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Evaluation of fracture resistance of AISI type 316LN stainless steel base and welded pipes with circumferential through-wall crack
International Journal of Pressure Vessels and Piping 178 (2019) 104008
Contents lists available at ScienceDirect
International Journal of Pressure Vessels and Piping journal homepage: http://www.elsevier.com/locate/ijpvp
Evaluation of fracture resistance of AISI type 316LN stainless steel base and welded pipes with circumferential through-wall crack S.A. Krishnan a, *, R. Nikhil a, G. Sasikala a, A. Moitra a, Shaju K. Albert a, A.K. Bhaduri a, C. Lakshmana Rao b, S. Vishnuvardhan c, M. Saravanan c, P. Gandhi c, G. Raghava c a b c
Indira Gandhi Centre for Atomic Research, HBNI, Kalpakkam, Tamil Nadu, India Department of Applied Mechanics, Indian Institute of Technology Madras, Chennai, India Structural Engineering Research Centre - CSIR, Chennai, India
A R T I C L E I N F O
A B S T R A C T
Keywords: Fracture experiment Pipe Pipe weld C(T) specimen J-R curve Limit load
Fracture resistance of 316LN SS base and welded straight pipes with circumferential through-wall cracks has been evaluated. The pipe J-R curves have been determined using existing experimental methods and analytical expressions available in literature. The J-R curves of 88.9 mm OD pipe and pipe welds have been compared with the J-R curves obtained from C(T) specimen tests. The results indicate that the C(T) specimen data is not suitable for fracture assessment of pipes due to difference in constraint level and the fracture resistance of the base pipe is inferior to welded pipe. Further, the limit load analyses based on expressions available in literature reveals that the initiation and crack propagation are observed prior to the maximum bending moment.
1. Introduction AISI Type 316LN stainless steel (316LN SS) is used as the major piping material for liquid sodium cooled fast breeder reactor programme in India. Due to exothermic affinity of sodium for water, sodium-water/ steam reaction may occur in the event of leakage from high pressure steam generator tubes. This reaction leads to high-pressure waves and corrosive products, which may affect the structural integrity of inter connecting piping/components of secondary sodium main circuit, Sumathi et al. [1]. Further, any leakage due to undetected crack in piping may react with air and produce corrosion and remain undetect able while progressing and finally leads to rupture when an incidental loading gets imposed on the normal load. For any initial defect in the sodium circuit, either undetected or undetectable by the available methods of inspection, one must be able to compute the growth of defect/crack during plant operation, taking into account “design basis event” loads which includes normal, upset, emergency and faulted conditions. In spite of having well established procedures [2] for assessment of fracture behavior from laboratory scale specimens like C(T) and three point bend specimens; the applicability of these data for integrity assessment of piping components still remains a challenge due to limited crack growth data. This has been addressed by many researchers for the
piping components [3–7]. Experimental methods to evaluate the J-R curve for straight pipes have been established by Zahoor and Kanninen €rster [3]. For 316L pipes, the fracture resistance has been evaluated by Fo et al. [4]. The differences in J-R curves obtained from pipes and C(T) specimens have been reported. Similar observations were also made by Moulin and Delliou [5] and Chattopadhyay et al. [6] for circum ferentially through-wall cracked 316L and SA 333 Gr.6 carbon steel pipes respectively. This observation also holds good for 304LN SS as reported by P. K.Singh et al. [7]. Overall, this is related to the differences in crack tip constraints for the C(T) and piping components. The influ ence of crack tip constraint on material fracture toughness at specimen level has been discussed in Tatafder et al. [8] and the comparison of fracture toughness from laboratory specimens with the fracture resis tance of components for SA 333 Grade 6 carbon steel has been discussed in Pavankumar et al. [9]. To the author’s knowledge, for the 316LN SS, which can be classified as a low yield stress but high hardening material, the fracture results for the pipes and pipe welds, is not available in open literature. Further, the transferability problem could be more prominent due to following reasons: (1) most of the piping components used in sodium cooled fast breeder reactor are thinner than the minimum thickness required to evaluate the fracture resistance from standard specimens, (2) limited crack extension data (1.5–3 mm) is available from fracture tests on laboratory specimens, whereas the crack growth
* Corresponding author. E-mail address: [email protected] (S.A. Krishnan). https://doi.org/10.1016/j.ijpvp.2019.104008 Received 20 March 2019; Received in revised form 17 October 2019; Accepted 27 October 2019 Available online 31 October 2019 0308-0161/© 2019 Elsevier Ltd. All rights reserved.
