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Dynamic fracture toughness properties of a 9Cr1Mo weld from instrumented impact and drop-weight tests
Inr. J. Prrs. Ves. & Piping 69 (1996) 149%159 Copyright IQ 1996 Elsevier Science LimIted Prmted in Northern Ireland. All rights reserved 030X-0161/96/$15.00
0308-0161(95)00126-3
ELSEVIER
Dynamic fracture toughness properties of a 9Cr-1Mo weld from instrumented impact and drop-weight tests P. R. Sreenivasan, Materials
Development
Division,
Indira
A. Moitra,
S. K. Ray, S. L. Mannan
(Gandhi Centre for Atomic
Research, Kalpakkam-603102,
India
& R. Chandramohan Quality
Assurance
Division,
Indira
Gandhi
Centre for Atomic
Research, Kalpakkam-603102,
India
(Received 1.5August 1995; accepted 16 November 1995) This paper reports the RTNDT and K,, results obtained from instrumented impact and drop-weight tests of 9Cr-1Mo welds. RTNDT results for welds prepared usingelectrodesof diameter 2.5, 3.15 and 4 mm agreewithin 5K, the respective values being 264, 269 and 268K. For all the three welds,269K can be taken as the conservative RT,,,. The procedurespresentedin this paper enable estimationof reasonablyconservative values of K,, from instrumented impact test of unprecracked Charpy V-notch specimens.For the range of temperatures 193-303K, the lower bound K,, estimates obtained for the 9Cr-1Mo welds are higher than the ASME K,, curve; hence, the ASME K,, curve will give conservative K,, values for the welds of this 9Cr-1Mo steel in the above temperature range; at higher temperatures,the applicability of the ASME K,, curve to the welds of the present 9Cr-IMo steel needs verification/validation. Microcleavage fracture stress or cleavage fracture strength, (TV,estimatedfor the 9Cr-1Mo welds is about 2160MPa and is much lower than the 2550-285OMPa reported for a normalised and tempered 9Cr-1Mo plate material. Copyright 0 1996Elsevier ScienceLtd.
deciding the hydro-test temperature. RT,,, is the higher of (T,, - 33K) and TNDT, where TNDT is the drop-weight (DW) nil-ductility temperature (NDT) and T,, is the temperature at which a minimum Charpy V-notch (CVN) specimen energy (Cd (average of three tests) of 68 J and a lateral expansion (LE) (average of three tests) of at least 0.89 mm are obtained. For application of linear elastic fracture mechanics (LEFM), Appendix G, Section III of the ASME Code” provides an empirical relation describing the variation of K,, (the lower bound of toughness for static, dynamic and crack arrest tests) with (T - RT,,,), where T is the test temperature. The K,, curve is specifically applicable to steels with a minimum specified yield strength of 34.5MPa or less at room temperature: for steels with minimum specified yield strength of
1 INTRODUCTION
Superior creep-rupture properties and improved weldability including resistance to hot cracking and better resistance to hydrogen and temper embrittlement have made 9Cr-1Mo steels (and their modified versions containing V and Nb) major candidates for superheater and reheater tubings and thick section tube-sheet of fast breeder reactor steam generators.lm3 These steels have been widely used in conventional power plants, and it has been found that their properties, in contrast to those of other low-alloy ferritic steels, are remarkably tolerant to variations in heat treatment.‘,3 As welds are known to be regions of weakness, reference nil-ductility transition temperature (Z?TNDT) results for welds are required for 149
P. R. Sreenivasan et al.
150
345-621 MPa at room temperature, the applicability of the ASME KIR curve requires verification by K,, measurements on a minimum of three heats of the material covering the temperature range of interest; testing must include weld metal and the heat affected zone (HAZ).’ This paper reports the results from DW and CVN impact tests carried out using an instrumented test system on three different welds of a 9Cr-1Mo steel, each of which was prepared using electrodes of a different diameter. The load-time (P-t) data obtained were analysed for determining conservative estimates of dynamic fracture toughness values (K,, or Jr,). The results have been compared with the ASME K,, curve. The procedures presented here enable conservative K,, or Jld estimates to be made from the analysis of P-t traces obtained from instrumented impact tests of unprecracked CVN specimens.
the backing strip and base plate with the weld angle being 45”. Three different weld pads, designated as welds I, II and III, were used in the present investigation; each of the three different pads was prepared using an electrode of a different diameter, namely, 2.5, 3.15 and 4 mm, respectively. All electrodes were predried for 1 h at 673K, and during welding the interpass temperature was maintained below 533K. The average welding conditions employed for each diameter electrode are reported in Table 1. The welds were given a post-weld heat treatment (PWHT) at 1008 f 5K for 1 h with a maximum heating/cooling rate of 150K/h. The chemical compositions and the ambient temperature tensile and impact properties of the weld (as given in the test certificate)6 are reported in Table 2; the three different weld pads show similar properties, but slightly higher tensile ductility and impact values are shown by weld I.
2 MATERIAL DETAILS
2.2 Drop-weight preparation
2.1 Preparation, weld
AND EXPERIMENTAL properties
and composition
of
The weld details are given in Fig. 1. IS 2062 steel (C: O-13; S: O-023; P: 0.016; Si: 0.09) was used as
and impact
specimen
The layout of the DW and CVN impact specimens is shown in Fig. 1. The crack plane orientations for the DW and CVN specimens were T-S and T-L, respectively7 (T here stands
r WELD
L
METAL
WELD METAL DROP WEIGHT
z
WELD
METAL
CHARPYl
NOTE :
WELD
DETAILS
AND
SPECIMEN
ALL DIMENSIONS
ARE IN mm.
LAYOUT
Fig. 1. Weld detailsand specimenlayout for drop-weight and Charpy specimensof 9Cr-kMo welds.
of a 9Cr-1Mo
Properties
overtempered zone, if any. All electrodes were predried at 423-453K for 1 h and a current of 190-200 A was employed during weld-bead deposition.
Table 1. Average welding conditions employed during welding using different diameter electrodes of 9Cr-1Mo steel*
Electrode Voltage diameter (mm) (V) 2.5 3.15 4.0
20-27 20-26 19-28
Current (A)
Welding speed (mm/min)
go-105 135-155 135-160-180
50-100-130 60-100-150 122-125
1.51
weld
2.3 Drop-weight
and impact
Drop-weight and impact tests were carried out using a Dynatup Model 8000A drop-tower and a Tinius Olsen Model 74 (358 J capacity) impact machine, respectively; both machines were provided with a Dynatup Model 500 instrumentation system. The P-t traces captured by an analogue storage oscilloscope were photographed using a Polaroid camera for subsequent analysis. Drop-weight and impact tests were performed following the ASME Section III, Div. I guidelines4 and ASEM E 208’ and E 239 procedures and other practices relevant to instrumented impact and drop-weight testing.‘3,‘4 For above-ambient temperature tests, a heated quenching oil bath was used, and for below ambient temperature tests a bath consisting of a methanol-liquid nitrogen mixture was used following the ASTM E 208’ or E 239 requirements as applicable.
* Interpass temperature was maintained below 533K.
for transverse to the weld bead, L for along the weld bead and S in the thickness direction to the weld bead) as required by ASME Section III, DIV. I, NG-2322*2.4 P-3 DW and standard CVN specimens were prepared following the relevant ASTM standards.8.9 Following ASTM E 208 specifications,* to reduce the extent of th.e HAZ, short-weld beads (approximately 25 mm long and 13 mm wide) were deposited on the P-3 specimens using 5 mm diameter BOR-C (a Cr-MO-V air-hardening steel giving approximately a Rockwell C hardness HRc of 52) hard-facing electrode supplied by M/s D&H Secheron Electrodes Pvt. Ltd, Indore. Suitability of these electrodes for preparing drop-weight specimens had been verified earlier.” Also to reduce the HAZ effect further, following some of the latest research reports,““* the specimens were placed on a ‘chiller block’ (approximately 150 X 150 X 150 mm steel block) when depositing the hard-facing weld-bead; use of this chiller block is supposed to reduce the extent of
2.4 RTNDT determination Following the ASME Section III, Div. I guidelines, for each weld-pad (i.e. corresponding
Table 2. Chemical Chemical
compositions and mechanical properties of 9Cr-1Mo prepared using different diameter electrodes composition (wt%)*
Electrode diameter (mm) 2.5 3.15 4.0 Mechanical
C
Mn
Si
testing
Cr
Ni
MO
v
0.10 0.93 0.39 9.11 0.25 0.87 Nil 0.09 0.90 0.43 9.21 0.24 0.87 Nil 0.087 095 0.46 9.08 0.25 0.92 Nil
cu Nil Nil Nil
s
welds
P
0.004 0.004 0.004 0.004 0.003 0.004
properties?
Electrode diameter (mm)
O-2% proof stress(MPa)
LJTs (MPa)
2.5 3.15 4.0
551.7 579.9 548.2
695.6 699.1 712.7
% Elongation Charpy V-notch impact (5.65VA) energiesat 293K (km) 20.3 18.9 19.0
13, 10.4, 11.2 13.9, 9.0, 8.4 10.0, 8.6, 8.2
* Balance-Fe. t All welds were stress relieved at 1003f 5K for 1 h with rates of heating and cooling 150K/h max. Also, prior to welding, all electrodeswere baked for 1 h at 673K giving moisture contents of 0.205, 0.310 and 0.17% in the welds using the 2.5, 3.15 and 4mm diameter electrodes, respectively. Tensile properties are for tests at room temperature.
