

The results of the prior overload sequence on the subsequent behaviour of the T-L orientation specimens, expressed in terms of the reference stress at failure, are shown in Fig. 17.

Polynomial curves have been fitted to the data using eqn (6). Although the goodness of fit was poor, the curves serve to illustrate the approximate

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Fig. 18.

I 0 7 •

100 Preloaded 5i - f . . . . . Vg = O.2 m m / / |

~oy--~,.~/,l~ +J / ' / . _ ~ lb 7

~ J I Envelope of

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Comparison of curve-fitted CTOD transition curves for non-preloaded and preloaded 0.2a/W L-T orientation A533B SENB specimens.

behaviour. Because the fit was poor, the individual data points have been retained in Fig. 17. There was a benefit at low temperature. At higher temperatures, contrary to the C T O D results, there was no loss in the load- bearing capacity (in terms of reference stress) even for the high prior overload.

4.3.3 Effect of initial crack length For specimens with a/W=0.2 in the L - T orientation, the effect of prior overload is to provide a benefit over the temperature range - 196 to + 20°C,

100 over/oad , ~

/ / - ~ - 6 i - -

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stress.


as shown in Fig. 18. The benefit is greatest at the lower temperatures, but remains constant for temperatures greater than about -160°C.

For specimens with long cracks ( a / W = 0.7) in the L-T orientation as shown in Fig. 19(a), there is a beneficial effect on the lower shelf from the prior overload. However, contrary to the behaviour of specimens with shorter cracks in this orientation, there is no benefit from prior overload at temperatures above about -140°C. Indeed, prior overloading is clearly detrimental in terms of CTOD within the transition regime. This is also reflected in the load-bearing capacity of the specimen in terms of the reference stress, as shown in Fig. 19(b), particularly at temperatures around -100°C. However, at a slightly higher temperature of -80°C, although there is a significant drop in CTOD after overloading, as shown in Fig. 19(b), the reference stress values are approximately the same.

5 CONCLUSIONS

Based on the experimental study on the mechanical and fracture behaviour of a ferritic steel, A533B, the following can be concluded.

(1) Valid K~c fracture toughness was obtained only for temperatures below - 140°C, for T -L and L-T orientations and crack lengths a / W of 0.2, 0"5 and 0.7. On the lower shelf, at -170°C, the fracture toughness variability, based on a sample of 14 specimens, exhibited a bimodal distribution. This remains unexplained. The initiation of tearing temperature for specimens with crack lengths a / W = 0"5 in the L-T orientation was about -60°C. For similar specimens in the T - L orientation the initiation of tearing temperature was about -80°C. The initiation of tearing as well as the maximum load toughness was found to be dependent on the crack length for the L-T orientation. At temperatures close to and below the initiation of tearing temperature the transition curves for each crack length were found to be practically independent of crack length. However, there was a tendency for higher CTOD values to be obtained in the transition regime for short and long cracks.

(2) On the lower shelf at - 170°C the fracture toughness variability after prior overloading, based on a sample of 14 specimens, was found to exhibit a bimodal distribution similar to that for the non-preloaded material. However, the steeper distribution of the two linear distributions, exhibited in each data set, extended further for the preloaded material. The benefit in toughness compared with the as- received material at -170°C from a prior overload at room temperature ranged from 17% to 89%, with a median value of 70%.


(3) The effect of the prior overload on the fracture toughness at lower temperatures was found to depend significantly on orientation. In the L-T orientation for short and medium length cracks ( a / W of 0.2 and 0"5) there was a benefit throughout the transition regime. In the T - L orientation, however, no benefit was obtained for temperatures greater than the initiation of tearing temperatures. It was found that for these temperatures the prior overloading sequence had a detrimental effect on the fracture toughness. Contrary to the CTOD results, there was no loss in the load-bearing capacity of the specimens. For long cracks (a /W = 0.7) in the L-T orientation it was found that a benefit from proof testing occurred only for temperatures lower than - 140°C. At higher temperatures there was a loss in fracture toughness even though failure was cleavage dominated up to -80°C.

