Low temperature fracture properties of DIN 22NiMoCr37 steel in fine-grained bainite and coarse-grained tempered embrittled martensite microstructures

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Engineering Fracture Mechanics 75 (2008) 2480–2513 www.elsevier.com/locate/engfracmech

Low temperature fracture properties of DIN 22NiMoCr37 steel in fine-grained bainite and coarse-grained tempered embrittled martensite microstructures M.J. Balart *, J.F. Knott Department of Metallurgy and Materials, School of Engineering, The University of Birmingham, Edgbaston, Birmingham B15 2TT, UK Received 13 December 2006; received in revised form 31 July 2007; accepted 31 July 2007 Available online 22 September 2007

Abstract An as-received (AR) DIN 22NiMoCr37 nuclear reactor pressure vessel steel has been heat treated for 1 h at austenitising temperatures of 1373 and 1473 K to obtain different austenite grain sizes. After austenitising, the samples were water quenched, tempered for 2 h at 923 K, water quenched and then held isothermally at 793 K for 180 h before final air-cooling. The AR condition had a tempered bainite microstructure and a prior austenite grain size of 30 lm, whereas the heat treated conditions were tempered martensite and had a prior austenite grain size of approximately 100 lm for the 1373 K condition and ‘extraordinary’ large austenite grains (>1 mm diameter) for the 1473 K condition. Their low temperature fracture properties were determined and were related to the susceptibility to segregation induced embrittlement. Despite the heat treated conditions having a larger prior austenite grain size compared to the AR condition, at a given testing temperature, the tempered martensitic 1373 K condition generally exhibited higher strength and higher fracture toughness values at 123 K. The heat treated conditions generally exhibited higher local fracture stress (rf) values in 0.2 mm blunt notch SE(B)-0.4T specimens at 123 and 77 K.  2007 Elsevier Ltd. All rights reserved. Keywords: Intergranular fracture; Transgranular fracture; Toughness testing; Power plants; Pressurised components

1. Introduction The steel DIN 22NiMoCr37 is an European forging grade of a type used for the fabrication of nuclear reactor pressure vessels (RPVs). It is similar to the ASTM forging grade A508 Cl.2 [1] which has generally replaced ASTM A533B plate material for RPV manufacture. In such manufacture, ring forgings (or plates), domes and nozzles are welded together and it is therefore of interest to assess the fracture resistance of the weld heataffected-zones (HAZ), in particular the coarse-grained HAZ (CGHAZ). A particular heat of DIN 22NiMoCr37 was used to generate the Euro dataset for fracture toughness: 779 compact tension (C(T)) tests *

Corresponding author. Tel.: +44 (0)121 4143265; fax: +44 (0)121 4145232. E-mail addresses: [email protected] (M.J. Balart), [email protected] (J.F. Knott).

0013-7944/$ - see front matter  2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.engfracmech.2007.07.023

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ranging from C(T)-0.5T to C(T)-4T tested at temperatures from 119 to 293 K [1,2]. We have been able to cut blocks from undeformed pieces of some of the larger Euro dataset testpieces, subject these to furnace heat treatments designed to simulate the CGHAZs in service welds and determine their resistance to fracture at low temperatures. One issue is that the cooling rate in the centres of large blocks of quenched as-received (AR) material is such as it generates bainite, rather than martensite microstructures. Earlier work on A533B [3] showed that tempered bainites possess lower yield strengths, fracture toughnesses and values of local fracture stress in notched bars than do tempered martensites. It could therefore be argued that the simulated CGHAZ samples would exhibit properties superior to those of AR material. The fracture mode observed by Bowen et al. [3] was, however, in all cases, transgranular cleavage. It has been demonstrated [4] that in A533 class 1 a tempered coarse grain martensite is more susceptible to intergranular embrittlement than is a tempered bainite of the same grain size. Phosphorus segregation to prior austenite grain boundaries and subsequent detrimental effects on fracture resistance have been demonstrated in A508 class 3 [5] and in A533B [6]. Segregation could occur in practice during the slow cooling after the stress-relief heat treatment or, possibly, as a result of irradiation enhanced segregation of trace impurity elements to prior austenite grain boundaries during service operation at approx. 570 K [7]. The heat of DIN 22NiMoCr37 used for the Euro dataset was designed to be ultraclean, from the point of view of trace impurities, but it was thought sensible to include an embrittlement step in the CGHAZ simulation, to discover whether intergranular fracture could be induced and whether this lowered the fracture resistance. Early heat treatment trials reported in [8] showed that an austenitising treatment of 1 h at 1373 K gave a grain size of approx. 100 lm, typical of many CGHAZ regions in service welds, but that 1 h at 1473 K gave a grain size of over 1 mm. More detail is given later, but, one consequence is that care has to be exercised in deciding how best to assess the fracture resistance of such very large grain size material. The work of Ritchie et al. [9] indicates that sharp crack fracture toughness tests can be misleading and non-conservative compared with notched bar tests containing large root radii. In the present paper, we have carried out tensile tests, fracture toughness tests and blunt notched tests combined with detailed fractography to compare and contrast features on the fracture surfaces in the different types of test. 2. Experimental procedure 2.1. As-received DIN 22NiMoCr37 steel, heat treatment and specimen location The chemical composition of the DIN 22NiMoCr37 steel is given in Table 1. The steel was taken from the C(T)-4T broken halves SX20.1 (1 half), SX25.2 (2 halves) and SX25.3 (1 half) used in the Euro dataset [1,2]. Fig. 1 depicts how blocks for heat treatment and, subsequently, specimens, were sampled from the original and heat treated steel blocks. The dimensions of the blocks for heat treatment were B/3 = 33.3 mm, E/2 = 120 mm and G = 70 mm. Based on the homogeneity checks and fracture toughnesses described in Ref. [2], tensile testpieces and the fatigue crack tips and blunt notch roots were located within a 100 mm thick ‘homogeneous’ layer inside the wall of the ring segment as drawn in Fig. 1. The identification system of the specimens in the present investigation starts with the C(T)-4T specimen designation of Ref. [2], followed by ‘-’ and subsequently the positions i, j and k given in Fig. 1. For example, specimen SX20.1-a.2.3 stands for C(T)-4T specimen location from the steel block SX20, position 1 of Ref. [2] and positions i = a, j = 2 and k = 3 of Fig. 1. Specimen SX25.2-a.2 0 + 1 0 . 7 stands for C(T)-4T specimen location from the steel block SX25, position 2 of Ref. [2] and positions i = a, j = 2 0 and 1 0 (not cut) and k = 7 of Fig. 1. The AR condition had a tempered bainitic microstructure, and its fracture properties have been previously determined in the generation of the Euro fracture toughness dataset [2]. Table 1 Chemical composition (wt.%) of DIN 22NiMoCr37 (see Ref. [8] for details) C

Si

P

S

Cr

Mn Ni

Cu

Mo Sn

Al

As

B

Sb

Ti

Nb

V

N

O

0.21

0.22

0.005 0.004 0.41 0.87 0.87 0.06 0.52 0.007 0.014 0.006 <0.0005 0.002 0.0004 0.002 0.004 0.008 0.0005

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INSIDE 0

120

OUTSIDE 280

220 155

190

250

L

S

1’

T (B) G

2’ 2 1 j = {1’, 2’, 2, 1}

30

0 INSIDE

i = {a, b, c}

i = {a, b, c}

150 185

120

a b c a b c

250

220

280 OUTSIDE

1 3 5 7 9 2 4 6 8 10

k = {1, 2, 3, 4, 5, 6, 7, 8, 9, 10}

B B = W = 10 mm

1 3 5 7 2 4 6 8

k = {1, 2, 3, 4, 5, 6, 7, 8}

B B = W = 12.7 mm

W (mm)

A (mm)

B (mm)

C (mm)

D (mm)

E (mm)

F (mm)

G (mm)

200 ± 0.2

250 ± 0.2

100 ± 0.2

50 H11

105 ± 0.1

240 ± 0.2

150 ± 0.1

70

Fig. 1. Location of specimens machined from the C(T)-4T used in the Euro dataset [1,2].