P. R. Sreeniuusanet al.
152
to each diameter electrode), TNDT was determined. Then, CVN specimens were tested in triplicate at successively higher temperatures, at 5K intervals, starting the first batch of tests at TNDT+ 33K. According to the ASME Section III, Div. I, NG 2330 and NG 2350 requirements on Charpy test values,4 for acceptance, the average value from the triplicate tests should meet the minimum requirements; more than one specimen should not show values below the minimum requirements and, for the specimen not meeting the minimum requirements, the C, and LE values should not be below the minimum specified requirements by more than 10 ft-lb (13.4 J) and 5 mils (O-127 mm), respectively: in cases not meeting these criteria, two additional tests performed should both meet the minimum requirements.
3 INSTRUMENTED ANALYSIS
L
P t
I
t
TYPE-I
P max = PF
TEST DATA
General procedures for analysing instrumented test data are given in Server and co-workers’3”4 and Varga et al.” 3.1 Drop-weight
P max = P,
TYPE
-II
pGY
test data analysis
By assuming that the elliptical thumbnail formed by the HAZ of the weld-bead to be the starter crack, and using the stress-intensity factor analysis of Scott and Thorpe,16 Sreenivasan et al. ‘O and Moitra et aLI7 estimated Kid from P-t traces of ASTM E 208 drop-weight tests at and below TNDT. The same procedure has been adopted here. The results on crack-profile and Kid are given in Section 4. 3.2 Impact test data analysis 3.2.1 General The P-t traces from instrumented tests of CVN specimens can be categorised into types I, II and III as shown in Fig. 2; the general yield, maximum and fast/brittle fracture loads are represented by Pcy, Pm,, and PF, respectively. In type I, brittle or fast running fracture is characterised by a sudden drop in load occurring before or at general yield, indicating linear-elastic material behaviour; this is the situation in the
TYPE-III Fig. 2. Three types of load-time (P-t) traces from instrumented impact tests of Charpy specimens.
lower shelf of the transition region of ferritic steels. In type II, fast fracture occurs after general yielding at (PF = Pm,,) or after the maximum load (PF < P,,,,,); at higher temperatures in the transition region, fast fracture arrests after some propagation and this is reflected in the P-t traces by a tail after some sudden drop. Type III refers to complete ductile fracture with no sudden load drop, signifying fast fracture, and occurs in the upper shelf region. For reliable interpretation of P-t traces from instrumented impact tests, the time to any load measurement point, t, should be greater than 2.32 (or more conservatively 32), where r is the apparent period of oscillation of the three-pointbend (3PB) specimen.‘3,‘4 Moreover, to avoid
Properties of a 9Cr-1Mo weld excessive signal attenuation, f should be greater than or equal to l.lT,, where T, is the response time of the instrumentation system. The value of TR in the present tests was mostly 90 ps (in some tests it was 10 ps), yielding a value of 100 (or 11) ps for l.lT,. The 32 period calculated for CVN steel specimens was about 100 ps: this corresponds to the first three oscillations on the P-t traces. The dynamic yield stress, (TYd, was computed from PGy using the following relation for a standard CVN specimen:13 cr,,(MPa) = 46*7P,,(kN)
(1) The cleavage fracture strength or micro-cleavage fracture stress, nf, was obtained from: o,(MPa) = C,cr,,(MPa)
(2) where (Tyd is the dynamic yield stress at the brittleness transition temperature (T,) at which PF= Pcy (see load-temperature (P-T) diagram results and discussion given later) and the stress intensification factor C, = 2.57 for the standard CVN specimen at T,.” 3.2.2 Kid estimation procedures Krd was estimated using the ASTM E 399’ formula for precracked 3PB specimens taking a/W = O-2 and the final expression for standard CVN specimen is: K,, = 4.67P,
(3) where PF is in kilonewtons. An error analysis for K,, reported elsewhere’0,‘7 shows that ,the error band can be taken as &20%. For type I cases, the application of a reduction factor of 0.8 (the lower bound of the 20% error band) to the results from CVN specimens using eqn (3) can be expected to give results in the error band of precracked Charpy K1d tests.” Wherever the P-t traces showed time-to-failure, tf, between z and 22, PF should be replaced by PC, a critical fracture load obtained by Varga’s procedure.‘5.‘0 Turner,*’ based on beam-on-foundation analysis of instrumented CVN tests, has given graphical correction factors to be applied to PF derived from P-t traces showing only one or two oscillations. In the present case, these inertia correction factors were found to be smaller than 20% of those obtained by Varga’s procedure. For type II and III test traces (Fig. 2), an elastic-plastic procedure has to be applied. Usually such traces have been analysed using a
153
J-integral or crack tip opening displacement (CTOD) procedure assuming that crack initiation occurs at the maximum load.*‘,“* For impact tests of CVN specimens of ferritic steels, Ghoneim and Hammadz3 proposed that fracture initiation occurs at a load equal to (Pm,, + P,,)/2, while Norris,24 based on a comparison of finite element results with experimental data, reported that the time to crack initiation, ti, equals 40% of the time to reach the maximum load, t,,,. For CVN specimens of ASTM A 533 Class 1 steel, Kobayashi” has reported a value of 0.8 for the ratio of initiation to maximum load energy. In the present case, taking tiltmax = O-4 gave slightly smaller values for the initiation energy compared with Ghoneim and Hammad’s procedure and much smaller values compared wtih Kobayashi’s procedure. Hence, J or CTOD estimates were computed assuming initiation to occur at a tiltmax value of O-4. The average velocity, V, and the hammer travel (i.e. load point displacement, LPD), d, at any time, t, during impact were obtained using the usual relations:21~22~26 d = tV = tVo(l - E,/4E,)
(4) where V, is the velocity of the pendulum at the point of impact, E, is the initial pendulum energy and E, = V,J P dt, J P dt being the total area under the P-t trace. From this the plastic part of the LPD, d,, is obtained by subtracting d,,, the total system elastic deflection given by:13 d,, = C,P
(54
where P is the load at t (corresponding to d) and C, is the total system compliance given by:13 CT = (VohxIPGY> - W2,,l@Eo)
w
where the subscript GY signifies values at general yield. For a standard CVN specimen (the same holds for a half-thick CVN specimen, for the relation is independent of thickness), following the derivation in Mutoh et al.,*’ and assuming that the 3PB specimen during loading undergoes rotation by a plastic hinge mechanism with a rotational factor of 0.4, the relation between the plastic part of CTOD, apl, and d,, had been given as:*’ S,, = 0.32d,,
(64 However, according to more recent results’8Jy for shallow cracks (a/W = 0.1%0.3) r values are in
P. R. Sreenivasan et al.
154
the range 0.2-O-3, while for deeply cracked specimens Y = 0.4-0.45. Hence, in the present paper taking Y = O-2 modifies eqn (Sa) to: 6,, = O.l6d,’
w
Then total CTOD at initiation, 6i, is given as the sum of the elastic and plastic parts:29
Table 3. TNDT, RTNDT and TD transition-temperatures and cleavage fracture strength (~3 for 9Cr-1Mo welds
Electrode diameter (mm) 2.5 3.15 4.0
TNDT Gv G-33) WI (J-9 WI 263 297 268 302 268 301
R&m W
264 269 268
264 269 268
TD W ($4 264 2112 255 2166 -
6i = 6, + Spl = (K2(1 - u*)/(2a,,E))
+ 6,,
(7)
where K is calculated using eqn (3) wtih P = Pi at ti, oys is replaced by cyd and Young’s modulus E = 190GPa for the range of temperatures involved. Then Kid(S) is given by Kld(S)
=
d(EaYd6i)
(8)
The value of Ji at initiation is determined from the following formula due to Rice et aZ.:30 Jj = n Ei/(Sb)
Pa>
where n = 2 for deep-cracked 3PB specimens, B is the thickness, b is the remaining ligament depth (= W - a) and Ei is the total energy (excluding the machine contribution) up to initiation. Based on a critical consideration of the expressions given in Sumpter,28 for a specimen with a/W = 0.2, n can be taken as 1.45. Then J, = 1’45Ei/(Bb)
cw
In eqns (9a) and (9b), Ei is obtained from the total absorbed energy E as follows:‘3 Ei = E - (P2/2) * [CT - CND/(EB)]
(11)
In the literature Kid has been estimated by the Charpy energy (C,) correlation approach.3’ One of the most successful correlations applicable over a wide range of C, (3-95 J) and room temperature static yield strength (gys = 270815 MPa) values is the following: K’,(C”) = 15*5(cv)“~37s
AND DISCUSSION
4.1 RT,,, Table 3 reports the RTNDr results; clearly, for all the welds RT,,, agrees with the respective TNDT within lK, which is within the accuracy of temperature measurement. These TNDT values are worth comparing with those reported for a modified 9Cr-1Mo plate steel in the normalized and tempered condition: 268.6-279*7K.32 In the plots of Kid versus RTNDT reported later, RT,,, of welds I, II and III have been taken as 263,268 and 268K, respectively (i.e. equal to the respective TNDT values). The lower RTNDT (i.e. tougher behaviour) of weld I is in line with the tensile and impact results shown in Table 2. This may be due to the greater refinement in structure occurring in the weld prepared using a smaller diameter electrode. Since the resolution of TNDT is only 5K, referring to Table 3, the RTNDT of all the three welds can be conservatively taken as 269K.
(10)
where CND is the non-dimensional specimen compliance given in ServerI (equal to 24.39 for a standard CVN specimen) and P is the load at t (equal to ti in present case). Then Kid(J) = q(EJi)
4 RESULTS
(12)
The above was used in the present paper to estimate Kld(Cv) from C, values.