A C K N O W L E D G E M E N T S

The work was funded jointly by research members of The Welding Institute and the Minerals and Metals Division of the UK Department of Trade and Industry. The authors would like to thank their colleagues in the fracture department for helpful discussions and comments. The laboratory staff are thanked for conducting the large number of tests, particularly Mr Neaves for supervising the tests.

REFERENCES

1. Nichols, R. W., The use of overstressing techniques to reduce the risk of subsequent brittle fracture, Parts 1 and 2. Br. Welding J. (January, February 1988).

2. Loss, F. J., Gray, R. A. & Hawthorne, J. R., Significance of ware prestress to crack initiation during thermal shock. NRL/NUREG Report 8165, 1977.

3. Pickles, B. W. & Cowan, A., A review of warm prestressing studies. Int. J. Pres. lies. & Piping, 14 (1983) 95-131.

4. Sutcliffe, D., Jolley, G. & Hopkins, P., Effect of preload on the low temperature toughness properties of a pressure vessel steel. Paper presented at the 2nd Irish Durability and Fracture Conference, Limerick, Ireland, March 1984.

5. Smith, D. J. & Garwood, S. J., The significance of prior overload on fracture resistance: A critical review. Int. J. Pres. Ves. & Piping, 41 (1990) 255-96.

6. ASTM E208-81, Standard Method for Conducting Drop-Weight Test to Determine Nil-ductility Temperature for Ferritic Steels. American Society for Testing and Materials, Philadelphia, PA, 1981.

7. ASTM E23-82, Standard Method for Notched Bar Impact Testing of Metallic Materials. American Society for Testing and Materials, Philadelphia, PA, 1982.


8. British Standards Institution, Method for Crack Opening Displacement (COD) Testing. BS 5762. BSI, London, 1979.

9. Sumpter, J. D. G. & Turner, C. E., Method for laboratory determination o f J c. In Cracks and Fracture Proc. 9th National Syrup. Fracture Mechanics. ASTM STP601. American Society for Testing and Materials, Philadelphia, PA, 1976, pp. 3-18.

10. Anderson, R. G., Some observations on reference stresses, skeletal points, limit loads and finite elements. In Creep in Structures, IUTAM 3rd Syrup., 1980.


APPENDIX A: RESULTS FROM SENB FRACTURE TOUGHNESS TESTS

A1 As-received material

(i) 0.2a/W, L - T orientation

Specimen Temperature Crack Fracture parameters number (°C) growth, Aa

(mm) ~o Kmax Jo crR~ (MPa) (MPa~/m) (MJ/m 2) (MPa)

129 - 196 0 0-003 7 41"03 0-007 2 119"0 130 - 160 0 0-005 3 43-48 0-008 1 125"9 1 3 1 - 120 0 0"034 9 94"69 0-007 1 274"0 132 - 8 0 0 0-2204 168-10 0"2133 490"6 133 - 4 0 0 0"889 1 152"00 0"858 5 442-9 134 0 5"16 2'1790 185'60 4"0890 544"5 141 - 180 0 0.004 5 42'39 0-007 7 122-7 142 - 140 0 0"032 8 94-20 0"043 9 272"2 143 - 1 0 0 0 0"1243 152"70 0"1225 441"5 144 - 6 0 0"70 0"7458 185'00 0"701 9 534"6 145 - 20 3"91 3"026 0 199"50 2"928 0 577"3 146 20 6"46 4'1730 197"90 4"1590 571"7

(ii) 0.5a/W, L - T orientation

Specimen Temperature Crack Fracture parameters number (° C) growth, Aa

(mm) 3 o Kma x (mm) (MPa~/m)

Jo O'Rc (M J/m) (MPa)

70 - 130 0 0"080 8 136"2 0"092 8 443-7 74 - 1 5 0 0 0.0046 38"7 0"0065 126-1 75 - 140 0 0"0160 69"1 0-021 6 225"8 78 - 120 0 0"058 2 117'2 0"066 9 383"1 91 - 160 0 0"011 5 54'2 0"0134 177"6 93 - 1 7 0 0 0.0036 36'8 0"0058 121-1 95 - 196 0 0-002 6 34"4 0"005 1 112-6