Some samples of the steel were austenitised for 1 h at 1373 K to simulate the grain size of an observed CGHAZ microstructure and determine its fracture properties; others were austenitised for 1 h at 1473 K to simulate ‘extra large’ austenite grains in the HAZ and determine their fracture properties for reference. After austenitising, samples were water quenched, tempered at 923 K for 2 h to simulate stress-relief, water quenched and held isothermally at 793 K for 180 h, followed by air cooling, to maximise any susceptibility to intergranular embrittlement. Metallographic samples were etched in a solution of 1 gr picric acid, 10 ml HCl, 10 ml HNO3 and 80 ml ethanol to reveal the prior austenite grain boundaries. In reporting observations, samples are identified by austenitisation temperature and the specimen location. 2.2. Tensile testing Tensile specimens were machined with their length parallel to the L direction [10]. Tensile tests were carried out using a Zwick 1484 screw-driven machine of 200 kN capacity in displacement control at a crosshead speed of 0.5 mm/min at temperatures of 123 K and liquid nitrogen. The initial testpiece

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diameter was 5 mm. The initial gauge length was 28 mm and the displacement was monitored continuously by a linear variable displacement transducer (LVDT). The yield strength (YS) and ultimate tensile strength (UTS) were measured. The AR condition tested at both temperatures and the heat treated steel tested at 77 K exhibited upper and lower yield points and the lower yield point value was taken as the yield strength. The heat treated steel tested at 123 K exhibited continuous yielding and the YS was taken as the 0.2% proof strength. From the instantaneous measurements of load, true strain (and hence instantaneous cross-sectional area), the strain hardening exponent (n) was calculated from the true stress (r)– true strain (e) relationship r = Ken in the region of uniform plastic deformation, where K is a constant, assuming constant volume deformation. The elongation and reduction of area values were obtained after fracture by putting the specimen back together and taking measurements of the gauge length marks and the diameter of the smallest cross-section [11]. 2.3. Fracture toughness testing Precracked Charpy specimens were prepared from single-edge notched bend bars, with thickness (B) = width (W) = 10 mm (SE(B)-0.4T) and the span between outer loading points, S = 4W = 40 mm, machined in the LS direction [10]. Here L is the direction normal to the subsequent crack plane and S is the direction of subsequent crack propagation. L is the longitudinal direction and S is the short transverse direction. The SE(B)-0.4T samples were precracked in fatigue at room temperature to 0.45 6 a0/W 6 0.55 using a machine of 20 kN capacity. A ratio of minimum to maximum stress intensity factor, Kmin and Kmax, respectively, of 0.1 was used for all the fatigue precracking. Kmax did not exceed 0.6 (YS1/YS2)KQ, where YS1 and YS2 are the yield strengths at the fatigue precracking temperature T1 and the yield strengths at the testing temperature T2 in accordance with the requirements given in ASTM E 399-90 [10]. Precracked SE(B)-0.4T fracture toughness tests, for the AR and 1373 conditions, were carried out using a DMG screw-driven machine of 50 kN capacity in displacement control at a crosshead speed of 0.5 mm/min in three-point bending. The testing temperature was 123 K, obtained using different amounts of liquid nitrogen in a bath below the specimen. The temperature was monitored using a Type N Nicrosil–Nisil thermocouple to a temperature control of ±2 K. Values of fracture toughness were obtained from values of fracture load using standard compliance functions. Probability values were calculated from median ranking. This encompassed ordering the data, for N number of specimens, by rank (i) and gave the estimate of the rank probability: Pf ¼

ði  0:3Þ ðN þ 0:4Þ

ð1Þ

notch radius 0.2 mm notch angle 45o

a

3.33 mm

10 mm 10 mm

10 mm

30 mm

10 mm notch radius 1.00 mm notch angle 90o

b

12.7 mm 4.23 mm 12.7 mm 16 mm

32 mm

16 mm

Fig. 2. The blunt notch testpiece designs: (a) subsized Griffiths and Owen’s specimen geometry type and (b) specimen dimensions similar to Griffiths and Owen’s specimen geometry but larger notch root radius (1 mm) and larger notch angle (90) [14].

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The non-dimensional deformation parameter M, defined from Eq. (2), was determined: M¼

bo r y Jc

ð2Þ

where bo is the ligament size in the testpiece and ry is the YS at the toughness testing temperature. Here we use the identity sffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi EJ c K Jc ¼ ð3Þ ð1  m2 Þ where E is Young’s modulus and m is Poisson’s ratio. The value of E at 123 and 77 K was estimated as 217.5 and 219.8 GPa using the expression [12] E ¼ 210  0:05T

ð4Þ

Fig. 3. The variation of stress intensification with applied load in blunt notch testpieces: (a) of Griffiths and Owen’s [13] and (b) of Fig. 2b and other two notch geometries for Ref. [14]. The curve for the notch radius of 1 mm and notch angle 90 has been extrapolated.

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Fig. 4. Macrograph of the 1473 condition.

Table 2 Tensile properties Specimen

123 K YS (MPa)

AR SX25.3c.2 + 1.3

Mean

779 (830)

Specimen UTS (MPa)

El (%)

RA (%)

n

866

18.4

54

0.136

77 K YS (MPa)

AR SX25.3c.2 + 1.1 SX25.3c.2 + 1.2

UTS (MPa)

El (%)

RA (%)

n

917 (978)

978

18.0

42

0.121

917 (988)

988

18.7

45

0.134

917 (983)

983

18.4

44

0.128

779 (830)

866

18.4

54

0.136

Mean 1373 SX25.2a.2 0 + 1 0 .7

1182 (1192)

1227

17.1

36

0.064

1373 SX25.2c.2 + 1.1 SX25.2a.2 0 + 1 0 .8 SX25.2c.2 + 1.2

978

1070

16.6

49

0.071

1013

1100

15.0

45

0.072

1044

1108

16.0

46

0.053

Mean

1012

1093

15.9

47

0.065

Mean

1182 (1192)

1227

17.1

36

0.064

988

1080

11.6

29

0.066

1473 SX25.2c.2 0 + 1 0 .3

1192 (1207)

1233

14.5

15

0.074

993

1070

13.8

23

0.063

1014

1080

12.9

23

0.066

998

1077

12.8

25

0.065

Mean

1192 (1207)

1233

14.5

15

0.074

1473 SX25.2c.2 0 + 1 0 .6 SX25.2c.2 0 + 1 0 .5 SX25.2c.2 0 + 1 0 .4 Mean

(Upper) and lower yield points values. YS – 0.2% yield strength; UTS – ultimate tensile strength; El – elongation to failure; RA – reduction of area; n – strain hardening exponent.

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where E is in GPa and T is in C. Poisson’s ratio is taken as 0.3. We have determined values of M using our KQ values rather than KJc. 2.4. Blunt notch four-point bend testing SE(B)-0.4T specimens, corresponding to a subsized Griffiths and Owen specimen geometry [13] with B = W = 10 mm, a 45 notch, a 1/3 of notch length-to-width (a/W) ratio (i.e. 3.33 mm notch depth) and 0.2 mm notch root radius, Fig. 2a, were machined in the LS orientation [10]. The blunt notch specimens were tested in slow bend for the AR and the 1373 and 1473 conditions. The span between the outer loading points was 50 mm and that between the inner loading points was 30 mm. An alternative testpiece design of similar dimensions to the Griffiths and Owen geometry 12.7 mm · 12.7 mm with a 64 mm outer span and a 32 mm inner span, but larger notch root radius (1 mm) and larger notch angle (90) was tested for the 1373 and 1473 conditions. This testpiece design is shown in Fig. 2b [14]. The slow notch bend tests were carried out using a Servohydraulic ESH machine of 200 kN capacity in displacement control at a crosshead speed of 0.5 mm/min at temperatures of 123 and 77 K in four-point bending. The local fracture stresses rf were calculated from normalized plots of the variation of the maximum tensile stress intensification (r1max) with the ratio of nominal bending stress at the notch root to general yield stress of Fig. 3a, for the subsized Griffiths and Owen’s specimen geometry type, and Fig. 3b, for the specimen with larger notch root radius (1 mm) and larger notch angle (90). The rnom/ry values at general yield are 2.24 ± 0.06 and 2.18 ± 0.01, respectively. For the second geometry fracture generally occurred beyond general yield, somewhat outside the range of the finite element analysis (FEA); values of r1max/ry here have been obtained by extrapolation as indicated by the dashed line. It is assumed that the local fracture stress rf is the peak stress r1max ahead of the notch root at failure. To calculate more precise rf values, it would be necessary to input the true stress–strain curves for this material into the FEA.