4.2 Load-temperature
diagram
Load-temperature (P-T) diagrams for welds are shown in Fig. 3. Paucity of specimens did not permit testing of CVN specimens of weld III (i.e. prepared using 4 mm diameter electrodes) at low temperatures. The results show considerable scatter, hence there is likely to be more uncertainty in the TD values determined. However, the loads corresponding to PF = PGy are not likely to be in error by more than lo%, which is only within the uncertainty of instrumented impact test load values. T, and gf values determined are reported in Table 4. The values of 2112-2160 MPa for the (T’ are reasonable, but are much lower than the reported values of 2550-2850MPa for a normalised and tempered plate metal. 33 This indicates that the fracture
Properties of a 9Cr-1Mo 9Cr-1Mo 20
2.5mm
WELD
Q ELECTRODE
/
/
A-
10
1
173
B 1
1
1
1
1
1
223
1
1
1
273
1
PGY
’
1
1
1
323
T/K Fig. 3. Load-temperature diagrams from instrumented impact tests of Charpy V-notch specimens of 9Cr-1Mo welds I and II (2.5 and 3.15 mm diameter electrodes, respectively).
resistance of the welds is poorer than that of the base metal.
Kid values estimated by various procedures and the computed gyd values are reported in Table 4. For drop-weight tests, crack profile measurements on fracture surfaces of broken specimens yielded average values of 0~241(*0~037) and O-296 (*O-056) for (a/2c) and (a/t), respectively (values in parentheses are two standard deviation values; n and c are the semi-minor and semi-major axes of the elliptical thumbnail crack bounded by the HAZ of the crack-starter weld-bead on the broken specimen surface and t is the specimen thickness, respectively). The above ratios compare very well with the values of (a/2c) = 0.25 f 0.02 and (a/t) = 0.27 f 0.04 reported for an AISI 403 stainless steel.” Final fOrIIIUkE for Kid from drop-weight tests Of P-3 specimens obtained in the present test campaign were as given below: K,,(MPadm) = 0.41Pcdu at TNDT! = 0-38P,~u at T < ThIDT! (13)
155
weld
where PC (kN) is the critical fracture load derived using Varga’s procedure” and a is in millimetres; eqn (11) shows excellent agreement with the expressions reported in Sreenivasan et al.;” these Kid estimates are accurate to &20% and have been reported as Kld(DW) in Table 4. The K&%d ratio at TNDT evaluated from crack profile measurements on broken specimens is O-074 f 0,016 drn and is in good agreement with that reported in the literature.” The various Kid values reported in Table 4 have been plotted against (T - RTNDT) in Figs 4-6. The ASME KIR curve4 has also been plotted for comparison. An error band of *20% was applied to K,,(DW), Kid (PJP,) and Kld(CV) for reasons mentioned earlier. The error bands for Kid(8) and Kid(J) were taken as *25% and &35%, respectively (see the Appendix). From Figs 4-6, it is evident that Charpy correlation K,, values (Kld(CV)) are not conservative with respect to the ASME K,, curve. It must be borne in mind that most of the Charpy correlations reported have been based on results for carbon and low-alloy steels. As such, the applicability of these correlations to high alloyed steels like 9Cr-lMo, 12Cr-steels, etc. requires verification. James and Carlson3’ had applied C,-K,, correlation and a temperature shift procedure to obtain the K,, versus temperature curve of a modified 9Cr-1Mo steel; a single KIc determined at the upper shelf was much higher than that predicted by the above method. In the present instance also it is likely that the KId(CV) values are unduly pessimistic. The lower bound Vahes from all the other Kid estimates plotted in Figs 4-6 are conservative (i.e. give higher values) with respect to the ASME K,, curve over the range of temperatures covered in this investigation: 193-303K. It must be noted that taking the R Lx- as 269K for all the welds (as suggested earlier under the discussion on RTNDT) would increase the conservatism of the ASME KIR curve with respect to the K,, estimates. 4.4 Reliability
of Kid values
Since the available Charpy correlations seem to give unduly pessimistic values, it is desirable to develop more reliable correlations between C, and K,c/K,, results based on proper tests (proper according to established and accepted standards and procedures for Ki,-/Kid determination).‘3-‘5 The alternative procedures for estimating K,,
P. R. Sreenivasan et al.
156 Table
Electrode diameter (mm) 4
19.5 223 268 293 297 297 302
3.15
191 226 253 263 268 273 300 302
2.5
191 227 253 263 273 293 297
4. Dynamic
yield strength
(uYd) and various K,, estimates for 9th1Mo
welds
UYd
K,,(J)
K,.,(S)
KdP~lf’c)*
KM(G)
f MW
(MPadm)
(MPadm)
(MPagm)
(MPaqm)
(MPat/m)
-
67.6 65.8 80.2 80.6
56.4 54.6 744 -
805.1 797.6 786.3 783.4 786.9 792 748.2 780.6 779.4
KdDV
157 169.6 181 180.7
93.4 94.6 99.2 96.8
142.5 195-5 184
88.2 117.4 98.1
52.21 68.2$ 81.2 -
24.11 28.33 36.7 59.3 75 80.4
61.4 56.3 61 76 70.8 -
195.6 195.1
109 105
75.21 86.48 77.5 61.5 -
28.3t 33%$ 36 47.7 77.1 81.2
45.7 56 67 -
*K,, from PF/Pc obtained from linear-elastic load-time trace from instrumented test of unprecracked Charpy specimen usingASTM E 399 formula (seetext for further explanation). ‘r Resultsat 193K. $ Results at 223K.
from instrumented test P-t traces of CVN specimens discussed above are promising when used with sufficient engineering judgement. The lower bound values of K,,(DW) give reliable estimates of Kid. In most cases, K,, from PF/Pc of CVN specimens giving type I P-t traces give values in agreement with K,,(DW). A minimum of three or four CVN specimens should be tested at each temperature and over a range of temperatures and a lower bound trend line taken. Appropriate statistical procedures, though not applied in the present case, are available to account for scatter.3b36 The test temperatures should not be too low to produce type I P-t traces with only a single inertial load peak that violates the 22-32 requirements. In the case of type II and type III P-t traces, it is seen that Kid(S) is less than K&). However, Duffy et d3’ state that crack opening displacements (CODS) are not meaningful in an absolute sense when the material is very tough. For type II traces, up to temperatures where PF = Pm,,, K&6) can be taken as a conservative estimate of K,,. In making use of these estimates, a smooth lower bound curve may be drawn. ‘The error bands
used in this paper and discussed in the Appendix are based on three or more weld CVN specimens tested at each temperature. Application of these error bands to homogeneous specimen results can be expected to give conservative values. Modification of these error bands may be done based on further CVN tests and comparison with precracked Charpy tests.
5 CONCLUSIONS 1. The RT,,* of 9Cr-1Mo welds prepared using electrodes of diameter 2.5, 3.15 and 4 mm agree within 5K; the respective values being 264, 269 and 268K; these are well within the 5K resolution expected for TNDT. Hence, with utmost conservatism, the RTNDT of all the three welds can be taken as 269K. 2. The procedures presented in this paper enable estimation of reasonably conservative values of K,, from instrumented impact tests of unprecracked Charpy V-notch specimens. 3. For the range of temperatures over which
of a 9Cr-1Mo
Properties
9Cr-1Mo
200 9Cr-1Mo
200
150
WELD:
2Smm
0 ELECTRODE
v
K,.,
FROM
DROP
WEIGH1
0 A
Ytd
FRQM
PF
K,,,
FROM
6
El m
Kid FROM J LOWEST VALUES IN THE ERROR BAND OF KldtJl
ta
K,,, FROM
1.57
weld
130
TEST
WELD:
0 ELECTRODE
v A
Kid FROM 6 K,* FROM J LOWEST VALUES IN THE ERROR BAND OF Kld(Jl K,, FROM Cy CORRELATION
q q 0
FROM
&mm
Kid
DROP WEIGHT
TEST 88
El
150
CV CORRELATION
i
ki 0 &
100
\ 0 Y
r 50
LASME
L ASME
K,,
0 0
I
I
I
I
I
100
I
I
I
223
I
I
I
I1
1
I
RT,,,VK
Fig. 4. Variation of K,, with (T-RTNDT) for 9Cr-IMo weld I (2.5 mm diameter electrode).
tests were carried out (193-303K), the lower bound K,, estimates obtained for thLe 9Cr1Mo welds were higher than the ASME K,, 200
-
9Cr-1Mo
-
v
-
-
150
Kid
WELD: FROM
3.15mm
0
Kid
FROM
PF
A
Kid
FROM
6
FROM
J
0
Kid
q
LOWEST VALUES IN THE ERROR BAND OF K,,,(J)
I3
K,.,
-
FROM
d ELECTRODE
DROP WEIGHT
El
TEST
q
Cv CORRELATION
CURVE
El
“‘([“I”““’ 173
323
273
(T-
K,,
CURVE
T
223
273 (T-
RT,,
323
11 K
Fig. 6. Variation of Kid with (T-RTNDT) for 9Cr-1Mo weld III (4 mm diameter electrode).
curve. Hence the ASME KIR curve provides a conservative lower bound over these temperatures. For this material, applicability of the ASME K,, curve at higher temperatures needs validation. 4. Microcleavage fracture stress, gf, estimated for the 9Cr-1Mo welds is about 2160MPa; this is much lower than the 2550-2850 MPa reported for a normalised and tempered 9Cr-1Mo plate material, and indicates poorer fracture resistance for the welds.