111 - 1 8 0 0 0"0030 35"0 0"0053 114"8 1 1 2 - 80 0 0"127 1 148"4 0"134 1 486"1 113 - 4 0 2'35 1"7720 178"3 1'3620 578'7 114 0 2"08 1-9110 178"7 1"934 0 580"5 1 1 9 - 100 0 0.028 4 91'4 0"036 2 298"0 120 - 6 0 0 0"267 3 163"2 0"271 7 544"8 121 - 2 0 4"04 1"981 0 186-8 2-0920 616"0 122 20 1"76 1"479 0 181"8 1-544 0 592"4 127 40 4"87 2"1030 178"5 3"0460 582"4 128 - 130 0 0'008 4 51"5 0-011 4 167"4

c o n t i n u e d on n e x t p a g e


(iii) 0"5a/W, T-L orientation

Specimen Temperature Crack number (° C) growth, Aa

(mm)

Fracture parameters

60 K~ go , ~ (ram) (MPax/m) (Md/m 2) (MPa)

27 - 3 5 2.46 0"9290 164.9 0'884 6 542'9 28 - 7 0 0,06 0-2103 162.8 0'2089 536'5 29 - 1 2 0 0 0'0763 139.0 0"0929 459'5 30 - 8 0 0 0'119 6 144.5 0-123 5 475'9 31 - 5 0 0"04 0"235 1 158.5 0'223 7 521.7 39 - 1 4 0 0 0'003 8 43.7 0'008 2 143.5 41 - 1 9 6 0 0.005 7 51'0 0'001 1 167.1 42 - 1 0 0 0"05 0.129 4 155.6 0.143 8 508-5 45 - 1 8 0 0 0'0049 44.4 0"008 6 147.8 49 - 4 0 0.15 0'301 1 163.4 0"279 3 530"2 50 - 6 0 0,08 0"3046 167-5 0"292 3 551"8 51 - 1 6 0 0 0-007 9 53"2 0"012 3 175-5 52 - 3 0 2,67 0"455 6 168-4 0"703 3 533"3 53 - 9 0 0 0"1104 144.8 0.117 4 478.4 54 20 3.97 0.441 4 155.1 0,745 6 513.1

192 - 1 7 0 0 0.013 0 66.1 0'018 9 218-4 193 - 1 5 0 0 0'0090 55-6 0,0134 185-3 194 - 1 3 0 0 0"051 0 111-5 0,053 9 365"8 195 - 1 1 0 0 0-072 0 123.7 0'066 3 404.8

(iv) 0.7a/W, L-T orientation

Specimen Temperature Crack number (°C) growth, Aa

(mm)

Fracture parameters

6o Km~x Jo ~Rc {mm) (MPax/m) (M Jim 2) (MPa)

79 20 2"96 80 - 4 0 2"13 81 - 80 0"15 82 - 1 2 0 0 83 0 2'33 84 20 2'09 89 - 160 0

101 - 2 0 1'73 102 - 6 0 0'33 103 -- 100 0 104 - 140 0 105 - 100 0 109 - 180 0

1"227 0 131" 1 1"420 0 577"9 1"1550 145'2 1'3370 620"3 0'232 4 132"4 0"234 9 564"9 0-119 6 151"9 0"172 8 669"8 1"161 0 143'4 1.495 0 640"4 1"272 0 138-9 1'399 0 611"3 0"007 6 52"0 0"011 7 249"8 1"230 0 136"7 1'269 0 604"0 0"475 0 136-3 0'487 3 607 5 0-031 5 96.2 0"040 1 415'7 0"008 1 51"4 0"011 4 223'3 0"258 5 36'6 0'113 1 594'2 0-004 7 43"7 0'008 6 184"7

Prior overload effects on fracture toughness: Experimental study

(v) 0.5a/W, L-T orientations, all tests at -170°C

321

Specimen Fracture parameters number

~0 gmax O'Rc ( m m ) (MPax/m) (MPa)