Fig. 5. SEM micrographs of the fracture surfaces from longitudinal tensile tests of the AR condition (a and b) at 123 K; and (c) at 77 K.

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3. Results 3.1. Prior austenite grain size and microhardness The microstructure of the AR condition consists of tempered bainite whereas that of the heat treated steels is tempered martensite. The AR sample exhibited a fine prior austenite grain size of 30 lm. The 1373 condition showed a prior austenite grain size of approx. 100 lm. The 1473 condition exhibited an extremely large, uniform equiaxed prior austenite grain size (diameter 1 mm), indicating unrestrained grain growth at high temperature, Fig. 4. This was attributed to the absence in the samples of steel of ‘microalloying’ elements such as Ti, Nb, V which are able to form carbides, nitrides and/or carbonitrides, which can pin grain boundaries at high temperatures (Table 1) [8]. The average and standard deviation (s.d.) Vickers microhardness

Fig. 6. SEM micrographs of the fracture surfaces from longitudinal tensile tests of the 1373 condition (a–c) at 123 K and (d–f) at 77 K.

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measurements using a load of 0.5 kg (HV 0.5), for the AR, the 1373 and 1473 conditions, were reported in Ref. [8]. The values obtained parallel to the S direction were 210 ± 6 for AR, 307 ± 7 for 1373 and 307 ± 7 for 1473. The values for the AR condition are comparable to the upper range reported in Ref. [2] for the SX4

Fig. 7. SEM micrographs of the fracture surfaces from longitudinal tensile tests of the 1473 condition (a and b) at 123 K and (c) at 77 K. Table 3 Fracture toughness, the non-dimensional deformation parameter M determined in three-point bending testing at 123 K and distance from precrack tip to the position of fracture initiation site Specimen

KQ (MPa m1/2)

M

Xf (lm)

AR SX25.3-a.2.7 SX25.3-a.1.6 SX25.3-a.1.7 SX25.3-a.1.1 SX25.3-a.1.3 SX25.3-a.2.6 SX25.3-a.2.3 SX25.3-a.2.1 SX25.3-a.1.8 SX25.3-a.1.4 SX25.3-a.2.2 SX25.3-a.2.4 SX25.3-a.1.5 SX25.3-a.1.2

47.4 50.4 51.9 52.5 55.2 56.3 57.6 59.0 59.8 61.5 62.7 64.9 66.2 72.4

446 367 365 353 287 298 300 293 272 256 243 225 219 176

19 100 32 85 30 77 11 47 35 25 45 33 10 61

Mean s.d. CoV

58.4 6.85 11.7

cp – close to precrack.

Specimen

KQ (MPa m1/2)

M

Xf (lm)

1373-QTQE SX25.3-b.1.6 SX25.3-b.1.8 SX25.3-b.2.6 SX25.3-b.1.10 SX25.3-b.2.8 SX25.3-b.2.4 SX25.3-b.2.1 SX25.3-b.1.2 SX25.3-b.2.5 SX25.3-b.1.7

66.8 84.6 84.8 88.8 90.9 97.8 98.9 105.8 109.3 115.6

265 148 159 136 143 121 116 103 99 90

cp 50 cp 49 38 cp 40 65 cp 48

Mean s.d. CoV

94.3 14.3 15.1

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block, 100 mm thick ‘homogeneous’ layer inside the wall of the ring segment as drawn in Fig. 1. The values for the 1373 and 1473 conditions are within the range reported in Ref. [15], 295 ± 25 for tempered martensitic microstructures in MnMoNi steels.

99.9 SE(B)-0.4T AR SE(B)-0.4T 1373

99 90

Pf (%)

70 50 30 10 1 0.1 0.01

0

10

20

30

40

50

60

70

80

90

100 110 120

1/2

KQ (MPa m

)

Fig. 8. Normal cdfs of fracture toughness values for the AR and 1373 conditions in precracked SE(B)-0.4T specimens at 123 K.

Fig. 9. SEM micrographs of the main cleavage initiation site in the precracked SE(B)-0.4T AR 56.3 MPa m1/2 at 123 K showing: (a) the cleavage initiation area ahead of the precrack tip and (b) cleavage facets around initiation site.

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3.2. Tensile tests Tensile test results obtained at 123 and 77 K for the AR, the 1373 and 1473 conditions are summarised in Table 2. The tensile test results for the AR and 1373 conditions were reported in Ref. [16]. On decreasing the testing temperature from 123 to 77 K for a given steel condition (AR, 1373 or 1473), the YS and UTS increased, whereas within the experimental scatter, both n and the El% values remained practically unchanged; the RA% values decreased slightly. For a given test temperature, comparing AR with the 1373 and 1473 conditions, the YS and UTS increased for the heat treated steel, whereas the n, El% and RA% values decreased. The 1373 and 1473 conditions exhibited similar strength levels (YS and UTS) and strain hardening exponent, however, the ductility of the 1473 condition (12.8% El and 25% RA at 123 K; 14.5% El and 15% RA at 77 K) is lower than that of the 1373 condition (15.9% El and 47% RA at 123 K; 17.1% El and 36% RA at 77 K). Associated with these differences in tensile ductility, differences in failure mode were observed as follows. Fractographic studies showed that on decreasing the testing temperature from 123 to 77 K, the failure mode changed from transgranular ductile/cleavage fracture to transgranular cleavage fracture with isolated ductile areas, for the AR condition (Fig. 5), and from intergranular/transgranular ductile/cleavage fracture to an increased area fraction of intergranular/transgranular cleavage fracture, for the 1373 condition (Fig. 6). The failure mode of the 1473 condition at 123 and 77 K was a mixture of intergranular and transgranular cleavage (Fig. 7). 3.3. Fracture toughness tests Values of fracture toughness and M are given in Table 3 and normal cumulative distribution functions (cdfs) are plotted in Fig. 8. Those results were reported in Ref. [16]. The mean fracture toughness

Fig. 10. SEM micrographs of the main cleavage initiation site in the precracked SE(B)-0.4T specimen 1373 condition 66.8 MPa m1/2 (min.) at 123 K showing: (a) the cleavage initiation area ahead of the precrack tip, (b) cleavage facets around initiation site and (c) detail of the fracture initiation site.

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increased from 58.4 ± 6.9 MPa m1/2 for AR bainitic steel to 94.3 ± 14.3 MPa m1/2 for the 1373 condition. These figures show that the means are significantly different at the 5% level of significance, assuming two normal populations having equal variance at the 1% level of significance. From Table 2 and Fig. 8, it is clear that the 1373 condition exhibits both higher YS and higher fracture toughness values than those for the AR condition. Figs. 9–11 show the fracture surfaces for selected samples, for the AR and 1373 conditions in precracked SE(B)-0.4T specimens fractured at 123 K. It can be seen that the fracture mode is predominantly transgranular with the 1373 samples showing large cleavage facets. The fractures have initiated from a microstructural feature at a critical distance (Xf) ahead of the precrack tip, or apparently at the tip itself for some 1373 specimens. For each specimen, the distance from the precrack tip to the fracture initiation site (Xf) as measured in the SEM is given in Table 3, together with the corresponding fracture toughness KQ value. 3.4. Blunt notch four-point bend tests 3.4.1. 0.2 mm blunt notch SE(B)-0.4T specimens The values of the local fracture stress (rf), fracture load (Lf) and nominal bending stress (rnom) are given in Tables 4 and 5. The rf values are plotted as normal cdfs in Fig. 12. The rf for the AR and 1373 conditions were reported in Ref. [16]. From Tables 4 and 5 and Fig. 12, it can be seen that, for a given condition, there is generally an increase in the values of rf and the coefficient of variation (CoV) at 77 K compared to those at 123 K. For a given testing temperature, the AR condition had lower rf values (1948–2041 MPa, at 123 K), (2017–2192 MPa, at 77 K), than those for the 1373 and 1473

Fig. 11. SEM micrographs of the main cleavage initiation site in the precracked SE(B)-0.4T specimen 1373 condition 88.8 MPa m1/2 at 123 K showing: (a) the cleavage initiation area ahead of the precrack tip, (b) cleavage facets around initiation site and (c) detail of the fracture initiation site.

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conditions. At 123 K, the 1373 condition generally exhibited higher local fracture stresses values (2591– 2641 MPa) than the 1473 condition (2106–2595 MPa), although there was a small overlap for the lowest value 1373 value. At 77 K, the range of rf values in both conditions tend to overlap (2364–3014 MPa for the 1373 condition and 2265–2968 MPa for the 1473 condition), but the lower bound value for the 1473 condition was 99 MPa (4%) lower than that for 1373 condition.