REFERENCES 100
1. Hudson, J. A., Druce, S. G., Gage, G. & Wall, M., Thermal aging effects in structural steels. Theor. Appl. Fracf. Mech, 10 (1988) 123-33. 2. Fidler, R. S., The creep of normalised and tempered 9Cr-1Mo: effects of stress,temperature and grain size on creep and rupture properties. CERL Report No. RD/L/R 1949,1976. 3. Fidler, R. S. & Middleton, C. J., The impact and hot tensile properties of 9Cr-1Mo steel in various heat treatment conditions. In Proc. IAEA Specialists
50
173
223
273
323
(T-RT,DT)/ K Fig. 5. Variation of K,, with (T-RTNDT) for 9Cr-1Mo weld
II (3.15 mm diameter electrode).
Meeting-Mechanical Properties of Structural Materials Including Environmental Effects, Chester,UK, 1983,pp.
519-44. 4. ASME Boiler and PressureVessel Code, Section III Nuclear Power Plant Components,Division 1, Appendix G. ASME. New York, 1989.
158
P. R. Sreenivasan
5. Lidbury, D. P. G. & Morland, G., Review of fracture toughness requirements and data relevant to LWR reactor pressurevessels,Znt. J. Pres. Ves. & Piping, 29, (1987) 343-428. 6. Test Certificate SerialsNos 18474,19228and 18471,dtd. 05-01-1993 from M/s. Advani Oerlikon Limited, Bombay on CITOCHROM-9R SFA-5.4, E 505-15. 7. ASTM E 399-83,Standard test method for plane-strain fracture toughnessof metallic materials. 1990 Annual Book of ASTM Standards, Vol 03.01, 1990, pp. 488-512. 8. ASTM E 208-87a,Standard test method for conducting drop-weight test to determine nil-ductility transition temperature of ferritic steels. 1990 Annual Book of ASTM Standards, Vol. 03.01, 1990,pp. 360-71. 9. ASTM Standard E 23-88,Standardmethodsfor notched bar impact testing of metallic materials. 1990 Annual Book of ASTM Standards, Vol. 03.01, 1990, pp. 197-212. 10. Sreenivasan, P. R., Ray, S. K., Mannan, S. L. & Rodriguez, P., Determination of K,, at or below NDTT using instrumenteddrop-weight testing, Znt. J. Fract. 55 (1992) 273-83.
11. Holt, J. M. & Puzak, P. P. (eds), Drop-weight test for determination of nil-ductility transition temperature of ferritic steels:user’sexperience with ASTM method E 208. ASTM STP 919, American Society for Testing and Materials, 1986. 12. Satoh, M., Funada, T. & Tomimatsu, M., Evaluation of valid nil-ductility transition temperatures for nuclear vesselsteels.ASTM STP 919, pp. 16-33. 13. Server, W. L., Impact three-point bend testing for notched and precracked specimens,J. Test. Eval., 6 (1978) 29-34. 14. Ireland, D. R., Server, W. L. & Wullaert, R. A., Proceduresfor testing and data analysis,effects. Tech. Inc., Technical Report, TR 75-43, 1975. 15. Varga, T., Junker, M., Njo, D. H. & Prandtl, G., Proposedmethod for instrumented precracked Charpy type tests: ASK-AN-425, Rev 1, December 1978. In C.S.N.I. SpecialistMeeting on Instrumented Precracked Charpy Testing. EPRI NP-2102-LD, November 1981, pp. 1-85-l-115. 16. Scott, P. M. & Thorpe, P. W., A critical review of crack tip stress intensity factors for semi-elliptic cracks, Fatigue Eng. Mater. Struct., 4 (1981) 291-309. 17. Moitra, A., Sreenivasan,P. R., Ray, S. K. & Mannan, S.
L., Instrumented drop-weight test resultsfor normalised and tempered 9Cr-1Mo steel, communicatedto Znt. J. Fracture (1995). 18. Ray, S. K., Sreenivasan, P. R., Samuel, K. G. &
Rodriguez, P., Characterisation of irradiation-induced embrittlement in an ASTM A 203D steel using instrumentedimpact testing, Znt. J. Pres. Ves. h Piping, 54 (1993)481-96. 19. Krompholz, K., Tipping, P. & Ullrich, G., ‘Resultsfrom investigationswith an instrumentedimpact machine on a molybdenum base alloy, nickel base alloys, and Incoloy 800’, Z. Werkstoftech., 15 (1984) 117-23. 20. Turner, C. E., Measurement of fracture toughnessby instrumented impact testing, Impact Testing of Metals, ASTM STP 466 (1970) 93-114. 21. Samuel, K. G., Sreenivasan, P. R., Ray, S. K. &
Rodriguez, P., Evaluation of ageing-inducedembrittlement in an austenitic stainlesssteel by instrumented impact testing, J. Nucl. Mater., 150 (1987) 78-84. 22. Iyer, K. R. & Miclot, R. B., Instrumented Charpy
et al. testing for determination of the J-integral, Instrumented Zmpact Testing, ASTM
STP 563 (1974) 146-65.
23. Ghoneim, M. M. & Hammad, F. H., Instrumented impact testing of an irradiated 20MnMoNi55 PVS weld material, J. Nucl. Mater., 186 (1992) 196-202. 24. Norris, D. M., Jr, Computer simulation of the Charpy V-notch test, Eng. Fract. Mech, 11 (1979) 261-74. 25. Kobayashi, T., Analysis of impact properties of A533 steel for nuclear reactor pressurevesselby instrumented impact test. Eng. Fract. Mech., 19 (1984) 49-65. 26. Ireland, D. R., Proceduresand problemsassociatedwith reliable control of instrumented impact testing, Instrumented Impact Testing, ASTM STP 563 (1974) 3-29.
27. Mutoh, Y., Toyoda, M. & Satoh, K., On the relation between load-point displacement and stretch zone width, Znt. J. Fract., 16 (1980) R171-4. 28. Sumpter, J. D. G., Jc determination of shallow notch welded bend specimen, Fatigue Fract. Eng. Mater. Struct., 10 (1987) 479-93.
29. Sorem, W. A., Dodds, R. H., Jr & Rolfe, S. T., Effects of crack depth on elastic-plastic fracture toughness,Znt. J. Fract., 47 (1991) 105-26. 30. Rice, J. R., Paris, P. C. & Merkle, J. G., Some further resultson J-integral analysisand estimates,Progress in Flaw Growth and Fracture STP 536 (1973) 231-63.
Toughness
Testing,
ASTM
31. Roberts, R. & Newton, C., Interpretive report on small-scaletest correlation with KIC data, Welding Research Council Bull. 265, 3 February 1981. 32. James, L. A. & Carlson, K. W., The fatigue-crack growth and ductile fracture toughness behaviour of ASTM A387 Grade 91 steel, J. Pres. Ves. Tech. (Trans ASME),
107 (1985) 271-8.
33. Packiaraj, C. C., Sreenivasan,P. R., Moitra. A., Ray, S. K., Mannan, S. L. & Arivazhagan, B., Effect of simulated post weld heat treatment on toughnessand transition temperature characteristicsof normalisedand tempered 9Cr-1Mo ferritic steels. Presented at the Poster Session of 49th ATM, IIM, Calcutta, 14-17 November 1995. 34. Bishop, T. A., Markworth, A. J. & Rosenfield, A. R., Analysing statistical variability of fracture properties, Metall.
Trans., 14A (1983) 687-93.
35. Oldfield, W., Fitting curves to the toughnessdata, J. Test. Eval.. 7 (1979)326-33. 36. Koons, G. ‘F., ‘Statical analysisof notch toughnessdata, J. Test. Eval., 7 (1979) 96-102. 37. Duffy, A. R., Eiber, R. J. & Maxey, W. A., Recent work on flaw behaviour in pressurevessels.In Practical Fracture Mechanics for Structural Steel, Proc. Symp. on Fracture Toughness Concepts for Weldable Structural Steel, Risley, UK, ed. M. 0. Dobson. UKAEA and
Chapmanand Hall, Risley, 1969pp. Ml-M34. 38. Burdekin, F. M., Crack opening displacement-a review of principlesand methods.ibid. pp. Cl-C12. 39. Heerens,J., Cornet, A. & Schwalbe,K.-H., Resultsof a round robin on stretch zone width determination, Fatigue Fract. Eng. Mater. Struct. 11 (1988) 19-29.
40. Sreenivasan, P. R., Ray, S. K., Vaidyanathan, S. & Rodriguez, P., Measurementof stretch zone height and its relationship to crack tip opening displacementand initiation J-value in an AISI 316 stainless steel, communicated to Fatigue Fract. Eng. Mater. Struct. (1995). 41. Broek, D., The Practical use of Fracture Mechanics. Kluwer, Dordrecht (1989).
Properties oi ~1K’r-1Mo
APPENDIX: ERROR BANDS VARIOUS K,, ESTIMATES
FOR
weld
‘lherefore 2SKIK
The basis for assuming an error band of +20% for Kld(DW), Kld(PF/PC) and Kld(CV) estimates has sufficient literature support.‘0,38 In the following an error analysis has been given for Kid(8) and K&) estimates. An assumption of &40% for the error in 6 is reasonable;3x.3Qresults on 6 for a 316 stainless steel obtained from stretch zone measurements4’ show a two standard deviation varying from 28 to 54%, with most of the values lying in the range 40-45%. Similarly, an error of 50-70% for J is reasonable.4’ Moreover, J estimates in the present tests, based on three or more Charpy tests of welds in the same condition at each temperature, yieldled two standard deviation values in the range 25-70%. So an error value of 70% for J has been taken in the following analysis. K(S) = v(EaydS)
159
(Al)
= 6EIE
+ 6ay,/a,,
+ &S/6
(A2)
Taking error values of 10% for Young’s modulus E and flyd is reasonable. Then the root mean
square error in K from eqn (A2) computes to &21%. Therefore, an error of *25% has been assumed for Kid(a). It must be noted that the error bands taken are assumed to account for notch sharpness effects also. Similarly: K(J) = q(EJ)
WI
Then 26KIK
= 8ElE
+ SJlJ
Thus the root mean square error computes to +35%.