163 0"014 70"9 235'1 164 0"005 43-5 142'3 165 0"008 54-5 180"4 166 0'015 74'5 243"7 167 0"008 53"9 178"0 168 0-014 73'1 242'6 169 0"008 54"8 179'9 170 0-011 65'0 214"7 171 0"010 61"8 205-0 172 0'007 50'1 167-4 173 0"012 67"7 233'0 174 0-014 71"9 237'3 175 0"006 45"7 151"5 93 0'004 36"9 121"0

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81"5

0-

028

8 26

6"6

0 0-

033

7 10

5"9

0-05

0 3

344"

7 ,~

0

0"06

5 1

128"

0 0"

078

1 42

3"0

0 0"

074

4 13

5"3

0"08

9 0

447-

0 ~a

- 0

0"01

2 8

71'5

0-

022

1 23

5"6

0 0

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63

16

6-9

0"19

96

549"

4 2"

94

1"94

70

192"

9 2"

1000

62

4"8

t~

2"86

1-

8150

18

4-0

2"05

70

595'

2 0

0-06

9 7

128'

5 0"

082

2 41

1'5

-~"

0"50

0"

5190

17

5"0

0-52

95

570"

1

2"93

1-

9560

17

8-5

1'99

1 0

591-

4

a V

g = 0

"8 m

m.

con

tin

ued

on

nex

t p

ag

e

(iii

) V

, = 1

.8 m

m, 0.54 K L-T o

rien

tati

on

Spe

cim

en

nu

mbe

r P

relo

ad p

aram

eter

s

aRl

WfW

Tem

pera

ture

C

rack

(“C

) gr

owth

, A

u

(mm

)

71

0.37

4 2

160.

9 0.

341

1 52

1.5

-196

0.

16

@01

14

72

0.36

1 9

155.

5 0.

325

7 50

6.2

-140

0.

15

0.03

2 1

73

0.37

0 9

1544

0.

336

3 49

8.9

-160

0.

20

0.02

16

98

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153.

3 0.

3137

50

7.9

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0.

14

0.10

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293

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11

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153-

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317

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16

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152

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158.

6 o-

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20

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63

153

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389

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10

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0.34

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517.

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77

0.50

15

7 03

618

132.

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320

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25

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Fra

ctur

e pa

ram

eter

s

4 (m

m)

(M.$

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5 s

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T - 57

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8.0

&

615.

5 62

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5865

56

1.0

608.

3

(iv)

V

~ = 0

"5 m

m,

0"5a

/W, T

-L o

rien

tati

on

Spec

imen

P

relo

ad pa

ram

eter

s nu

mbe

r

(mm

) (M

Pax

/m)

(MJ/

m 2)

~Y

R1

(MP

a)

Tem

pera

ture

(°

C)

Cra

ck

grow

th, A

a (r

am)

Fra

ctur

e par

amet

ers

61 "

Kfm

ax

Jf

o'Rf

(mm

) (M

Pax

/m)

(M Ji

m 2)

(M

Pa)

E

"

32

0-03

3 86

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0"03

4 75

33

0"

029

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0-03

2 22

34

0"

030

7 86

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0"03

4 68

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0

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02

86

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40

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031

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343

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037

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19

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0469

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301

197

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32"1

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0365

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330"

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7-0

285.

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32

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70

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128"

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199

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20

551"

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29

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04

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149"

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118

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0-2

"~,~

309"

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170

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082

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0-04

2 4

316"

6 29

3"2

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00

106"

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069

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130

0 0"

0400

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6-8

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24

354"

8 ,~

, 30

8-6

- 11

0

° V

~ = 0

"55

mm

.

con

tin

ued

on

nex

t p

ag

e ta

~

(v)

Vs=

1.8

mm

, 0.