Table 4 Macro- and micro-mechanical parameters of 0.2 mm blunt notch four-point bending tests at 123 K Specimen

Lf (kN)

rnom (MPa)

AR SX20.1-c.1.5 SX20.1-c.2.5 SX20.1-c.1.6 SX20.1-c.2.6 SX20.1-c.1.7 SX20.1-c.2.7 SX20.1-c.1.8 SX20.1-c.2.8

23.08 23.81 26.69 25.74 25.89 23.51 23.08 22.93

1556 1606 1800 1736 1746 1585 1556 1546

Min Max Mean s.d. CoV

22.93 26.69 24.34 1.51 6.21

1373 SX25.2-b.1.1 SX25.2-b.1.2 SX25.2-b.1.3 SX25.2-b.1.4 SX25.2-b.1.5 SX25.2-b.1.6 SX25.2-b.1.7 SX25.2-b.1.8 SX25.2-b.1.9 SX25.2-b.1.10

rnom/ry

Xf (lm)

Xf/q

r1max/rys

rf (MPa)

2.00 2.06 2.31 2.23 2.24 2.04 2.00 1.98

147 205 256 253 287 221 138 677

0.74 1.03 1.28 1.27 1.44 1.11 0.69 3.39

2.50 2.51 2.62 2.59 2.60 2.53 2.50 2.50

1948 1955 2041 2018 2025 1971 1948 1948

1546 1800 1641 102 6.21

1.98 2.31 2.11 0.13 6.21

138 677 273

0.69 3.39 1.37

2.50 2.62 2.54 0.05 2.00

1948 2041 1982 40 2.00

34.59 33.59 33.14 34.55 33.70 33.20 33.33 34.33 33.78 32.26

2332 2265 2234 2330 2272 2238 2248 2315 2278 2175

2.30 2.24 2.21 2.30 2.25 2.21 2.22 2.29 2.25 2.15

2.61 2.60 2.58 2.61 2.60 2.58 2.58 2.61 2.60 2.56

2641 2631 2611 2641 2631 2611 2611 2641 2631 2591

Min Max Mean s.d. CoV

32.26 34.59 33.65 0.72 2.15

2175 2332 2269 49 2.15

2.15 2.30 2.24 0.05 2.15

2.56 2.61 2.59 0.02 0.66

2591 2641 2624 17 0.66

1473 SX25.2-a.2.1 SX25.2-a.2.2 SX25.2-a.2.3 SX25.2-a.2.4 SX25.2-a.2.5 SX25.2-a.2.6 SX25.2-a.2.7 SX25.2-a.2.8 SX25.2-a.2.9 SX25.2-a.2.10

32.27 33.13 30.74 31.18 27.48 32.15 30.53 32.97 19.23 32.54

2176 2234 2073 2103 1853 2168 2059 2223 1297 2194

2.18 2.24 2.08 2.11 1.86 2.17 2.06 2.23 1.30 2.20

2.57 2.60 2.54 2.55 2.43 2.57 2.51 2.59 2.11 2.58

2565 2595 2535 2545 2425 2565 2505 2585 2106 2575

Min Max Mean s.d. CoV

19.23 33.13 30.22 4.20 13.9

1297 2234 2038 283 13.9

1.30 2.24 2.04 0.28 13.9

2.11 2.60 2.51 0.15 5.88

2106 2595 2500 147 5.88

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Fig. 13 shows a fracture surface at 77 K for an AR sample, from which it can be seen that, the fracture mode is predominantly transgranular and that the fractures have initiated from microstructural features at a critical distance. The critical distance measurements (Xf) are given in Tables 4 and 5 and the range of values are marked in a copy of the original Griffiths and Owen’s stress and strain distributions ahead of the blunt

Table 5 Macro- and micro-mechanical parameters of 0.2 mm blunt notch four-point bending tests at 77 K Specimen

Lf (kN)

rnom (MPa)

AR SX20.1-c.1.1 SX20.1-c.2.1 SX20.1-c.1.2 SX20.1-c.2.2 SX20.1-c.1.3 SX20.1-c.2.3 SX20.1-c.1.4 SX20.1-c.2.4

24.03 23.80 24.14 23.65 19.54 21.47 20.82 23.10

1620 1605 1628 1595 1318 1448 1404 1558

Min Max Mean s.d. CoV

19.54 24.14 22.57 1.73 7.68

1373 SX20.1-a.1.1 SX20.1-a.2.1 SX20.1-a.1.2 SX20.1-a.2.2 SX20.1-a.1.3 SX20.1-a.2.3 SX20.1-a.1.4 SX20.1-a.2.4

rnom/ry

Xf (lm)

Xf/q

r1max/ry

rf (MPa)

1.77 1.75 1.78 1.74 1.44 1.58 1.53 1.70

356 113 149 150 73 300 123 170

1.78 0.57 0.75 0.75 0.37 1.50 0.62 0.85

2.39 2.38 2.39 2.37 2.20 2.28 2.25 2.35

2192 2182 2192 2173 2017 2091 2063 2155

1318 1628 1522 117 7.68

1.44 1.78 1.66 0.13 7.68

73 356 179

0.37 1.78 0.90

2.20 2.39 2.33 0.07 3.14

2017 2192 2133 67 3.14

30.95 33.12 35.19 26.00 19.96 27.54 32.82 37.18

2087 2233 2373 1753 1346 1857 2213 2507

1.77 1.89 2.00 1.48 1.14 1.57 1.87 2.12

2.37 2.45 2.50 2.22 2.00 2.27 2.44 2.55

2801 2896 2955 2624 2364 2683 2884 3014

Min Max Mean s.d. CoV

19.96 37.18 30.35 5.59 18.4

1346 2507 2046 377 18.4

1.14 2.12 1.73 0.32 18.4

2.00 2.55 2.35 0.18 7.66

2364 3014 2778 213 7.66

1473 SX20.1-b.1.1 SX20.1-b.2.1 SX20.1-b.1.2 SX20.1-b.2.2 SX20.1-b.1.3 SX20.1-b.2.3 SX20.1-b.1.4 SX20.1-b.2.4 SX20.1-b.1.5 SX20.1-b.2.5

33.02 24.20 26.82 22.84 25.85 29.37 32.88 35.02 24.97 17.37

2227 1632 1809 1540 1743 1980 2217 2361 1684 1171

1.87 1.37 1.52 1.29 1.46 1.66 1.86 1.98 1.41 0.98

2.44 2.15 2.25 2.11 2.21 2.32 2.43 2.49 2.17 1.90

2908 2563 2682 2515 2634 2765 2897 2968 2587 2265

Min Max Mean s.d. CoV

17.37 35.02 27.23 5.41 19.9

1171 2361 1836 365 19.9

0.98 1.98 1.54 0.31 19.9

1.90 2.49 2.25 0.18 8.00

2265 2968 2678 214 8.00

Lf – fracture load; rnom – nominal bending stress 6M/B(W  a)2, where M is the bending moment M = Pf/2 · B; ry – yield strength; Xf – distance from blunt notch root to the position of fracture initiation site; q – blunt notch root radius; rf – local fracture stress; s.d. – standard deviation; CoV – coefficient of variation = s.d./mean · 100.

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notch, Fig. 14 [13]. In contrast, Figs. 15–21 show selected fracture surfaces at 123 and 77 K for the 1373 and 1473 conditions. These exhibit intergranular/transgranular cleavage facets in the region near the notch root, except for the lowest value of the data set 1473 condition 2106 MPa (Fig. 17) where cleavage initiated at matrix microstructural features in an extremely coarse cleavage facet. The crack propagation in the bulk of the fracture surface away from the notch root was transgranular cleavage with a few isolated intergranular cleavage facets at 123 K (Figs. 15–18) and exhibited an increased proportion of intergranular facets at 77 K (Figs. 19–21). 3.4.2. 1 mm blunt notch SE(B)-0.5T specimens The values of the local fracture stress (rf), fracture load (Lf) and nominal bending stress (rnom) are given in Table 6 and rf values are shown in Fig. 12. On increasing the notch root radius from 0.2 to 1 mm, for a given heat treatment (1373 or 1473), the corresponding lower bound value of the rf range at 77 K increased and the upper bound value decreased. While the scatter band narrowed in range, the mean rf values remained relatively constant in the range between 2622 and 2778 MPa. Again at 77 K, the range of rf values in both steels tend to overlap (2660–2719 MPa for the 1373 condition and 2473– 2742 MPa for the 1473 condition), but there was a decrease of 187 MPa (7%) in the lower bound value for the 1473 condition. Figs. 22 and 23 show the fracture surfaces for selected samples in 1 mm blunt notch SE(B)-0.5T specimens at 77 K for the 1373 and 1473 conditions. It can be seen that the failure mode is intergranular/transgranular cleavage fracture in the region near the notch root and predominantly transgranular cleavage crack propagation as observed in 0.2 mm blunt notch SE(B)-0.4T specimens for this steel condition at the testing temperatures of 123 and 77 K (Fig. 15–21).