S.A. Krishnan et al.
International Journal of Pressure Vessels and Piping 178 (2019) 104008
Table 1 Chemical composition in wt (%). Base
C
Si
Mn
Cr
Ni
N
Mo
P
S
Fe
0.027
0.22
1.7
17.5
12.22
0.07
2.49
0.01
0.01
Bal.
seamless and welded specimens. The chemical composition and me chanical properties of 316LN SS are presented in Table 1 and Table 2. The details of specimens are shown in Table 3.
Table 2 Mechanical properties at room temperature. Material
Yield strength (YS), MPa
Ultimate tensile strength (UTS), MPa
Modulus of elasticity (E), GPa
Poisson’s ratio (μ)
316LN SS Base
250
570
200
0.3
316LN SS Weld
420
665
200
2.2. C(T) specimen A 30 mm thick weld pad with butt weld groove of total 20� bevel angle and 14 mm root gap was prepared using multi-pass SMAW pro cess. X-ray radiography has been carried out to ensure the welds are free from detectable defects like cracks, inclusions etc. The C(T) specimens are fabricated as per sketch shown in Fig. 2 (a & b). In case of weld specimen, the crack plane is kept parallel to the weldment. The notch was introduced by electric discharge machining process at the weld centre line (see Figure 2).
0.3
analysis requires large extensions (40–50 mm) for components and (3) scatter in fracture test data of welded specimens. Hence, to ensure the integrity of 316LN SS piping and fracture transferability, a detailed fracture analysis of various components, e.g. straight pipes, elbows and branch tees have been proposed in parallel with the conventional lab oratory specimen tests. The present work is a part of experimental and numerical studies being conducted at IGCAR Kalpakkam to establish the transferable fracture parameters from specimens to components. Experiments have been carried out to investigate the fracture behavior of through-wall cracked 316LN SS base and welded straight pipes. The J-R curves evaluated from pipes have been compared with that from laboratory C (T) specimens. Further, the comparison between experimental maximum load and analytical limit load has been presented.
3. Experiments 3.1. Fatigue pre-cracking Fracture tests are conducted on the fatigue pre-cracked base and weld pipe specimens at room temperature. Similar type of tests on pipes and its analysis have been reported by several researchers [3–7,10,11]. The fatigue pre-cracking of the specimens was carried out under four-point bending to produce sharp crack front as shown in Fig. 3.
2. Material properties and specimen details 2.1. Pipe specimen The seamless pipe specimens were fabricated from 316LN SS pipe of 80 NB, Sch. 40 category (OD 88.9 mm � 5.49 mm thick). The circum ferential through-wall notch was introduced at mid span of length. The girth welded pipe specimens are prepared using 16-8-2 filler wire for root pass and modified E316-15 electrode for subsequent passes with weld configuartion as shown in Figure 1. The root passes are made by gas tungsten arc welding (GTAW) procedure and subsequent passes by shielded metal arc welding (SMAW) procedure. X-ray radiography of the welds has been carried out to ensure absence of weld defects. Circumferential U-notch of root radius 1.5 mm has been introduced in
Fig. 1. Details of pipe weld configuration.