(-w
in Kid(J)
0308-0161(95)00126-3
ELSEVIER
Dynamic fracture toughness properties of a 9Cr-1Mo weld from instrumented impact and drop-weight tests P. R. Sreenivasan, Materials
Development
Division,
Indira
A. Moitra,
S. K. Ray, S. L. Mannan
(Gandhi Centre for Atomic
Research, Kalpakkam-603102,
India
& R. Chandramohan Quality
Assurance
Division,
Indira
Gandhi
Centre for Atomic
Research, Kalpakkam-603102,
India
(Received 1.5August 1995; accepted 16 November 1995) This paper reports the RTNDT and K,, results obtained from instrumented impact and drop-weight tests of 9Cr-1Mo welds. RTNDT results for welds prepared usingelectrodesof diameter 2.5, 3.15 and 4 mm agreewithin 5K, the respective values being 264, 269 and 268K. For all the three welds,269K can be taken as the conservative RT,,,. The procedurespresentedin this paper enable estimationof reasonablyconservative values of K,, from instrumented impact test of unprecracked Charpy V-notch specimens.For the range of temperatures 193-303K, the lower bound K,, estimates obtained for the 9Cr-1Mo welds are higher than the ASME K,, curve; hence, the ASME K,, curve will give conservative K,, values for the welds of this 9Cr-1Mo steel in the above temperature range; at higher temperatures,the applicability of the ASME K,, curve to the welds of the present 9Cr-IMo steel needs verification/validation. Microcleavage fracture stress or cleavage fracture strength, (TV,estimatedfor the 9Cr-1Mo welds is about 2160MPa and is much lower than the 2550-285OMPa reported for a normalised and tempered 9Cr-1Mo plate material. Copyright 0 1996Elsevier ScienceLtd.
deciding the hydro-test temperature. RT,,, is the higher of (T,, - 33K) and TNDT, where TNDT is the drop-weight (DW) nil-ductility temperature (NDT) and T,, is the temperature at which a minimum Charpy V-notch (CVN) specimen energy (Cd (average of three tests) of 68 J and a lateral expansion (LE) (average of three tests) of at least 0.89 mm are obtained. For application of linear elastic fracture mechanics (LEFM), Appendix G, Section III of the ASME Code” provides an empirical relation describing the variation of K,, (the lower bound of toughness for static, dynamic and crack arrest tests) with (T - RT,,,), where T is the test temperature. The K,, curve is specifically applicable to steels with a minimum specified yield strength of 34.5MPa or less at room temperature: for steels with minimum specified yield strength of
1 INTRODUCTION
Superior creep-rupture properties and improved weldability including resistance to hot cracking and better resistance to hydrogen and temper embrittlement have made 9Cr-1Mo steels (and their modified versions containing V and Nb) major candidates for superheater and reheater tubings and thick section tube-sheet of fast breeder reactor steam generators.lm3 These steels have been widely used in conventional power plants, and it has been found that their properties, in contrast to those of other low-alloy ferritic steels, are remarkably tolerant to variations in heat treatment.‘,3 As welds are known to be regions of weakness, reference nil-ductility transition temperature (Z?TNDT) results for welds are required for 149
P. R. Sreenivasan et al.
150
345-621 MPa at room temperature, the applicability of the ASME KIR curve requires verification by K,, measurements on a minimum of three heats of the material covering the temperature range of interest; testing must include weld metal and the heat affected zone (HAZ).’ This paper reports the results from DW and CVN impact tests carried out using an instrumented test system on three different welds of a 9Cr-1Mo steel, each of which was prepared using electrodes of a different diameter. The load-time (P-t) data obtained were analysed for determining conservative estimates of dynamic fracture toughness values (K,, or Jr,). The results have been compared with the ASME K,, curve. The procedures presented here enable conservative K,, or Jld estimates to be made from the analysis of P-t traces obtained from instrumented impact tests of unprecracked CVN specimens.
the backing strip and base plate with the weld angle being 45”. Three different weld pads, designated as welds I, II and III, were used in the present investigation; each of the three different pads was prepared using an electrode of a different diameter, namely, 2.5, 3.15 and 4 mm, respectively. All electrodes were predried for 1 h at 673K, and during welding the interpass temperature was maintained below 533K. The average welding conditions employed for each diameter electrode are reported in Table 1. The welds were given a post-weld heat treatment (PWHT) at 1008 f 5K for 1 h with a maximum heating/cooling rate of 150K/h. The chemical compositions and the ambient temperature tensile and impact properties of the weld (as given in the test certificate)6 are reported in Table 2; the three different weld pads show similar properties, but slightly higher tensile ductility and impact values are shown by weld I.
2 MATERIAL DETAILS
2.2 Drop-weight preparation
2.1 Preparation, weld
AND EXPERIMENTAL properties
and composition
of
The weld details are given in Fig. 1. IS 2062 steel (C: O-13; S: O-023; P: 0.016; Si: 0.09) was used as
and impact
specimen
The layout of the DW and CVN impact specimens is shown in Fig. 1. The crack plane orientations for the DW and CVN specimens were T-S and T-L, respectively7 (T here stands
r WELD
L
METAL
WELD METAL DROP WEIGHT
z
WELD
METAL
CHARPYl
NOTE :
WELD
DETAILS
AND
SPECIMEN
ALL DIMENSIONS
ARE IN mm.
LAYOUT
Fig. 1. Weld detailsand specimenlayout for drop-weight and Charpy specimensof 9Cr-kMo welds.
of a 9Cr-1Mo
Properties
overtempered zone, if any. All electrodes were predried at 423-453K for 1 h and a current of 190-200 A was employed during weld-bead deposition.
Table 1. Average welding conditions employed during welding using different diameter electrodes of 9Cr-1Mo steel*
Electrode Voltage diameter (mm) (V) 2.5 3.15 4.0
20-27 20-26 19-28
Current (A)
Welding speed (mm/min)
go-105 135-155 135-160-180
50-100-130 60-100-150 122-125
1.51
weld
2.3 Drop-weight
and impact
Drop-weight and impact tests were carried out using a Dynatup Model 8000A drop-tower and a Tinius Olsen Model 74 (358 J capacity) impact machine, respectively; both machines were provided with a Dynatup Model 500 instrumentation system. The P-t traces captured by an analogue storage oscilloscope were photographed using a Polaroid camera for subsequent analysis. Drop-weight and impact tests were performed following the ASME Section III, Div. I guidelines4 and ASEM E 208’ and E 239 procedures and other practices relevant to instrumented impact and drop-weight testing.‘3,‘4 For above-ambient temperature tests, a heated quenching oil bath was used, and for below ambient temperature tests a bath consisting of a methanol-liquid nitrogen mixture was used following the ASTM E 208’ or E 239 requirements as applicable.
* Interpass temperature was maintained below 533K.
for transverse to the weld bead, L for along the weld bead and S in the thickness direction to the weld bead) as required by ASME Section III, DIV. I, NG-2322*2.4 P-3 DW and standard CVN specimens were prepared following the relevant ASTM standards.8.9 Following ASTM E 208 specifications,* to reduce the extent of th.e HAZ, short-weld beads (approximately 25 mm long and 13 mm wide) were deposited on the P-3 specimens using 5 mm diameter BOR-C (a Cr-MO-V air-hardening steel giving approximately a Rockwell C hardness HRc of 52) hard-facing electrode supplied by M/s D&H Secheron Electrodes Pvt. Ltd, Indore. Suitability of these electrodes for preparing drop-weight specimens had been verified earlier.” Also to reduce the HAZ effect further, following some of the latest research reports,““* the specimens were placed on a ‘chiller block’ (approximately 150 X 150 X 150 mm steel block) when depositing the hard-facing weld-bead; use of this chiller block is supposed to reduce the extent of
2.4 RTNDT determination Following the ASME Section III, Div. I guidelines, for each weld-pad (i.e. corresponding
Table 2. Chemical Chemical
compositions and mechanical properties of 9Cr-1Mo prepared using different diameter electrodes composition (wt%)*
Electrode diameter (mm) 2.5 3.15 4.0 Mechanical
C
Mn
Si
testing
Cr
Ni
MO
v
0.10 0.93 0.39 9.11 0.25 0.87 Nil 0.09 0.90 0.43 9.21 0.24 0.87 Nil 0.087 095 0.46 9.08 0.25 0.92 Nil
cu Nil Nil Nil
s
welds
P
0.004 0.004 0.004 0.004 0.003 0.004
properties?
Electrode diameter (mm)
O-2% proof stress(MPa)
LJTs (MPa)
2.5 3.15 4.0
551.7 579.9 548.2
695.6 699.1 712.7
% Elongation Charpy V-notch impact (5.65VA) energiesat 293K (km) 20.3 18.9 19.0
13, 10.4, 11.2 13.9, 9.0, 8.4 10.0, 8.6, 8.2
* Balance-Fe. t All welds were stress relieved at 1003f 5K for 1 h with rates of heating and cooling 150K/h max. Also, prior to welding, all electrodeswere baked for 1 h at 673K giving moisture contents of 0.205, 0.310 and 0.17% in the welds using the 2.5, 3.15 and 4mm diameter electrodes, respectively. Tensile properties are for tests at room temperature.