99.9 99 90

Pf (%)

70 50 30 10 1 0.1 0.01 1800

2000

2200

2400

2600

2800

3000

3200

Local fracture stress (MPa) SE(B)-0.4T SE(B)-0.4T SE(B)-0.4T SE(B)-0.4T SE(B)-0.4T SE(B)-0.4T SE(B)-0.4T SE(B)-0.4T

0.2 mm notch radius AR 123 K 0.2 mm notch radius AR 77 K 0.2 mm notch radius 1373 123 K 0.2 mm notch radius 1373 77 K 0.2 mm notch radius 1473 123 K 0.2 mm notch radius 1473 77 K 1 mm notch radius 1373 77 K 1 mm notch radius 1473 77 K

Fig. 12. Normal cdf of local fracture stresses values for the blunt notched SE(B)-0.4T specimens at 123 K – filled symbols – and at 77 K – unfilled symbols –, showing linear and non-linear behaviour.

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4. Discussion 4.1. Prior austenite grain size The AR sample exhibited a fine prior austenite grain size of 30 lm, but a relatively coarse bainitic microstructure. The 1373 K treatment gave an average prior austenite grain size of approximately 100 lm, such as those encountered in simulated CGHAZ of a SA 533 weld [17]. The 1473 K treatment produced extremely large austenite grains (>1 mm diameter) which would not generally be thought to be representative of a CGHAZ in practice, although a prior austenite grain size of 410 lm was measured at a peak temperature of 1623 K after double thermal cycling an ASME SA 508 steel to simulate an unaltered CG HAZ [18] The 1473 heat treatment was investigated primarily to elucidate the fracture behaviour of an extremely coarse prior austenite grain size in the steel.

4.2. Tensile tests For the AR condition, the average values of YS at 123 and 77 K are 779 and 917 MPa. These may be compared with the values derived from the temperature dependence of the YS (Eq. 5) obtained by Wallin [1] for the DIN 22NiMoCr37 steel and the values given in Ref. [19] for other steels DIN 22NiMoCr37 and SA 508. YS ¼ 450 þ 1294 expð0:0147T Þ

ð5Þ

where YS is in MPa, T is in K and 119 K < T < 293 K. In the work from which Eq. (5) was derived, specimens were loaded in the L direction at a cross-head displacement rate of 0.5 mm/min [2], as in the present work. The

Fig. 13. SEM micrographs of the main cleavage initiation site in the 0.2 mm blunt notch SE(B)-0.4T specimen AR at 77 K: (a) Xf = 73 lm (min.) and rf = 2017 MPa (min.) and (b) Xf = 356 lm (max.) and rf = 2192 MPa (max.).

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YS value at 123 K obtained from Eq. (5) is 662 MPa. The YS values at 123 and 77 K given in Ref. [19] are 740 and 902 MPa, respectively, in agreement with the present values. The values of UTS (866 MPa at 123 K and 983 MPa at 77 K) are similar to the values of 807 ± 18 MPa at 119 K [2] and 840 MPa at 123 K, 951 MPa at 77 K [19]. The decrease in RA values for the AR condition with decreasing test temperature from 123 to 77 K, was associated with a change from transgranular ductile/cleavage fracture to transgranular cleavage fracture with isolated ductile areas (Fig. 5), although the El values remained practically unchanged. El values for SA 508 steel are 20.9% at 133 K, 12.6% at 77 K [20]. The

Fig. 14. The variation of (a) the maximum principal stress and (b) the maximum strain distribution, below the notch root at various loads [13]. The region between the continuous lines corresponds to Xf/q = 0.69 and 1.44, rnom/ry = 2.00 and 2.24 (left position of maximum principal stress) and the point Xf/q = 3.39 and rnom/ry = 1.98 (right position of maximum principal stress), for the AR steel at 123 K. The region between the dashed lines corresponds to Xf/q = 0.37 and 1.78, rnom/ry = 1.44 and 1.77 (left position of maximum principal stress), for the AR steel at 77 K.

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RA values reported were in the range 68–71% between 293 and 119 K [2]. The n values reported were in the range 0.15–0.13 between 198 and 77 K [20]. It was found that prior austenite grain size per se in the tempered martensites has no effect on the yield strength of the steel. At 123 K, the mean values were 1012 MPa for the 1373 K (100 lm grain size) and 998 MPa for the 1473 K (1 mm grain size). At 77 K, the values were 1182 and 1192 MPa, respectively. These values are identical, within experimental scatter, and values for the UTS similarly did not vary with grain size. The reason for this is that resistance to dislocation movement is provided by the fine scale tempered carbide distribution: dislocation tangle around the particles and no simple ‘pile-up’ arrays traversing the length of a grain are created. In this very clean steel, having no significant level of microalloying elements, all carbon is taken into solution at both 1373 and 1473 K, so that the volume fraction, size and spacing of carbides particles is governed solely by the tempering conditions, since this is the same for both austenitising treatments, the yield strength are identical (and greater than that for the coarser tempered bainitic condition). The reduced amount of grain boundary area per unit volume for the coarse grain material, however, made it more susceptible to intergranular embrittlement resulting from the 180 h at 793 K exposure. The RA at 123 K was 47% for the 1373 condition and 25% for the 1473 condition. At 77 K, the values were 36% and 15%, respectively. A decrease in RA values with decreasing test temperature from 123 to 77 K can also be observed for the 1373 and 1473 conditions. However, the failure mode for the 1373 condition changed from intergranular/ transgranular ductile/cleavage fracture to an increased area of intergranular/transgranular cleavage fracture (Fig. 6). The lower ductility values of the 1473 condition compared with those of the 1373 condition were associated with the occurrence of coarse intergranular/transgranular cleavage facets (1473 condition) compared with coarse intergranular/transgranular ductile/cleavage facets (1373 condition), as shown in Figs. 6 and 7.

4.3. Fracture toughness tests Extensive fracture toughness datasets in Charpy size precracked specimens (PCH-dataset) [21] and in 12.5, 25, 50 and 100 mm thickness compact tension specimens (Euro dataset) [1,2], are available for the AR (tempered bainite) microstructure. The PCH-dataset results at 119 K of Ref. [21] for the AR condition and the values of sharp crack fracture toughness determined at 123 K, for the AR and 1373 conditions, plotted as

Fig. 15. SEM micrographs of the main cleavage initiation site in the 0.2 mm blunt notch SE(B)-0.4T specimen 1373 condition at 123 K, rf = 2591 MPa (min.) showing (a) the cleavage initiation area ahead of the notch, (b) a mixture of intergranular/transgranular cleavage fracture in the region near the notch root, and (c) predominantly transgranular cleavage crack propagation in the bulk of the fracture surface marked in (a). A few isolated intergranular cleavage facets have been marked in (c).