Table 3 Experimental details of pipe specimens. Specimen ID
SP4-90TWC-M1 SP4-120TWCM2 SP4-150TWCM3 SP4-60TWC-M4 SP4-60TWCWM5 SP4-90TWCWM6 SP4-150TWCWM7 SP4-120TWCWM8
Crack location
Inner Span (mm)
Outer Span (mm)
Fatigue pre-cracking load (kN)
Frequency (Hz)
Max. (Pmax)
Min. (Pmin)
10 7.5
1 0.75
1.0–2.0 1.0–2.0
5
0.5
12.5 15
Weld
Number of cycles
Pre-cracked crack length (mm)
Notch angle (pre� cracked (2θ ))
Tip “A”
Tip “B”
5,05,000 5,12,880
2.45 1.75
2.80 2.20
96.89 124.83
1.5–2.5
9,70,682
3.05
2.50
159.08
1.25 1.5
2.5–5.0 5.0–7.5
9,66,652 10,86,367
2.65 3.00
2.80 2.85
68.82 77.06
15
1.5
5.0–10.0
17,03,671
3.65
2.30
108.09
Weld
10
1
5.0–10.0
20,23,897
2.90
2.55
161.52
Weld
14
1.4
10
6,44,846
3.10
2.95
126.24
Base Base
400
1000
Base Base Weld
350
900
2
S.A. Krishnan et al.
International Journal of Pressure Vessels and Piping 178 (2019) 104008
Fig. 2. (a) C(T) specimen sketch. (b) Weld configuration of C(T) specimen.
Fig. 3. Geometry and loading of pipe specimen.
Pre-cracking was done under load control, using a 50 kN capacity servo-hydraulic actuator. The fatigue pre-cracking details for pipe is provided in Table 3. The pre-cracking was carried out under constant amplitude sinusoidal cyclic loading. The maximum load during fatigue pre-cracking was approximately 20% of the analytical limit load of the pipe specimen. The limit load was calculated using following equation.
PL ¼
� 16tσf R2m θ cos 2 Z L
1 sin θ 2
� (1)
where Rm is the mean radius of the pipe, t is the thickness of the pipe, σ y is the yield stress, σu is the ultimate tensile stress, σf ( ¼ (σ y þσu)/2) is the flow stress of the material, Z and L are outer and inner spans 3
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International Journal of Pressure Vessels and Piping 178 (2019) 104008
respectively, and θ is the half crack angle. The specimens were fatigue pre-cracked till the crack growth in circumferential direction reached nearly 2 mm at both the notch tips. The number of load cycles required for introducing 2 mm crack extension at each notch tip for selected R ratio and load range is given in Table 3. 3.2. Fracture test for pipes Fig. 4(a) and (b) shows the fracture test on a straight pipe. It consists of a servo hydraulic loading system and various instruments for the measurement of data while testing. A 500 kN capacity universal testing machine (UTM) was used for loading the specimens. The specimen has been supported using pedestals with a roller support on one side and a hinge support on the other side; the loading was applied at two points through rollers. This arrangement ensures that the specimen is subjected to pure four point bending. The specimens are subjected to monotonic loading after fatigue pre-cracking. The test is carried out under displacement control. The UTM consists of an in-built linear variable displacement transducer (LVDT) for measuring load-line displacement and a load cell for measuring the applied load. The load, load-line displacement (LLD) and crack mouth opening displacement (CMOD) data are acquired. Typical load vs. displacement and load vs. crack growth data for base and weld specimens with nominal circumferential crack angle (2θ) are shown in Figs. 5 and 6 [12, 13]. Cameras are used to capture the crack growth at both tips simul taneously (Fig. 7). From the recorded images and the grids marked along the crack growth direction, crack extension has been estimated. The periodic unloading is carried out to estimate the compliance at various crack depths. This is used to calculate the plastic area required for
Fig. 4. (a) Fatigue pre-cracking on a straight pipe. (b) shows a close-up view of instrumentation during fracture test.
Fig. 5. (a–d) Load vs. displacement record of base and weld pipes with nominal crack angle, 2θ� ¼ 60 & 120. 4
S.A. Krishnan et al.
International Journal of Pressure Vessels and Piping 178 (2019) 104008
Fig. 6. (a–d) Load vs. crack extension record of base and weld pipes with nominal crack angle, 2θ� ¼ 60 & 120.