5a/W

, T-L

ori

enta

tio

n

Spec

imen

P

relo

ad pa

ram

eter

s nu

mbe

r

(mm

) (m

Pa~

/m)

(M Ji

m 2)

(M

Pa)

Tem

pera

ture

(°

C)

Cra

ck

grow

th, A

a (r

am)

Fra

ctur

e pa

ram

eter

s

~,

Kfm

~x

J~

aR,

(ram

) (M

Pa~

/m)

(M Ji

m 2)

(M

Pa)

37

0-31

7 1

147"

7 0"

2922

49

1-1

-80

38

0"37

0 9

154"

6 0-

334

3 51

1"8

- 40

0"

58

0"07

5 9

137-

8 0"

086

6 45

8"2

44

0'33

2 1

162-

6 0"

3180

52

9'4

- 10

0 1"

28

0-17

52

160-

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2-2

47

0"33

25

151'

7 0"

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50

4"6

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0

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0"

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15

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0"10

80

507"

2 56

0"

3537

15

4"7

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23

508"

1 -1

60

0"

13

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54

78'6

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0268

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156"

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96

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030

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6"6

63

0"36

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154-

8 0-

3365

50

2-0

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6

1-28

0-

1450

16

8-4

0-16

23

559"

6 64

0-

3369

15

3"7

0"30

45

507"

3 -

140

0"04

0.

011

9 73

-5

0-02

34

238-

6 68

0"

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15

9"5

0"31

19

524"

5 -9

0

0"82

0"

1029

16

1-4

0"12

24

531"

0 20

0 0"

372

1 15

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0-33

78

516-

0 -

170

0"69

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0200

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0"03

1 !

280"

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1 0

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54

13

5"1

0-32

50

515"

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150

0"41

0"

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10

1"1

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7 1

335"

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2 0-

358

1 15

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0"32

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518"

8 -

130

0-47

0-

0640

13

1"5

0"08

1 8

435"

9 20

3 0-

3608

16

6"3

0-34

96

562"

4 -1

10

0"

46

0-09

70

156"

1 0-

1166

52

7"7

(vi)

V

g = 0

"7 m

m,

0.7a

/W, L

-T o

rien

tati

on

q~

e~

Spec

imen

P

relo

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aram

eter

s nu

mbe

r 61

K

~ max

J1

O

'R 1

(mm

) (M

eax/

m)

(M Ji

m 2)

(M

Pa

)

Tem

pera

ture

(°

C)

Cra

ck

grow

th, A

a (r

am)

Fra

ctur

e par

amet

ers

t~f

Kfma

x Jf

O'

Rf

(mm

) (M

Pa~/

m)

(M Ji

m 2)

(M

Pa)

q~

85

0-01

64

72"2

0"

024

1 32

4"8

-- 1

80

86

0-02

4 8

71"8

0"

025

4 31

2"7

-- 6

0 87

0"

023

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0-02

6 8

327"

0 --

100

88

0"

020

8 70

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0"02

2 5

328"

2 --

140

10

6 0-

024

2 73

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0-02

5 6

316-

7 --

80

107

0"02

1 4

71"8

0"

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7 31

7"1

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0 10

8 0-

020

8 71

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0-02

5 3

302"

3 --

160

15

9 0"

023

4 75

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0"02

4 8

334"

3 20

16

0 0"

028

1 75

-6

0"02

6 8

324"

9 0

161

0"02

74

73"0

0"

020

1 31

7-5

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16

2 0"

032

8 79

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0"03

1 5

340"

9 --

40

0 0"

013

6 73

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0-02

3 5

331"

2 '~

e~

0

0"16

45

132"

4 0"

1724

57

5-7

~"

0 0-

022

1 80

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0"02

8 2

349"

1

0 0"

039

6 10

0"3

0-04

8 5

464-

0 0

0-09

75

129-

5 0"

103

1 55

6"9

~ 0

0"02

1 8

81"6

0"

028

9 36

0"6

0 0-

0170

74

"7

0"02

49

311"

1 6"

56

1-08

13

4-1

1"15

50

593"

7 ~

1"36

1"

10

142"

2 1"

1630

61

0-6

5"

0"77

0"

81

140-

9 0"