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normal cdf, are compared in the Appendix and are represented in Fig. 24. Effects of geometry and constraint on fracture toughness are also discussed in the Appendix. The 1373 martensitic condition exhibited improved fracture toughness values at higher strength levels compared to the AR bainitic steel. In an A533B RPV steel, both fine and coarse-grained martensite microstructures

Fig. 15 (continued)

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Fig. 15 (continued)

exhibited superior combinations of yield strength and fracture toughness compared to those of coarse-grained bainitic microstructures [3]. The coarsest carbides in the distribution were found to be responsible for the most deleterious effects on fracture toughness. Higher combinations of YS and toughness values, for a

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coarse-grained martensitic microstructure compared with a fine-grained bainitic microstructure, have been reported for ASME SA 508 steel after postweld heat treatment (PWHT) [18] and for an A533B steel plate [22,23]. The increase in size of carbides in the bainitic microstructure was found to be the most important microstructural factor affecting Charpy toughness after PWHT. In an A533B steel plate [22,23], the fine-grained bainitic condition referred to the microstructure obtained after oil quenching to room temperature from 1 h at 1223 K and the coarse-grained martensitic condition referred to the microstructure obtained after oil quenching to room temperature from 1 h at 1523 K. The YS values determined at 193 K were 700 and 1400 MPa, respectively. The corresponding KIc values at 193 K, for the fine-grained bainitic steel, ranged from a minimum of 39.8 MPa m1/2 to a maximum of 50.3 MPa m1/2, mean 45.8 MPa m1/2 and s.d. of ±3.7 MPa m1/2. For the coarse-grained martensitic steel, they ranged from a minimum of 81.5 MPa m1/2 to a maximum of 101.6 MPa m1/2, mean 89.7 MPa m1/2 and s.d. of ±6.2 MPa m1/2. The mean fracture toughness increased from 45.8 ± 3.7 MPa m1/2 for fine-grained bainitic steel to 89.7 ± 6.2 MPa m1/2 for coarse-grained martensitic steel. These figures indicate a difference in means at the 5% level of significance, assuming two normal populations having equal variance at the 1% level of significance. Nevertheless, in a recent work on the effect of phosphorus segregation on fracture toughness in 2.25Cr– 1Mo pressure vessel steel having 0.013 wt.% P content and a tempered martensite microstructure [6], a direct relationship between grain boundary phosphorus concentration and fracture toughness at 77 K in precracked specimens has been demonstrated. This was associated with a transition from entirely transgranular cleavage to mixed intergranular/transgranular cleavage fracture. The decrease in fracture toughness was related to an increase in the area fraction of intergranular cleavage fracture [6]. It was found that P20% area fraction of intergranular fracture was needed to produce a significant reduction in fracture toughness. Near-tip stress fields in a SA 508 steel (class 3) having a tempered bainite microstructure have been calculated for the mid-section of precracked Charpy V-notch specimens [19]. The YS values at 133 and 77 K, for the SA 508 steel (class 3), given in Ref. [19] are 662 and 902 MPa, respectively. The computed results presented in Fig. 3 of Ref. [19] showed that the critical distance for the peak value of the maximum principal stress ahead of the precrack tip for KJc = 71 MPa m1/2 at 133 K is 38 lm and for KJc = 36.4 MPa m1/2 at 77 K is 9.9 lm, which compare reasonably (within scatter) with the results for Xf in the present investigation (Table 3).

Fig. 16. SEM micrographs of the main cleavage initiation site in the 0.2 mm blunt notch SE(B)-0.4T specimen 1373 condition at 123 K, rf = 2641 MPa (max.) showing (a) the cleavage initiation area ahead of the notch, (b) a mixture of intergranular/transgranular cleavage fracture in the region near the notch root; and (c) predominantly transgranular cleavage crack propagation in the bulk of the fracture surface marked in (a). A few isolated intergranular cleavage facets have been marked in (c).

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The failure mode in precracked SE(B)-0.4T specimens for the 1373 condition at 123 K is transgranular cleavage, contrasted with that in blunt notch SE(B)-0.4T specimens as discussed below.

4.4. Blunt notch four-point bend tests The microscopic cleavage fracture strength was found to depend on the steel microstructure in an A533B RPV steel [3]. Transgranular cleavage fracture strength values at 77 K were reported to be 1900 and 1950 MPa for coarse-grained (110 and 130 lm, respectively) bainitic microstructures and were found to lie in the range 3100–3800 MPa for both fine (8 lm) and coarse (120–130 lm) grained martensitic microstructures. The YS values determined at 77 K, for the bainitic microstructure were 1000 MPa (110 lm grain size) and 1030 MPa (130 lm grain size) and for the martensitic microstructures 1480 and 1570 MPa (8 lm grain size) and 1450 MPa (120 lm grain size) and 1660 MPa (130 lm grain size). At the test temperature of 77 K in the present investigation, transgranular cleavage fracture strength values for fine-grained bainitic microstructures were in the range of 2017–2192 MPa. Intergranular/transgranular cleavage fracture strength values for coarse-grained martensitic microstructures were in the range 2364–3014 MPa for the 1373 condition and in the range 2265–2968 MPa for the 1473 condition. The YS values determined at 77 K, for the bainitic

Fig. 16 (continued)

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Fig. 16 (continued)

microstructure were 917 MPa (30 lm grain size) and for the coarse-grained martensitic microstructure 1182 MPa (100 lm grain size) for the 1373 condition and 1192 MPa (>1 mm grain size) for the 1473 condition. On previous studies in martensitic steels, on the effect of phosphorus segregation on cleavage fracture stress in a model steel with a chemical composition representative of segregated zones of a RPV steel A508 class 3 [5]; and in a A533B steel [6], a direct relationship between grain boundary phosphorus concentration and microscopic fracture stress at 77 K has been demonstrated. The decrease in fracture stress is associated with a transition from 100% transgranular cleavage to partly intergranular cleavage fracture. It has been concluded that P20% area fraction of intergranular fracture is needed to produce a significant reduction in the microscopic fracture stress, as for the fracture toughness variation. Transgranular cleavage fracture strength values at 77 K were found to be 2392 ± 198 MPa for a martensitic A533B steel which was austenitised at 1473 K for 1 h (a prior austenite grain size of 110 lm), water quenched and then tempered at 923 K for 2 h followed by water quenching [6]. The A533B steel had a phosphorus content in the base composition of 0.005 wt.% similar to that in the present investigation (Table 1). If tempering was followed by embrittlement for 180 h at 793 K (as for the DIN 22NiMoCr37 steel), the intergranular/transgranular cleavage fracture strength decreased to 2192 ± 128 MPa, compared with the value of 2778 ± 213 MPa at 77 K (see Fig. 12). This difference cannot be explained in terms of grain size or phosphorus content. The carbon content in the A533B steel [6] was

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0.24 compared with 0.21 in the present steel, but this would not be expected to have a large effect. It is possible that the A533B steel contained other impurity elements such as Sn or Sb which were not analysed. In Fig. 14, it can be seen that transgranular cleavage fracture initiated at a critical distance from the notch root, with a bias of distances at positions between the notch root and the peak stress. This suggest that, for the AR condition, the critical event of cleavage fracture at 123 and 77 K in 0.2 mm blunt notch SE(B)-0.4T specimens, requires a combination of the attainment of a critical strain level and a critical stress level, i.e. cleavage

Fig. 17. SEM micrographs of the main cleavage initiation site in the 0.2 mm blunt notch SE(B)-0.4T specimen 1473 condition at 123 K, rf = 2106 MPa (min.) showing (a) the cleavage initiation area ahead of the notch and (b) detail of fracture initiation site.

Fig. 18. SEM micrograph of the main cleavage initiation site in the 0.2 mm blunt notch SE(B)-0.4T specimen 1473 condition at 123 K, rf = 2595 MPa (max.), showing a mixture of intergranular/transgranular cleavage fracture in the region near the notch root and predominantly transgranular cleavage crack propagation in the bulk of the fracture surface away from the notch root.

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fracture occurred under mixed nucleation/propagation control. For the 1373 and 1473 conditions, in 0.2 mm blunt notch SE(B)-0.4T specimens, the failure mode in the region near the notch root was intergranular/transgranular cleavage (15–30% intergranular fracture at 123 K and 20–40% intergranular fracture at 77 K) except

Fig. 19. SEM micrographs of the main cleavage initiation site in the 0.2 mm blunt notch SE(B)-0.4T specimen 1373 at 77 K (a) rf = 2364 MPa (min.), showing (a) a mixture of intergranular/transgranular cleavage fracture in the region near the notch root; and (b) predominantly transgranular cleavage crack propagation in the bulk of the fracture surface. Some isolated intergranular cleavage facets have been marked in (b).