Fig. 7. Photograph of crack growth of crack tip-A of fracture experiment no. (a) SP4-60TWC-M4, (b) SP4-120TWC-M2, (c) SP4-60TWCW-M5 and (d) SP4120TWCW-M8. 5
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International Journal of Pressure Vessels and Piping 178 (2019) 104008
Fig. 8. (a) J-R curves for base material. (b) J-R curves for weld material.
Fig. 9. (a) J-T curves for base material. (b) J-T curves for weld material.
evaluating the J-R curve. Table 4 Values of fitting constants.
3.3. C(T) specimen testing The CT specimens are fatigue pre-cracked using resonant type precracker machine and side grooved (10% net depth/2 mm on either side) along the crack propagation path. The tests are conducted con forming to ASTM E1820. Load-line displacement (LLD) was monitored using high resolution crack opening displacement (COD) gage. Online crack extension (Δa) measurement was carried out using a direct current potential drop (DCPD) device. Crack extension is determined from change in voltage drop manifested due to growing crack. The tested specimen is heat-tinted and pulled open to identify the crack extension regime for the fracture test. The P, LLD and Δa record is used to construct J-Δa curve as per ASTM procedure. 4. Determination of fracture resistance curve and tearing modulus
α
β
γ
SP4-90TWC-M1 SP4-120-TWC-M2 SP4-150TWC-M3 SP4-60-TWC-M4 SP4-60TWCW-M5 SP4-90TWCW-M6 SP4-120TWCW-M8 CT specimen (Base) CT specimen (Weld)
4041.26 2708.24 2603.81 3709.68 10213.12 6672.05 3891.82 2411.09 1299.05
0.015 0.043 0.031 0.066 0.005 0.018 0.086 0.228 0.437
0.265 0.454 0.317 0.469 0.398 0.476 0.633 0.692 0.738
point bending is given by � Jel ¼ K 2 E’
(3)
pffiffiffiffiffiffiffiffiffi K ¼ σ :ðFG Þ ðπaÞ;
(4)
where
The load-displacement and load-crack extension data has been used to generate J-R curve for pipe based on method proposed by Zahoor and Kanninen [3]. The J-integral for through-wall cracked pipe under four-point bending is given by the following expressions. J ¼ Jel þ Jpl
Specimen ID
ðFG Þ2 ¼ 0:7631
1:7602x þ 1:3511x2
0:3822x3
�� ð1
xÞ3 ;
x is the ratio of cracked area to the cross-sectional area of the pipe,
(2)
a is half of the circumferential crack length, and
The elastic solution for through-wall cracked pipe subjected to four6
S.A. Krishnan et al.
International Journal of Pressure Vessels and Piping 178 (2019) 104008
Table 5 Experimental and analytical limit moment. Specimen ID
Crack Location
Expt. (Mmax) expt./ 4R2m tσf
Predicted Eqn. (16)/4R2m tσf σf ¼
Eqn. (17)/4R2m tσf σf ¼
σy þ σu
σy þ σu
2 (% diff.)