982

1 61

2-9

~ 1-

31

1"25

15

3"0

0"96

8 7

652-

6

cont

inue

d on

nex

t p

ag

e ..~

328 D . J . Smith, S. J. Garwood

(vii) V~ = 0"5 mm, 0.5a/W, L - T orientation, fractured at - 1 7 0 ° C

Specimen Preload parameters number

Fracture parameters

1 KI max O'R 1 ~f gfmax ORf (ram) (MPa~/m) (MPa) (ram) (Mea~/m) (Mea)

179 0"030 3 80-8 279"7 0-027 99"6 344'9 177 0"022 0 89"7 296"4 0"035 111'5 368'3 178 0-036 6 89-7 296-9 0"033 110"7 366'7 179 0-038 8 90"2 298"9 0"029 102'5 339'6 180 0"037 8 88-8 293'9 0-026 96-8 326"3 181 0'0398 91-5 302'8 0"010 61"4 203'1 182 0'0390 89"2 298'0 0'032 108"3 361'9 183 0'040 2 90-8 302"6 0"026 97"7 325"7 184 0'037 2 90-5 300'4 0-033 111' 1 368"5 185 0"0405 92-5 306'4 0"029 104"3 345'4 186 0'037 1 86-7 289"5 0"018 81"3 271"3 204 0'040 0 90"4 299'3 0'030 105'4 348"9 205 0'035 8 86"4 289'7 0"023 92'3 309"4

69 0"055 5 83.1 283"5 0-020 79"4 264"6

APPENDIX B: FURTHER COMMENTS ON BASIC MATERIAL DATA

B1 Comparison of tensile data

For convenience an upper bound curve was drawn through the yield stress data illustrated in Fig. 2(a) and a fifth-order polynomial was fitted to the upper bound curve, to give

5

+ ) ' ciT i - 1 9 6 ° C < T < 20°C (B1) O'y Co

i = 1

where the coefficients Co and ci are given in Table B1. The tensile data

TABLE B1 Polynomial Coefficients for Upper Bound Yield Stress Data

i C i i C i

0 0 .542506×10 3 3 - 0 - 3 7 4 0 9 4 x 1 0 4 1 -0 '642481 4 -0 -45829 × 10 - 6

2 0'114938 × 10 -2 5 - 0 ' 2 2 8 864 × 10 -8


1000 , , , , , , , ,

~: 53381 HSST

~, U p p e r b o u n d • ~ , , , ~ ~

data I

/~0 I I I I i I I I I

-200 - 8 0 40

Tempera tu re . *E

Fig. BI. Comparison of yield stress data for A533B steel.

obtained in the present study can be compared with existing data B1 on A533B steel. Ritchie et al . s l used the following expression as a fit to yield strength data:

a v = 31 370.9[(9 + 49].7) in (]0a/~)] -°'431 (B2)

where the units of T are degrees centigrade and av is in MPa. A comparison between eqns (B1) and (B2) is shown in Fig. B1, where the strain rate ~ in eqn (B2) is taken as approximately 1 × 10 -3 s -~. Bearing in mind that eqn (B I ) represents an upper bound of the yield strength data obtained in this study, at the two extremes of the temperature range (20 and -196°C) there is reasonable agreement. At intermediate temperatures the yield strengths presented here are lower than those reported by Ritchie et al. BI

B2 Fracture toughness-- lower shelf validity

In the following, an indication will be given of the range over which valid KI¢ values have been obtained. To measure a valid Kit the specimen dimensions are required to be B2'B3

B,( W - ao),a o > 2.5(K[Jav) 2 (B3)

The ratio Kj/av is plotted in Fig. B2 for the range of initial crack lengths and specimen orientations examined in the present study, where Kj has been evaluated from

Kj = ( J E ) 1/2 (B4)

Also shown in Fig. B2 are the validity limits according to eqn (B3). It can be seen that for temperatures below -140°C valid Kic values are obtained for


Fig. B2.

%. .