P. R. Sreeniuusanet al.
152
to each diameter electrode), TNDT was determined. Then, CVN specimens were tested in triplicate at successively higher temperatures, at 5K intervals, starting the first batch of tests at TNDT+ 33K. According to the ASME Section III, Div. I, NG 2330 and NG 2350 requirements on Charpy test values,4 for acceptance, the average value from the triplicate tests should meet the minimum requirements; more than one specimen should not show values below the minimum requirements and, for the specimen not meeting the minimum requirements, the C, and LE values should not be below the minimum specified requirements by more than 10 ft-lb (13.4 J) and 5 mils (O-127 mm), respectively: in cases not meeting these criteria, two additional tests performed should both meet the minimum requirements.
3 INSTRUMENTED ANALYSIS
L
P t
I
t
TYPE-I
P max = PF
TEST DATA
General procedures for analysing instrumented test data are given in Server and co-workers’3”4 and Varga et al.” 3.1 Drop-weight
P max = P,
TYPE
-II
pGY
test data analysis
By assuming that the elliptical thumbnail formed by the HAZ of the weld-bead to be the starter crack, and using the stress-intensity factor analysis of Scott and Thorpe,16 Sreenivasan et al. ‘O and Moitra et aLI7 estimated Kid from P-t traces of ASTM E 208 drop-weight tests at and below TNDT. The same procedure has been adopted here. The results on crack-profile and Kid are given in Section 4. 3.2 Impact test data analysis 3.2.1 General The P-t traces from instrumented tests of CVN specimens can be categorised into types I, II and III as shown in Fig. 2; the general yield, maximum and fast/brittle fracture loads are represented by Pcy, Pm,, and PF, respectively. In type I, brittle or fast running fracture is characterised by a sudden drop in load occurring before or at general yield, indicating linear-elastic material behaviour; this is the situation in the
TYPE-III Fig. 2. Three types of load-time (P-t) traces from instrumented impact tests of Charpy specimens.
lower shelf of the transition region of ferritic steels. In type II, fast fracture occurs after general yielding at (PF = Pm,,) or after the maximum load (PF < P,,,,,); at higher temperatures in the transition region, fast fracture arrests after some propagation and this is reflected in the P-t traces by a tail after some sudden drop. Type III refers to complete ductile fracture with no sudden load drop, signifying fast fracture, and occurs in the upper shelf region. For reliable interpretation of P-t traces from instrumented impact tests, the time to any load measurement point, t, should be greater than 2.32 (or more conservatively 32), where r is the apparent period of oscillation of the three-pointbend (3PB) specimen.‘3,‘4 Moreover, to avoid
Properties of a 9Cr-1Mo weld excessive signal attenuation, f should be greater than or equal to l.lT,, where T, is the response time of the instrumentation system. The value of TR in the present tests was mostly 90 ps (in some tests it was 10 ps), yielding a value of 100 (or 11) ps for l.lT,. The 32 period calculated for CVN steel specimens was about 100 ps: this corresponds to the first three oscillations on the P-t traces. The dynamic yield stress, (TYd, was computed from PGy using the following relation for a standard CVN specimen:13 cr,,(MPa) = 46*7P,,(kN)
(1) The cleavage fracture strength or micro-cleavage fracture stress, nf, was obtained from: o,(MPa) = C,cr,,(MPa)
(2) where (Tyd is the dynamic yield stress at the brittleness transition temperature (T,) at which PF= Pcy (see load-temperature (P-T) diagram results and discussion given later) and the stress intensification factor C, = 2.57 for the standard CVN specimen at T,.” 3.2.2 Kid estimation procedures Krd was estimated using the ASTM E 399’ formula for precracked 3PB specimens taking a/W = O-2 and the final expression for standard CVN specimen is: K,, = 4.67P,
(3) where PF is in kilonewtons. An error analysis for K,, reported elsewhere’0,‘7 shows that ,the error band can be taken as &20%. For type I cases, the application of a reduction factor of 0.8 (the lower bound of the 20% error band) to the results from CVN specimens using eqn (3) can be expected to give results in the error band of precracked Charpy K1d tests.” Wherever the P-t traces showed time-to-failure, tf, between z and 22, PF should be replaced by PC, a critical fracture load obtained by Varga’s procedure.‘5.‘0 Turner,*’ based on beam-on-foundation analysis of instrumented CVN tests, has given graphical correction factors to be applied to PF derived from P-t traces showing only one or two oscillations. In the present case, these inertia correction factors were found to be smaller than 20% of those obtained by Varga’s procedure. For type II and III test traces (Fig. 2), an elastic-plastic procedure has to be applied. Usually such traces have been analysed using a
153
J-integral or crack tip opening displacement (CTOD) procedure assuming that crack initiation occurs at the maximum load.*‘,“* For impact tests of CVN specimens of ferritic steels, Ghoneim and Hammadz3 proposed that fracture initiation occurs at a load equal to (Pm,, + P,,)/2, while Norris,24 based on a comparison of finite element results with experimental data, reported that the time to crack initiation, ti, equals 40% of the time to reach the maximum load, t,,,. For CVN specimens of ASTM A 533 Class 1 steel, Kobayashi” has reported a value of 0.8 for the ratio of initiation to maximum load energy. In the present case, taking tiltmax = O-4 gave slightly smaller values for the initiation energy compared with Ghoneim and Hammad’s procedure and much smaller values compared wtih Kobayashi’s procedure. Hence, J or CTOD estimates were computed assuming initiation to occur at a tiltmax value of O-4. The average velocity, V, and the hammer travel (i.e. load point displacement, LPD), d, at any time, t, during impact were obtained using the usual relations:21~22~26 d = tV = tVo(l - E,/4E,)
(4) where V, is the velocity of the pendulum at the point of impact, E, is the initial pendulum energy and E, = V,J P dt, J P dt being the total area under the P-t trace. From this the plastic part of the LPD, d,, is obtained by subtracting d,,, the total system elastic deflection given by:13 d,, = C,P
(54
where P is the load at t (corresponding to d) and C, is the total system compliance given by:13 CT = (VohxIPGY> - W2,,l@Eo)
w
where the subscript GY signifies values at general yield. For a standard CVN specimen (the same holds for a half-thick CVN specimen, for the relation is independent of thickness), following the derivation in Mutoh et al.,*’ and assuming that the 3PB specimen during loading undergoes rotation by a plastic hinge mechanism with a rotational factor of 0.4, the relation between the plastic part of CTOD, apl, and d,, had been given as:*’ S,, = 0.32d,,
(64 However, according to more recent results’8Jy for shallow cracks (a/W = 0.1%0.3) r values are in
P. R. Sreenivasan et al.
154
the range 0.2-O-3, while for deeply cracked specimens Y = 0.4-0.45. Hence, in the present paper taking Y = O-2 modifies eqn (Sa) to: 6,, = O.l6d,’
w
Then total CTOD at initiation, 6i, is given as the sum of the elastic and plastic parts:29
Table 3. TNDT, RTNDT and TD transition-temperatures and cleavage fracture strength (~3 for 9Cr-1Mo welds
Electrode diameter (mm) 2.5 3.15 4.0
TNDT Gv G-33) WI (J-9 WI 263 297 268 302 268 301
R&m W
264 269 268
264 269 268
TD W ($4 264 2112 255 2166 -
6i = 6, + Spl = (K2(1 - u*)/(2a,,E))
+ 6,,
(7)
where K is calculated using eqn (3) wtih P = Pi at ti, oys is replaced by cyd and Young’s modulus E = 190GPa for the range of temperatures involved. Then Kid(S) is given by Kld(S)
=
d(EaYd6i)
(8)
The value of Ji at initiation is determined from the following formula due to Rice et aZ.:30 Jj = n Ei/(Sb)
Pa>
where n = 2 for deep-cracked 3PB specimens, B is the thickness, b is the remaining ligament depth (= W - a) and Ei is the total energy (excluding the machine contribution) up to initiation. Based on a critical consideration of the expressions given in Sumpter,28 for a specimen with a/W = 0.2, n can be taken as 1.45. Then J, = 1’45Ei/(Bb)
cw
In eqns (9a) and (9b), Ei is obtained from the total absorbed energy E as follows:‘3 Ei = E - (P2/2) * [CT - CND/(EB)]
(11)
In the literature Kid has been estimated by the Charpy energy (C,) correlation approach.3’ One of the most successful correlations applicable over a wide range of C, (3-95 J) and room temperature static yield strength (gys = 270815 MPa) values is the following: K’,(C”) = 15*5(cv)“~37s
AND DISCUSSION
4.1 RT,,, Table 3 reports the RTNDr results; clearly, for all the welds RT,,, agrees with the respective TNDT within lK, which is within the accuracy of temperature measurement. These TNDT values are worth comparing with those reported for a modified 9Cr-1Mo plate steel in the normalized and tempered condition: 268.6-279*7K.32 In the plots of Kid versus RTNDT reported later, RT,,, of welds I, II and III have been taken as 263,268 and 268K, respectively (i.e. equal to the respective TNDT values). The lower RTNDT (i.e. tougher behaviour) of weld I is in line with the tensile and impact results shown in Table 2. This may be due to the greater refinement in structure occurring in the weld prepared using a smaller diameter electrode. Since the resolution of TNDT is only 5K, referring to Table 3, the RTNDT of all the three welds can be conservatively taken as 269K.
(10)
where CND is the non-dimensional specimen compliance given in ServerI (equal to 24.39 for a standard CVN specimen) and P is the load at t (equal to ti in present case). Then Kid(J) = q(EJi)
4 RESULTS
(12)
The above was used in the present paper to estimate Kld(Cv) from C, values.