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Fig. 19 (continued)

for the lowest value of the dataset 1473 condition 2106 MPa (Fig. 17) where cleavage initiated at matrix microstructural features in an extremely coarse cleavage facet. The crack propagation in the bulk of the fracture surface away from the notch root was transgranular cleavage with a few isolated intergranular cleavage facets (1–5%) at 123 K (Figs. 15–18) and exhibited an increased proportion of intergranular facets (8–10%) at 77 K (Figs. 19–21). The observation that intergranular cleavage fracture was predominant in the region near the notch root, but transgranular cleavage fracture was predominant in the bulk of the fracture surface away from the notch root is consistent with that of Nanstad et al. [24]. This related to a (moderately) low P (0.007 wt.%) modified A302 grade B steel given the following heat treatment in a Gleeble simulation: austenitised at 1533 K for 10 s to produce an ASTM grain size in the range of 4–5 (equivalent to a feret diameter of 101 and 71 lm, respectively, as tabulated in ASTM Standard E112), followed by PWHT at 888 K for 24 h, oil quenched, aging at 723 K for 168 h and tested in Charpy V-notch impact, and then examined fractographically in an SEM. It is however, of interest to note that the lowest value of the dataset found for the 1473 condition suggests that in this steel, which has very low levels of trace impurity elements, a very coarse cleavage facet, when sampled ahead of a stress concentration, can have a significant effect in decreasing the fracture stress, and therefore, can be more detrimental to toughness. Note that, for the 1373 condition in precracked SE(B)-0.4T specimens, cleavage initiated at matrix microstructure features, whereas in 0.2 and 1 mm blunt notch SE(B)-0.4T and SE(B)-0.5T specimens,

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respectively, cleavage initiated at intergranular/transgranular cleavage facets. This different fracture behaviour may be attributed to sampling effects in the ‘process zone’ ahead of precracked and blunt notch specimens. In agreement with the results of Ritchie et al. [9] it appears that it is more difficult for a sharp crack to sample an embrittled grain boundary than it is for a blunt notch. It is possible that the reversed slip associated with prefatiguing partially disrupts the continuity of boundaries, but high proportions of intergranular facets have been seen on A533B tested at 77 K [6]. Druce [15] reported a cleavage initiation mechanism beneath the notch root in high strain rate Charpy impact testpieces to be at high angle boundaries, possibly prior austenite grain boundaries, when the cleavage initiation site could be identified, despite fractography showing brittle intergranular cleavage facets. An initiation site example is shown in Fig. 15 of Ref. [15]. Because of the significance of the sampling effects, the 1373 and 1473 conditions were tested in notched bars with a larger notch root radius (1 mm) and larger notch angle (90), Fig. 2b [14]. From Tables 5 and 6, it can be seen that for given macro- or micro-scopic parameters and steel condition (1373 or 1473), when comparing the 1 mm notch root radius observations with those made 0.2 mm specimens, the CoV

Fig. 20. SEM micrographs of the main cleavage initiation site in the 0.2 mm blunt notch SE(B)-0.4T specimen 1373 at 77 K (a) rf = 3014 MPa (max.), (a) a mixture of intergranular/transgranular cleavage fracture in the region near the notch root and (b) predominantly transgranular cleavage crack propagation in the bulk of the fracture surface. Some isolated intergranular cleavage facets have been marked in (b).

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Fig. 20 (continued)

Fig. 21. SEM micrographs of the main cleavage initiation site in the 0.2 mm blunt notch SE(B)-0.4T specimen 1473 condition at 77 K (a) rf = 2265 MPa (min.) and (b) rf = 2968 MPa (max.), showing a mixture of intergranular/transgranular cleavage fracture in the region near the notch root.

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Fig. 22. SEM micrographs of the main cleavage initiation site in the 1 mm blunt notch SE(B)-0.5T specimen 1373 condition at 77 K (a) rf = 2689 MPa and (b) rf = 2719 MPa (max.), showing a mixture of intergranular/transgranular cleavage fracture in the region near the notch root.

decreased and the scatter bands of the rf values narrowed in range; however, the mean rf values remained relatively constant. SEM micrographs shown in Figs. 22 and 23 confirmed that very coarse intergranular/transgranular cleavage facets, when sampled ahead of a stress concentration, are associated with a decrease in local fracture stress Fig. 23a. As observed previously for 0.2 mm SE(B)-0.4T specimens at 77 K, the fracture initiated from intergranular/transgranular facets (20–40% intergranular fracture) and the crack propagation in the bulk of the fracture surface away from the notch root was predominantly transgranular cleavage (8–10% intergranular fracture). The area fraction of intergranular facets in the region close to the notch root has increased in the 1 mm SE(B)-0.5T specimen, as can be compared in Figs. 19 and 22, for the 1373 condition, and in Figs. 21 and 23, for the 1473 condition. It has been reported for as-quenched martensitic AISI 4340 steel that when the steel had a fine-grained microstructure (diameter 24–32 lm) it exhibited a lower sharp crack fracture toughness, KIc, than when the steel had a coarse-grained microstructure (diameter 254–360 lm) [9]. However, the fine-grained microstructure showed superior toughness for fracture ahead of blunt notches once the notch root radius exceeded a critical value (q > 0.05 mm). The sharp crack behaviour was attributed to be due to a larger ‘effective’ root radius for fracture ahead of sharp cracks in the coarse-grained structure. The blunt notch effect was attributed to be due to a higher critical fracture stress for failure in the fine-grained microstructure. In the present experiments, the coarse (100 lm) grained martensitic 1373 condition, even though subjected to a temper embrittling treatment, exhibited improved local fracture stress values at 123 and 77 K (Fig. 12), and generally improved fracture toughness values at 123 K (Fig. 8), compared to the fine (30 lm) grained bainitic AR condition. There is no significant difference between the 1373 and 1473 rf values for 0.2 and 1 mm root radius notches.

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Fig. 23. SEM micrographs of the main cleavage initiation site in the 1 mm blunt notch SE(B)-0.5T specimen 1473 condition at 77 K (a) rf = 2473 MPa (min.) showing very coarse cleavage facets (arrowed) sampled ahead of the notch root and (b) rf = 2742 MPa (max.), showing a mixture of intergranular/transgranular cleavage fracture in the region near the notch root.

99.9 99 90

Pf (%)

70 50 30 10 1 0.1 0.01 0

10

20

30

40

50

60

1/2

KJc, KQ (MPa m

)

SE(B)-0.4T-AR exp KJc at 119 K Ref. [21] SE(B)-0.4T-AR exp KQ at 123 K C(T)-1T-AR exp KJc at 119 K Ref. [2] C(T)-1T-AR pred KJc at 119 K (Kmin = 0) Fig. 24. Normal cdfs of fracture toughness values.

70

80

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Table 6 Macro- and micro-mechanical parameters of 1 mm blunt notch four-point bending tests, at 77 K Lf (kN)

rnom (MPa)

rnom/ry

r1max/ry

rf (MPa)

62.46 59.08 61.82 60.56 57.84 66.66

3291 3112 3257 3190 3047 3301

2.76 2.61 2.73 2.68 2.56 2.77

2.30 2.25 2.30 2.28 2.25 2.30

2719 2660 2719 2689 2660 2719

Min Max Mean s.d. CoV

57.84 62.66 60.74 1.95 3.21

3047 3301 3200 103 3.21

2.56 2.77 2.68 0.09 3.21

2.25 2.30 2.28 0.02 1.08

2660 2719 2694 29.1 1.08

1473 SX25.2-b.2 0 + 1 0 .1 SX25.2-b.2 0 + 1 0 .2 SX25.2-b.2 0 + 1 0 .3 SX25.2-b.2 0 + 1 0 .4 SX25.2-b.2 0 + 1 0 .5 SX25.2-b.2 0 + 1 0 .6 SX25.2-b.2 0 + 1 0 .7 SX25.2-b.2 0 + 1 0 .8 SX25.2-c.2 0 + 1 0 .1 SX25.2-c.2 0 + 1 0 .2