03

0"2

0"I

v ~

o 121

6 9 ~ A

[3

13 9

. . . . 0_0 0

i I

{3 D Test data

a a/W=0.5) o a/W=~2~L-T A alk~0.7 w y a/W=0.5 PL

Solid symbols represen v a l i d KIc

a / W = 0 . 5

a/W=O.7 IKI;

a/W=0.2 i Kic VmyA• I i~l•V y

0 ~ ~ J J ~ ~ i 4L0 t i

-2~0 -160 -120 -BO - 0

Temperature, °C

Application of fracture toughness criterion for A533B SENB specimens.

the different initial crack lengths and the two orientations. There is, therefore, a significant temperature region over which there is, according to the UK and US Standards, 82"83 insufficient constraint to provide valid plane strain K~c results. Nevertheless, cleavage fracture occurred for a substantial temperature range above -140°C, particularly for the L-T orientation.

B3 Comparison of fracture properties with other A533B data

The mechanical and fracture behaviour of A533B has been well documented, particularly by the US Atomic Energy Commission Heavy Section Steel Technology (HSST) Program. B4 The fracture toughness data from the HSST Program has also been examined by Ritchie et aL B1 for the purposes of comparison with models for predicting upper and lower shelf fracture toughness. The data from the HSST Program is shown in Fig. B3, together with the data from the L-T and T-L orientation specimens with crack lengths of a /W= 0"5. The fracture toughness Kj has been evaluated for the present data using eqn (B4). The data (all valid ASTM K~c) from the HSST Program shown in Fig. B3 include 1-11 in thick compact tensile (CT) specimens from the L-T orientation and 1 in thick CT specimens in the T-L orientation.

On the lower shelf (below -140°C) it can be seen that there is agreement between the two data sets, with slightly higher fracture toughness values for the T-L orientation than for the L-T orientation.

From a temperature above - 140°C the present data rise above the HSST data. Earlier it was determined that valid K~c data were obtained in the


3001 ' ' ~ i I n r-LIsa m [ ,, L-r , "

=. I -~ Af- fTO*E median I Z sgm~,em

x o

~" I x ._.o."~" ~ ~,,

_

I I I I I I I I I

-200 - 100

Temperafuee,'C

Fig. B3. Comparison of fracture toughness data for A533B steel.

i i i 1600 o ' a - 3 ' ' ' • • - s B=25.45mm

] A B ' I O 1400 • B - z 5 ~ n . m m ~

o 8 - 4 6 ~ " " ' . . . . • a - ~ o ~o Aj~ A I

c'~ X Kc - KA= ] l e l a l I 1 0 0 0 fraraASrM v i i " / - / /

.o o l ; , ' / 1 1 7 / = : 3 o - - - o l l j / / / ,ooto.,oo /,Oi/

# A V Z -o[ '

01 I I I I I I

-200 -160 -120 -80 - l ,0 0 ~0

Tesf femperafure, T, of

Fig. !!4. Variation of fracture toughness with specimen thickness for A533B steel

(after Ref. B5).

present study only for temperatures below - 140°C. The deviation from the lower shelf exhibited by A533B steel has also been explored by Hagiwara. B~ The influence of thickness on the transition curves obtained by Hagiwara is shown in Fig. B4. The deviation from the lower shelf for thicknesses ranging from 25 to 40 mm also occurs at about -140°C.

References

B1. Ritchie, R. O., Server, W. L. & Wullaert, R. A., Critical fracture stress and fracture strain models for the prediction of lower and applied shelf toughness in nuclear pressure vessel steels. Met. Trans., A, 10A (1979) 1557-70.

B2. British Standards Institution, Methods of Test for Plane Strain Fracture Toughness (Kic) of Metallic Materials. BS 5447. BSI, London, 1977.

B3. ASTM E399-81, Standard Method of Test for Plane Strain Fracture Toughness of Metallic Materials. American Society for Testing and Materials, Philadel- phia, PA, 1981.

B4. US Atomic Energy Commission Heavy Section Steel Technology (HSST) Program, HSST Technical Reports Nos 1-36, 1968-1975.

B5. Hagiwara, Y., Evaluation of thickness effect on fracture toughness in heavy sectioned steels. In 2nd Japanese-German Joint Seminar on Non-Destructive Evaluation and Structural Strength in Nuclear Power Plants, October 1982.

experimental study of effects of prior overload on fracture toughness of a533b steel

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