4.2 Load-temperature
diagram
Load-temperature (P-T) diagrams for welds are shown in Fig. 3. Paucity of specimens did not permit testing of CVN specimens of weld III (i.e. prepared using 4 mm diameter electrodes) at low temperatures. The results show considerable scatter, hence there is likely to be more uncertainty in the TD values determined. However, the loads corresponding to PF = PGy are not likely to be in error by more than lo%, which is only within the uncertainty of instrumented impact test load values. T, and gf values determined are reported in Table 4. The values of 2112-2160 MPa for the (T’ are reasonable, but are much lower than the reported values of 2550-2850MPa for a normalised and tempered plate metal. 33 This indicates that the fracture
Properties of a 9Cr-1Mo 9Cr-1Mo 20
2.5mm
WELD
Q ELECTRODE
/
/
A-
10
1
173
B 1
1
1
1
1
1
223
1
1
1
273
1
PGY
’
1
1
1
323
T/K Fig. 3. Load-temperature diagrams from instrumented impact tests of Charpy V-notch specimens of 9Cr-1Mo welds I and II (2.5 and 3.15 mm diameter electrodes, respectively).
resistance of the welds is poorer than that of the base metal.
Kid values estimated by various procedures and the computed gyd values are reported in Table 4. For drop-weight tests, crack profile measurements on fracture surfaces of broken specimens yielded average values of 0~241(*0~037) and O-296 (*O-056) for (a/2c) and (a/t), respectively (values in parentheses are two standard deviation values; n and c are the semi-minor and semi-major axes of the elliptical thumbnail crack bounded by the HAZ of the crack-starter weld-bead on the broken specimen surface and t is the specimen thickness, respectively). The above ratios compare very well with the values of (a/2c) = 0.25 f 0.02 and (a/t) = 0.27 f 0.04 reported for an AISI 403 stainless steel.” Final fOrIIIUkE for Kid from drop-weight tests Of P-3 specimens obtained in the present test campaign were as given below: K,,(MPadm) = 0.41Pcdu at TNDT! = 0-38P,~u at T < ThIDT! (13)
155
weld
where PC (kN) is the critical fracture load derived using Varga’s procedure” and a is in millimetres; eqn (11) shows excellent agreement with the expressions reported in Sreenivasan et al.;” these Kid estimates are accurate to &20% and have been reported as Kld(DW) in Table 4. The K&%d ratio at TNDT evaluated from crack profile measurements on broken specimens is O-074 f 0,016 drn and is in good agreement with that reported in the literature.” The various Kid values reported in Table 4 have been plotted against (T - RTNDT) in Figs 4-6. The ASME KIR curve4 has also been plotted for comparison. An error band of *20% was applied to K,,(DW), Kid (PJP,) and Kld(CV) for reasons mentioned earlier. The error bands for Kid(8) and Kid(J) were taken as *25% and &35%, respectively (see the Appendix). From Figs 4-6, it is evident that Charpy correlation K,, values (Kld(CV)) are not conservative with respect to the ASME K,, curve. It must be borne in mind that most of the Charpy correlations reported have been based on results for carbon and low-alloy steels. As such, the applicability of these correlations to high alloyed steels like 9Cr-lMo, 12Cr-steels, etc. requires verification. James and Carlson3’ had applied C,-K,, correlation and a temperature shift procedure to obtain the K,, versus temperature curve of a modified 9Cr-1Mo steel; a single KIc determined at the upper shelf was much higher than that predicted by the above method. In the present instance also it is likely that the KId(CV) values are unduly pessimistic. The lower bound Vahes from all the other Kid estimates plotted in Figs 4-6 are conservative (i.e. give higher values) with respect to the ASME K,, curve over the range of temperatures covered in this investigation: 193-303K. It must be noted that taking the R Lx- as 269K for all the welds (as suggested earlier under the discussion on RTNDT) would increase the conservatism of the ASME KIR curve with respect to the K,, estimates. 4.4 Reliability
of Kid values
Since the available Charpy correlations seem to give unduly pessimistic values, it is desirable to develop more reliable correlations between C, and K,c/K,, results based on proper tests (proper according to established and accepted standards and procedures for Ki,-/Kid determination).‘3-‘5 The alternative procedures for estimating K,,
P. R. Sreenivasan et al.
156 Table
Electrode diameter (mm) 4
19.5 223 268 293 297 297 302
3.15
191 226 253 263 268 273 300 302
2.5
191 227 253 263 273 293 297
4. Dynamic
yield strength
(uYd) and various K,, estimates for 9th1Mo
welds
UYd
K,,(J)
K,.,(S)
KdP~lf’c)*
KM(G)
f MW
(MPadm)
(MPadm)
(MPagm)
(MPaqm)
(MPat/m)
-
67.6 65.8 80.2 80.6
56.4 54.6 744 -
805.1 797.6 786.3 783.4 786.9 792 748.2 780.6 779.4
KdDV
157 169.6 181 180.7
93.4 94.6 99.2 96.8
142.5 195-5 184
88.2 117.4 98.1
52.21 68.2$ 81.2 -
24.11 28.33 36.7 59.3 75 80.4
61.4 56.3 61 76 70.8 -
195.6 195.1
109 105
75.21 86.48 77.5 61.5 -
28.3t 33%$ 36 47.7 77.1 81.2
45.7 56 67 -
*K,, from PF/Pc obtained from linear-elastic load-time trace from instrumented test of unprecracked Charpy specimen usingASTM E 399 formula (seetext for further explanation). ‘r Resultsat 193K. $ Results at 223K.
from instrumented test P-t traces of CVN specimens discussed above are promising when used with sufficient engineering judgement. The lower bound values of K,,(DW) give reliable estimates of Kid. In most cases, K,, from PF/Pc of CVN specimens giving type I P-t traces give values in agreement with K,,(DW). A minimum of three or four CVN specimens should be tested at each temperature and over a range of temperatures and a lower bound trend line taken. Appropriate statistical procedures, though not applied in the present case, are available to account for scatter.3b36 The test temperatures should not be too low to produce type I P-t traces with only a single inertial load peak that violates the 22-32 requirements. In the case of type II and type III P-t traces, it is seen that Kid(S) is less than K&). However, Duffy et d3’ state that crack opening displacements (CODS) are not meaningful in an absolute sense when the material is very tough. For type II traces, up to temperatures where PF = Pm,,, K&6) can be taken as a conservative estimate of K,,. In making use of these estimates, a smooth lower bound curve may be drawn. ‘The error bands
used in this paper and discussed in the Appendix are based on three or more weld CVN specimens tested at each temperature. Application of these error bands to homogeneous specimen results can be expected to give conservative values. Modification of these error bands may be done based on further CVN tests and comparison with precracked Charpy tests.
5 CONCLUSIONS 1. The RT,,* of 9Cr-1Mo welds prepared using electrodes of diameter 2.5, 3.15 and 4 mm agree within 5K; the respective values being 264, 269 and 268K; these are well within the 5K resolution expected for TNDT. Hence, with utmost conservatism, the RTNDT of all the three welds can be taken as 269K. 2. The procedures presented in this paper enable estimation of reasonably conservative values of K,, from instrumented impact tests of unprecracked Charpy V-notch specimens. 3. For the range of temperatures over which
of a 9Cr-1Mo
Properties
9Cr-1Mo
200 9Cr-1Mo
200
150
WELD:
2Smm
0 ELECTRODE
v
K,.,
FROM
DROP
WEIGH1
0 A
Ytd
FRQM
PF
K,,,
FROM
6
El m
Kid FROM J LOWEST VALUES IN THE ERROR BAND OF KldtJl
ta
K,,, FROM
1.57
weld
130
TEST
WELD:
0 ELECTRODE
v A
Kid FROM 6 K,* FROM J LOWEST VALUES IN THE ERROR BAND OF Kld(Jl K,, FROM Cy CORRELATION
q q 0
FROM
&mm
Kid
DROP WEIGHT
TEST 88
El
150
CV CORRELATION
i
ki 0 &
100
\ 0 Y
r 50
LASME
L ASME
K,,
0 0
I
I
I
I
I
100
I
I
I
223
I
I
I
I1
1
I
RT,,,VK
Fig. 4. Variation of K,, with (T-RTNDT) for 9Cr-IMo weld I (2.5 mm diameter electrode).
tests were carried out (193-303K), the lower bound K,, estimates obtained for thLe 9Cr1Mo welds were higher than the ASME K,, 200
-
9Cr-1Mo
-
v
-
-
150
Kid
WELD: FROM
3.15mm
0
Kid
FROM
PF
A
Kid
FROM
6
FROM
J
0
Kid
q
LOWEST VALUES IN THE ERROR BAND OF K,,,(J)
I3
K,.,
-
FROM
d ELECTRODE
DROP WEIGHT
El
TEST
q
Cv CORRELATION
CURVE
El
“‘([“I”““’ 173
323
273
(T-
K,,
CURVE
T
223
273 (T-
RT,,
323
11 K
Fig. 6. Variation of Kid with (T-RTNDT) for 9Cr-1Mo weld III (4 mm diameter electrode).
curve. Hence the ASME KIR curve provides a conservative lower bound over these temperatures. For this material, applicability of the ASME K,, curve at higher temperatures needs validation. 4. Microcleavage fracture stress, gf, estimated for the 9Cr-1Mo welds is about 2160MPa; this is much lower than the 2550-2850 MPa reported for a normalised and tempered 9Cr-1Mo plate material, and indicates poorer fracture resistance for the welds.
REFERENCES 100
1. Hudson, J. A., Druce, S. G., Gage, G. & Wall, M., Thermal aging effects in structural steels. Theor. Appl. Fracf. Mech, 10 (1988) 123-33. 2. Fidler, R. S., The creep of normalised and tempered 9Cr-1Mo: effects of stress,temperature and grain size on creep and rupture properties. CERL Report No. RD/L/R 1949,1976. 3. Fidler, R. S. & Middleton, C. J., The impact and hot tensile properties of 9Cr-1Mo steel in various heat treatment conditions. In Proc. IAEA Specialists
50
173
223
273
323
(T-RT,DT)/ K Fig. 5. Variation of K,, with (T-RTNDT) for 9Cr-1Mo weld
II (3.15 mm diameter electrode).