52.28 60.43 57.94 51.23 48.39 59.97 64.19 57.29 57.34 52.96

2754 3183 3053 2699 2550 3160 3382 3018 3021 2790

2.33 2.69 2.58 2.28 2.16 2.67 2.86 2.55 2.56 2.36

2.14 2.28 2.25 2.14 2.08 2.28 2.30 2.20 2.20 2.14

2551 2712 2682 2551 2473 2712 2742 2622 2622 2551

Min Max Mean s.d. CoV

48.39 64.19 56.20 4.86 8.65

2550 3382 2961 256 8.65

2.16 2.86 2.51 0.22 8.65

2.08 2.30 2.20 0.07 3.39

2473 2742 2622 89 3.39

1373 SX25.2-a.2 0 SX25.2-a.2 0 SX25.2-a.2 0 SX25.2-a.2 0 SX25.2-a.2 0 SX25.2-a.2 0

+ 1 0 .1 + 1 0 .2 + 1 0 .3 + 1 0 .4 + 1 0 .5 + 1 0 .6

5. Conclusions The low temperature mechanical properties of DIN 22NiMoCr37 steel weldment in simulated ‘best estimate’ (1373 condition) and ‘extra large’ (1473 condition) CGHAZ microstructures were determined and compared to those of the parent steel: 1. The martensitic 1373 condition exhibited higher values of yield strength, fracture toughness and local fracture strength than the bainitic AR condition at 123 K. The mean YS values for 1373 condition were 1012 MPa at 123 K and 1182 MPa at 77 K compared with 779 MPa at 123 K and 917 MPa at 77 K for the AR condition. The mean fracture toughness at 123 K was 94.3 ± 14.3 MPa m1/2 for the 1373 condition compared with 58.4 ± 6.9 for AR bainitic steel. The mean rf values for the 1373 condition were 2624 MPa at 123 K and 2778 MPa at 77 K compared with 1982 MPa at 123 K and 2133 MPa at 77 K for the AR condition (i.e. 30% higher). 2. The 1373 condition, generally, exhibited higher local fracture stresses values at 123 K (2591–2641 MPa) than the 1473 condition (2106–2595 MPa), both steels having similar strength levels. At 77 K, the range of rf values in both conditions tend to overlap (2364–3014 MPa for the 1373 condition and 2265–2968 MPa for the 1473 condition), but there was a decrease in the lower bound value for the 1473 condition. 3. On increasing the notch root radius from 0.2 to 1 mm, the corresponding lower bound value of the rf range at 77 K increased and the upper bound value decreased, but the mean values were similar. Again, the range of rf values in both conditions tend to overlap (2660–2719 MPa for the 1373 condition at 77 K and 2473–2742 MPa for the 1473 condition at 77 K), but there was a decrease in the lower bound value for the 1473 condition.

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4. For the 1373 and 1473 conditions, both in 0.2 mm blunt notch SE(B)-0.4T and in 1 mm blunt notch SE(B)0.5T specimens, in general, the failure mode was intergranular/transgranular cleavage in the near notch region (15–30% intergranular fracture at 123 K and 20–40% intergranular fracture at 77 K), but transgranular cleavage fracture was predominant in the bulk of the fracture surface away from the notch root (1–5% intergranular fracture at 123 K and 8–10% intergranular fracture at 77 K). In contrast, for the 1373 condition, the failure mode was transgranular cleavage in precracked SE(B)-0.4T specimens. 5. When a very coarse intergranular/transgranular cleavage facet is sampled ahead of a stress concentration, it can have a significant effect in decreasing the cleavage fracture stress. This behaviour has also been observed in the 1 mm notch root radius specimens. Acknowledgements This work was carried out as part of a joint UK EPSRC-funded research project entitled ‘Role of interfaces in fracture resistance of power generation materials: combined multiscale modelling/experimental approach’ between Birmingham, Bristol, Surrey and Swansea Universities, UK. Thanks are due to BNFL Magnox Generation for its support of the research and provision of C(T)-4T steel specimens used to generate the Euro fracture toughness dataset, to British Energy for permission to reproduce the PCH-dataset of the Appendix and to the Department of Metallurgy and Materials at the University of Birmingham for the provision of research facilities. Appendix. Material variability and effects of geometry and constraint on the fracture toughness of the DIN 22NiMoCr37 reactor pressure vessel steel in the lower shelf region From our results at 123 K, the KQ values range from a minimum of 47.4 MPa m1/2 to a maximum of 72.4 MPa m1/2, mean 58.4 MPa m1/2 and s.d. of ±6.85. In Fig. 24, we show fracture toughness results converted from Jc values obtained on precracked Charpy samples at 119 K [21]. The Jc values range from a minimum of 3.7 N/mm (29.8 MPa m1/2) to a maximum of 22 N/mm (72.5 MPa m1/2), mean 9.7 N/mm (48.4 MPa m1/2) and s.d. of ±9.50, in 50 specimens tested. The mean fracture toughness is observed to be 48.4 ± 9.50 MPa m1/2 rather than 58.4 ± 6.85 MPa m1/2 for our results. These means are significant different at the 5% level of significance, assuming two normal populations having equal variance at the 5% level of significance. This difference in mean values for only 4 K difference in testing temperature, may be attributed in part to inherent spatial variability in material properties. This has been reported previously for the large ring segment of the AR steel blocks SX9 and SX10 tested in SE(B)-0.4T at 213 K [2]. Although, no significant material variability was found for the steel blocks SX1, SX2 and SX4 tested in C(T)-0.5T at 163 K [2]. It is of interest to note that if our data points are included together with those of Ref. [21], they do not appear to lie outside the overall distributions (see white circles in Fig. 24). In Section 10.1.4 of the ASTM E 1921-05 [25], specimen size effects on a single KJc value of the dataset, KJc(med) or K0 from test specimens SE(B)-0.4T to C(T)-1T are calculated using Eq. (6), which applies in the transition region (between lower shelf and upper shelf fracture toughness):   BSEðBÞ-0:4T 1=4 K JcðCðTÞ-1TÞ ¼ K min þ ½K JcðSEðBÞ-0:4TÞ  K min   ð6Þ BCðTÞ-1T where B is the thickness of the test specimen. Eq. (6) has been derived to take into account the weakest link size effect only, without constraint effects [26]. Eq. (6) is the simple weakest link scaling model, Eq. (7), modified for the threshold toughness [27]. Eqs. (6) and (7) are only applicable under plane-strain, SSY conditions:   BSEðBÞ-0:4T 1=4 K JcðCðTÞ-1TÞ ¼ K JcðSEðBÞ-0:4TÞ  ð7Þ BCðTÞ-1T On the lower shelf, size effects due to ‘weakest link’ sampling are not considered to apply, but, for small testpieces, constraint effects are likely to be exhibited. Although the size effect is due to loss of constraint, the

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‘weakest link’ (B0.4T/B1T)1/4 scaling factor to scale fracture toughness values from SE(B)-0.4T to C(T)-1T predicted the experimentally determined plane-strain fracture toughness KIc results reasonably well for a C–Mn steel, in two conditions, non-prestrained and 9% prestrained [28]. We have found it of interest to explore the efficacy of the factor implied in Eq. (7) for the DIN 22NiMoCr37 steel (noting that (0.4)1/4  0.8). The efficacy of the factor 0.8 of Eq. (7) for the DIN 22NiMoCr37 steel is shown in Fig. 24. The predicted values of fracture toughness for C(T)-1T, scaled from SE(B)-0.4T using the simple scaling model (Eq. 7) for the data given in Ref. [21] (black circles in Fig. 24) are compared with the experimental C(T)-1T values for data given in Ref. [2]. A good agreement can be seen for the predicted and experimental fracture toughness KJc results for C(T)-1T values using Eq. (7). From Table 3, it can be seen that the lowest value of M is 90, which is well above the ASTM E1921-05 minimum value of 30. Other work, however both computational [27] and experimental [29] suggest that in small specimens constraint is lost for M 6 170. If this more strict limit were to apply, doubt would be cast, in general, on the results for the 1373 condition. The M value for the AR condition range between 176 and 446. In Eq. (2), the minimum values of b0 correspond to maximum a/W values (0.55) giving a minimum b0 = 4.5 mm for SE(B)-0.4T and a minimum b0 = 22.5 mm for C(T)-1T (a maximum b0 = 5.5 mm for SE(B)-0.4T and a maximum b0 = 27.5 mm for C(T)-1T). Therefore, values of M calculated from the experimentally determined KJc at 119 K for C(T)-1T-AR are all above 1000. For the experimentally determined KJc at 119 K for SE(B)-0.4T-AR, when KJc 6 61.9 MPa m1/2 (corresponding to the majority of values of the dataset) then M(min.) P 190, and loss of constraint is not expected according to the above more strict limit criteria. Whereas, when KJc 6 72.5 MPa m1/2 (the maximum value of the dataset) then M(max.) 6 169 and lost of constraint is expected according the above more strict limit criteria. References [1] [2] [3] [4] [5] [6] [7] [8] [9] [10] [11] [12] [13] [14] [15] [16] [17] [18] [19] [20]

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