SP4-90TWC-M1 SP4-120TWCM2 SP4-150TWCM3 SP4-60TWC-M4 SP4-60TWCWM5 SP4-90TWCWM6 SP4-150TWCWM7 SP4-120TWCWM8
Jpl ¼ β
0.594 0.460
0.537 (þ9.67) 0.411 (þ10.63)
0.46 (þ23.22) 0.35 (þ24.03)
61.8 45
62.05 48.00
80.61 73.44
Base
0.310
0.277 (þ10.71)
0.24 (þ24.10)
52
32.38
64.79
Base Weld
0.692 0.574
0.672 (þ2.93) 0.633 ( 10.29)
0.57 (þ17.49) 0.54 (þ6.25)
68.4 53.7
72.25 79.26
87.72 123.70
Weld
0.423
0.486 ( 15.03)
0.41 (þ2.23)
41
58.35
112.30
Weld
0.253
0.27 ( 6.61)
0.23 (þ9.38)
26
34.98
92.79
Weld
0.408
0.407 (þ0.29)
0.35 (þ15.24)
43
56.37
105.64
θ
Pdδ þ 2 0
of Δa ¼ 3.927 mm and 4.547 mm for base and weld specimens respec tively which are small compared to crack growth in the case of base and weld pipes. The C(T) specimen J-R curves have been extrapolated for effective comparison with the component J-R curves. Beyond maximum J-resistance value, the J-R curve has been extrapolated in a tangentially linear manner up to a J-resistance value equal to twice the measured maximum J-resistance value on J-resistance vs. tearing modulus plane [15]. The equations of the extrapolated J-T curve and the corresponding J-R curves are as follows; For base C(T) specimen, the equations for the extended J-T curve and corresponding J-R curve are:
(5)
γ:Jpl dθ; θ0
where δ varies from 0 to final δ and θ varies from θ0 to final θ, 2P ¼ total bending load, δ - load line displacement, 2θ - total crack angle β¼
(6)
h’ ð2θÞ = Rthð2θÞ;
� ;
1 2 sinðθÞ
=
� hð2θÞ ¼ cosðθ = 2Þ
J ¼ 2343:77
(7)
γ ¼ h’’ ð2θÞ = h’ ð2θÞ;
h’ ð2θÞ ¼ dhð2θÞ = 2dθ
(8)
J ¼ 2343:77
J ¼ 1294:87
J ¼ 1294:87
(10)
E dJ
σ 2f da
(14)
3:2T
129:78*e
�
4:55 Δa 3:81
(15)
5.2. Limit load analysis
where α, β, γ are the fitting constants and Δa is the crack extension in mm. The tearing modulus is evaluated as per the procedure proposed by Paris et al. [14]. T¼
(13)
The fitting constants for pipe and CT specimens are given in Table 4. Figs. 8 (a–b) and 9 (a–b) shows the J-R and J-T curves for base and weld specimens with through-wall circumferential crack of various sizes. Due to insufficient synchronized experimental data for weld pipe of 150� crack angle, it has not been analyzed. With increase in initial crack size, the pipe J-R curve is found to shift downwards. Large crack size increases the crack tip constraint, leading to decrease in crack resistance. Though there is a uniform initial crack angle interval of ~30� , however, the J-R curve of 60� & 90� are close to each other and similarly 120� & 150� are close to each other in case of base pipes.
The J-R curves obtained from CT specimen and pipe tests are shown in Fig. 8. The J-R curves for both the pipes and CT specimens are fitted to exponential function. The equations are as follows; Þ
601:75*e
�
5.1. The J-R curves
βΔa γ
�
3:93 Δa 3:65
For weld-CT specimen, the equations are:
(9)
5. Results and discussion
e
(12)
3:06T �
The second term in Eq. (5) represents the correction due to crack growth. This term vanishes when the plastic J is computed at the point of crack initiation. For stable crack growth, the first term in Eq. (5) is calculated using area under the load versus LLD record to obtain an approximate value of Jpl .
J ¼ αð1
Analy. limit load (kN)
Base Base
Z
δ
Expt. Max. load (kN)
2:4 (% diff.)
σ is the tensile bending stress in the outer fibre Z
Expt. Crack initiation load (kN)
The experimental maximum moment of pipe is compared with analytical estimate. This will help to identify dominant failure mode i.e. plastic collapse or ductile tearing. The plastic collapse moment (ML) is calculated as follows:
(11)
ML ¼ 4R2m tσ f ½cosðθ = 2Þ
In the above equation, E is the Young’s modulus in N/mm2, σf is the flow stress in N/mm2, J is the J-resistance in kJ/mm2, da is the change in crack extension d (Δa) in mm. The J-R curves for C(T) specimens are obtained for crack extensions
0:5 sinðθÞ�
(16)
where, Rm is the mean radius of the pipe cross-section, t, wall thickness,
σf is flow stress (σf ¼ (σy þ σu)/2) and θ, the semi-crack angle.
Moulin and Delliou [5] modified equation (16) to account for the crack propagation at maximum moment and proposed the following
7
S.A. Krishnan et al.