Meeting-Mechanical Properties of Structural Materials Including Environmental Effects, Chester,UK, 1983,pp.
519-44. 4. ASME Boiler and PressureVessel Code, Section III Nuclear Power Plant Components,Division 1, Appendix G. ASME. New York, 1989.
158
P. R. Sreenivasan
5. Lidbury, D. P. G. & Morland, G., Review of fracture toughness requirements and data relevant to LWR reactor pressurevessels,Znt. J. Pres. Ves. & Piping, 29, (1987) 343-428. 6. Test Certificate SerialsNos 18474,19228and 18471,dtd. 05-01-1993 from M/s. Advani Oerlikon Limited, Bombay on CITOCHROM-9R SFA-5.4, E 505-15. 7. ASTM E 399-83,Standard test method for plane-strain fracture toughnessof metallic materials. 1990 Annual Book of ASTM Standards, Vol 03.01, 1990, pp. 488-512. 8. ASTM E 208-87a,Standard test method for conducting drop-weight test to determine nil-ductility transition temperature of ferritic steels. 1990 Annual Book of ASTM Standards, Vol. 03.01, 1990,pp. 360-71. 9. ASTM Standard E 23-88,Standardmethodsfor notched bar impact testing of metallic materials. 1990 Annual Book of ASTM Standards, Vol. 03.01, 1990, pp. 197-212. 10. Sreenivasan, P. R., Ray, S. K., Mannan, S. L. & Rodriguez, P., Determination of K,, at or below NDTT using instrumenteddrop-weight testing, Znt. J. Fract. 55 (1992) 273-83.
11. Holt, J. M. & Puzak, P. P. (eds), Drop-weight test for determination of nil-ductility transition temperature of ferritic steels:user’sexperience with ASTM method E 208. ASTM STP 919, American Society for Testing and Materials, 1986. 12. Satoh, M., Funada, T. & Tomimatsu, M., Evaluation of valid nil-ductility transition temperatures for nuclear vesselsteels.ASTM STP 919, pp. 16-33. 13. Server, W. L., Impact three-point bend testing for notched and precracked specimens,J. Test. Eval., 6 (1978) 29-34. 14. Ireland, D. R., Server, W. L. & Wullaert, R. A., Proceduresfor testing and data analysis,effects. Tech. Inc., Technical Report, TR 75-43, 1975. 15. Varga, T., Junker, M., Njo, D. H. & Prandtl, G., Proposedmethod for instrumented precracked Charpy type tests: ASK-AN-425, Rev 1, December 1978. In C.S.N.I. SpecialistMeeting on Instrumented Precracked Charpy Testing. EPRI NP-2102-LD, November 1981, pp. 1-85-l-115. 16. Scott, P. M. & Thorpe, P. W., A critical review of crack tip stress intensity factors for semi-elliptic cracks, Fatigue Eng. Mater. Struct., 4 (1981) 291-309. 17. Moitra, A., Sreenivasan,P. R., Ray, S. K. & Mannan, S.
L., Instrumented drop-weight test resultsfor normalised and tempered 9Cr-1Mo steel, communicatedto Znt. J. Fracture (1995). 18. Ray, S. K., Sreenivasan, P. R., Samuel, K. G. &
Rodriguez, P., Characterisation of irradiation-induced embrittlement in an ASTM A 203D steel using instrumentedimpact testing, Znt. J. Pres. Ves. h Piping, 54 (1993)481-96. 19. Krompholz, K., Tipping, P. & Ullrich, G., ‘Resultsfrom investigationswith an instrumentedimpact machine on a molybdenum base alloy, nickel base alloys, and Incoloy 800’, Z. Werkstoftech., 15 (1984) 117-23. 20. Turner, C. E., Measurement of fracture toughnessby instrumented impact testing, Impact Testing of Metals, ASTM STP 466 (1970) 93-114. 21. Samuel, K. G., Sreenivasan, P. R., Ray, S. K. &
Rodriguez, P., Evaluation of ageing-inducedembrittlement in an austenitic stainlesssteel by instrumented impact testing, J. Nucl. Mater., 150 (1987) 78-84. 22. Iyer, K. R. & Miclot, R. B., Instrumented Charpy
et al. testing for determination of the J-integral, Instrumented Zmpact Testing, ASTM
STP 563 (1974) 146-65.
23. Ghoneim, M. M. & Hammad, F. H., Instrumented impact testing of an irradiated 20MnMoNi55 PVS weld material, J. Nucl. Mater., 186 (1992) 196-202. 24. Norris, D. M., Jr, Computer simulation of the Charpy V-notch test, Eng. Fract. Mech, 11 (1979) 261-74. 25. Kobayashi, T., Analysis of impact properties of A533 steel for nuclear reactor pressurevesselby instrumented impact test. Eng. Fract. Mech., 19 (1984) 49-65. 26. Ireland, D. R., Proceduresand problemsassociatedwith reliable control of instrumented impact testing, Instrumented Impact Testing, ASTM STP 563 (1974) 3-29.
27. Mutoh, Y., Toyoda, M. & Satoh, K., On the relation between load-point displacement and stretch zone width, Znt. J. Fract., 16 (1980) R171-4. 28. Sumpter, J. D. G., Jc determination of shallow notch welded bend specimen, Fatigue Fract. Eng. Mater. Struct., 10 (1987) 479-93.
29. Sorem, W. A., Dodds, R. H., Jr & Rolfe, S. T., Effects of crack depth on elastic-plastic fracture toughness,Znt. J. Fract., 47 (1991) 105-26. 30. Rice, J. R., Paris, P. C. & Merkle, J. G., Some further resultson J-integral analysisand estimates,Progress in Flaw Growth and Fracture STP 536 (1973) 231-63.
Toughness
Testing,
ASTM
31. Roberts, R. & Newton, C., Interpretive report on small-scaletest correlation with KIC data, Welding Research Council Bull. 265, 3 February 1981. 32. James, L. A. & Carlson, K. W., The fatigue-crack growth and ductile fracture toughness behaviour of ASTM A387 Grade 91 steel, J. Pres. Ves. Tech. (Trans ASME),
107 (1985) 271-8.
33. Packiaraj, C. C., Sreenivasan,P. R., Moitra. A., Ray, S. K., Mannan, S. L. & Arivazhagan, B., Effect of simulated post weld heat treatment on toughnessand transition temperature characteristicsof normalisedand tempered 9Cr-1Mo ferritic steels. Presented at the Poster Session of 49th ATM, IIM, Calcutta, 14-17 November 1995. 34. Bishop, T. A., Markworth, A. J. & Rosenfield, A. R., Analysing statistical variability of fracture properties, Metall.
Trans., 14A (1983) 687-93.
35. Oldfield, W., Fitting curves to the toughnessdata, J. Test. Eval.. 7 (1979)326-33. 36. Koons, G. ‘F., ‘Statical analysisof notch toughnessdata, J. Test. Eval., 7 (1979) 96-102. 37. Duffy, A. R., Eiber, R. J. & Maxey, W. A., Recent work on flaw behaviour in pressurevessels.In Practical Fracture Mechanics for Structural Steel, Proc. Symp. on Fracture Toughness Concepts for Weldable Structural Steel, Risley, UK, ed. M. 0. Dobson. UKAEA and
Chapmanand Hall, Risley, 1969pp. Ml-M34. 38. Burdekin, F. M., Crack opening displacement-a review of principlesand methods.ibid. pp. Cl-C12. 39. Heerens,J., Cornet, A. & Schwalbe,K.-H., Resultsof a round robin on stretch zone width determination, Fatigue Fract. Eng. Mater. Struct. 11 (1988) 19-29.
40. Sreenivasan, P. R., Ray, S. K., Vaidyanathan, S. & Rodriguez, P., Measurementof stretch zone height and its relationship to crack tip opening displacementand initiation J-value in an AISI 316 stainless steel, communicated to Fatigue Fract. Eng. Mater. Struct. (1995). 41. Broek, D., The Practical use of Fracture Mechanics. Kluwer, Dordrecht (1989).
Properties oi ~1K’r-1Mo
APPENDIX: ERROR BANDS VARIOUS K,, ESTIMATES
FOR
weld
‘lherefore 2SKIK
The basis for assuming an error band of +20% for Kld(DW), Kld(PF/PC) and Kld(CV) estimates has sufficient literature support.‘0,38 In the following an error analysis has been given for Kid(8) and K&) estimates. An assumption of &40% for the error in 6 is reasonable;3x.3Qresults on 6 for a 316 stainless steel obtained from stretch zone measurements4’ show a two standard deviation varying from 28 to 54%, with most of the values lying in the range 40-45%. Similarly, an error of 50-70% for J is reasonable.4’ Moreover, J estimates in the present tests, based on three or more Charpy tests of welds in the same condition at each temperature, yieldled two standard deviation values in the range 25-70%. So an error value of 70% for J has been taken in the following analysis. K(S) = v(EaydS)
159
(Al)
= 6EIE
+ 6ay,/a,,
+ &S/6
(A2)
Taking error values of 10% for Young’s modulus E and flyd is reasonable. Then the root mean
square error in K from eqn (A2) computes to &21%. Therefore, an error of *25% has been assumed for Kid(a). It must be noted that the error bands taken are assumed to account for notch sharpness effects also. Similarly: K(J) = q(EJ)
WI
Then 26KIK
= 8ElE
+ SJlJ
Thus the root mean square error computes to +35%.
(-w
in Kid(J)