International Journal of Pressure Vessels and Piping 178 (2019) 104008
equation. Mc ¼ 0:854 �
References R2m t f ½cosðθ = 2Þ
σ
0:5 sinðθÞ�
(17)
[1] V. Sumathi, S. Jalaldeen, P. Selvaraj, S. Murugan, Implications of large scale sodium water reactions in an LMFBR, Nucl. Eng. Des. 337 (2018) 364–377. [2] ASTM International, E1820-18a Standard Test Method for Measurement of Fracture Toughness, ASTM International, West Conshohocken, PA, 2018. [3] A. Zahoor, M.F. Kanninen, A plastic fracture mechanics prediction of fracture instability in a circumferentially cracked pipe in cracked pipe in bending -Part-I: JIntegral analysis, Press, Vessel Technol. Trans. ASME 103 (1981) 352–358. [4] K. Forster, L. Gruter, W. Setz, S. Bhandari, J.P. Debaene, C. Faidy, K.H. Schwalble, Crack resistance of austenitic pipes with circumferential through-wall cracks, Int. J. Press. Vessel. Pip. 65 (1996) 335–342. [5] D. Moulin, P. Delliou, French experimental studies of circumferentially throughwall cracked austenitic pipes under static bending, Int. J. Press. Vessel. Pip. 65 (1996) 343–352. [6] J. Chattopadhyay, B.K. Dutta, H.S. Kushwaha, Experimental and analytical study of three-point bend specimen and through-wall circumferentially straight pipe, Int. J. Press. Vessel. Pip. 77 (2000) 455–471. [7] P.K. Singh, K.K. Vaze, A.K. Ghosh, H.S. Kushwaha, D.M. Pukazhendi, D.S. R. Murthy, Crack resistance of austenitic stainless steel pipe and pipe welds with a circumferential crack under monotonic loading, Fatigue Fract Engg Mater Struct 29 (2005) 901–915. [8] S. Tarafder, S. Shivaprasad, M. Tarafder, P. Prasad, V.R. Raghunath, D. Swapan, Specimen Size and Constraint Effect on J-R Curves of SA333Gr-6 Steel, National Metallurgical Laboratory, Jamshedpur, India, 2000. Technical report. [9] T.V. Pavankumar, J. Chattopadhyay, B.K. Dutta, H.S. Kushwaha, Transferability of specimen J-R curve to straight pipes with throughwall circumferential flaws, Int. J. Press. Vessel. Pip. 79 (2002) 127–134. [10] J. Chattopadhyay, H.S. Kushwaha, E. Roos, Some recent developments on integrity assessment of pipes and elbows. Part II: experimental investigations, Int. J. Solids Struct. 43 (2006) 2932–2958. [11] S. Vishnuvardhan, D.M. Pukazhendhi, M. Saravanan, P. Gandhi, G. Raghava, Experimental Tests on 219mm OD Carbon Steel Pipes Having Circumferential Through-Wall Notch with Internal Pressure, 2009. Fracture test report number3. [12] S. Vishnuvardhan, M. Saravanan, P. Gandhi, G. Raghava, Fracture Studies on AISI Type 316L(N) Stainless Steel Straight Pipes Having Notch in the Base Metal, October 2016. Report No. R&D 03-SSP 17741-SR-01. [13] S. Vishnuvardhan, M. Saravanan, P. Gandhi, G. Raghava, Fracture Studies on AISI Type 316L(N) Stainless Steel Straight Pipes Having Notch in the Weld, March 2017. Report No. R&D 03-SSP 17741-SR-02 & Final. [14] P.C. Paris, H. Tada, A. Zahoor, H. Ernst, The theory of instability of the tearing mode of elastic-plastic crack growth, in: J.D. Landes, J.A. Begley, G.A. Clarke (Eds.), Elastic-plastic Fracture, ASTM STP 668 Philadelphia: American Society for Testing and Materials, 1979, pp. 5–36. [15] Report of the United States Nuclear Regulatory Commission Piping Review Committee, Evaluation of Potential for Pipe Breaks vol. 3, NUREG/CR-1061, 1984.
The above equation is obtained by modifying the flow stress term to
σf ¼ (σy þ σu)/2.4 instead of σf ¼ (σy þ σu)/2 in eq. (16) and is found to be valid for pipes with crack angle greater than 30� ; below which the pipes undergo ovalisation and fail by buckling. eqns. (16) and (17) are adopted to estimate the critical limit moment for 88.9 mm OD pipes. The estimated values and percentage difference with respect to experiment [(experiment-estimated) X 100/experiment values] are shown in Table 5. eqn. (16) is able to provide close prediction for base pipe and reasonable prediction for welds except 900. Eqn. (17) is able to provide close prediction for weld pipes except for 1200. The positive difference indicates the prediction is conservative. 6. Conclusions 1. The fracture resistance of the pipes found to be decreasing with the increase in initial crack size (angle). The effect of initial crack angle on J-R curve in case of weld pipe is negligible, especially in the initial regime of crack growth (up to 15 mm). 2. In case of weld pipes, the crack has deviated into interface/base material region which could be due to extensive plastic deformation in the interface region. 3. Crack initiation occurs much earlier than maximum bending moment. 4. Existing limit load expressions are able to provide reasonable pre diction for base and weld pipes. The expression based on multipli cation factor of 0.85 for flow stress is able to provide better prediction for weld pipe. Appendix A. Supplementary data Supplementary data to this article can be found online at https://doi. org/10.1016/j.ijpvp.2019.104008.
8
Contents lists available at ScienceDirect
International Journal of Pressure Vessels and Piping journal homepage: http://www.elsevier.com/locate/ijpvp
Evaluation of fracture resistance of AISI type 316LN stainless steel base and welded pipes with circumferential through-wall crack S.A. Krishnan a, *, R. Nikhil a, G. Sasikala a, A. Moitra a, Shaju K. Albert a, A.K. Bhaduri a, C. Lakshmana Rao b, S. Vishnuvardhan c, M. Saravanan c, P. Gandhi c, G. Raghava c a b c
Indira Gandhi Centre for Atomic Research, HBNI, Kalpakkam, Tamil Nadu, India Department of Applied Mechanics, Indian Institute of Technology Madras, Chennai, India Structural Engineering Research Centre - CSIR, Chennai, India
A R T I C L E I N F O
A B S T R A C T
Keywords: Fracture experiment Pipe Pipe weld C(T) specimen J-R curve Limit load
Fracture resistance of 316LN SS base and welded straight pipes with circumferential through-wall cracks has been evaluated. The pipe J-R curves have been determined using existing experimental methods and analytical expressions available in literature. The J-R curves of 88.9 mm OD pipe and pipe welds have been compared with the J-R curves obtained from C(T) specimen tests. The results indicate that the C(T) specimen data is not suitable for fracture assessment of pipes due to difference in constraint level and the fracture resistance of the base pipe is inferior to welded pipe. Further, the limit load analyses based on expressions available in literature reveals that the initiation and crack propagation are observed prior to the maximum bending moment.
1. Introduction AISI Type 316LN stainless steel (316LN SS) is used as the major piping material for liquid sodium cooled fast breeder reactor programme in India. Due to exothermic affinity of sodium for water, sodium-water/ steam reaction may occur in the event of leakage from high pressure steam generator tubes. This reaction leads to high-pressure waves and corrosive products, which may affect the structural integrity of inter connecting piping/components of secondary sodium main circuit, Sumathi et al. [1]. Further, any leakage due to undetected crack in piping may react with air and produce corrosion and remain undetect able while progressing and finally leads to rupture when an incidental loading gets imposed on the normal load. For any initial defect in the sodium circuit, either undetected or undetectable by the available methods of inspection, one must be able to compute the growth of defect/crack during plant operation, taking into account “design basis event” loads which includes normal, upset, emergency and faulted conditions. In spite of having well established procedures [2] for assessment of fracture behavior from laboratory scale specimens like C(T) and three point bend specimens; the applicability of these data for integrity assessment of piping components still remains a challenge due to limited crack growth data. This has been addressed by many researchers for the
piping components [3–7]. Experimental methods to evaluate the J-R curve for straight pipes have been established by Zahoor and Kanninen €rster [3]. For 316L pipes, the fracture resistance has been evaluated by Fo et al. [4]. The differences in J-R curves obtained from pipes and C(T) specimens have been reported. Similar observations were also made by Moulin and Delliou [5] and Chattopadhyay et al. [6] for circum ferentially through-wall cracked 316L and SA 333 Gr.6 carbon steel pipes respectively. This observation also holds good for 304LN SS as reported by P. K.Singh et al. [7]. Overall, this is related to the differences in crack tip constraints for the C(T) and piping components. The influ ence of crack tip constraint on material fracture toughness at specimen level has been discussed in Tatafder et al. [8] and the comparison of fracture toughness from laboratory specimens with the fracture resis tance of components for SA 333 Grade 6 carbon steel has been discussed in Pavankumar et al. [9]. To the author’s knowledge, for the 316LN SS, which can be classified as a low yield stress but high hardening material, the fracture results for the pipes and pipe welds, is not available in open literature. Further, the transferability problem could be more prominent due to following reasons: (1) most of the piping components used in sodium cooled fast breeder reactor are thinner than the minimum thickness required to evaluate the fracture resistance from standard specimens, (2) limited crack extension data (1.5–3 mm) is available from fracture tests on laboratory specimens, whereas the crack growth
* Corresponding author. E-mail address: [email protected] (S.A. Krishnan). https://doi.org/10.1016/j.ijpvp.2019.104008 Received 20 March 2019; Received in revised form 17 October 2019; Accepted 27 October 2019 Available online 31 October 2019 0308-0161/© 2019 Elsevier Ltd. All rights reserved.
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International Journal of Pressure Vessels and Piping 178 (2019) 104008
Table 1 Chemical composition in wt (%). Base
C
Si
Mn
Cr
Ni
N
Mo
P
S
Fe
0.027
0.22
1.7
17.5
12.22
0.07
2.49
0.01
0.01
Bal.
seamless and welded specimens. The chemical composition and me chanical properties of 316LN SS are presented in Table 1 and Table 2. The details of specimens are shown in Table 3.
Table 2 Mechanical properties at room temperature. Material
Yield strength (YS), MPa
Ultimate tensile strength (UTS), MPa
Modulus of elasticity (E), GPa
Poisson’s ratio (μ)
316LN SS Base
250
570
200
0.3
316LN SS Weld
420
665
200
2.2. C(T) specimen A 30 mm thick weld pad with butt weld groove of total 20� bevel angle and 14 mm root gap was prepared using multi-pass SMAW pro cess. X-ray radiography has been carried out to ensure the welds are free from detectable defects like cracks, inclusions etc. The C(T) specimens are fabricated as per sketch shown in Fig. 2 (a & b). In case of weld specimen, the crack plane is kept parallel to the weldment. The notch was introduced by electric discharge machining process at the weld centre line (see Figure 2).
0.3
analysis requires large extensions (40–50 mm) for components and (3) scatter in fracture test data of welded specimens. Hence, to ensure the integrity of 316LN SS piping and fracture transferability, a detailed fracture analysis of various components, e.g. straight pipes, elbows and branch tees have been proposed in parallel with the conventional lab oratory specimen tests. The present work is a part of experimental and numerical studies being conducted at IGCAR Kalpakkam to establish the transferable fracture parameters from specimens to components. Experiments have been carried out to investigate the fracture behavior of through-wall cracked 316LN SS base and welded straight pipes. The J-R curves evaluated from pipes have been compared with that from laboratory C (T) specimens. Further, the comparison between experimental maximum load and analytical limit load has been presented.
3. Experiments 3.1. Fatigue pre-cracking Fracture tests are conducted on the fatigue pre-cracked base and weld pipe specimens at room temperature. Similar type of tests on pipes and its analysis have been reported by several researchers [3–7,10,11]. The fatigue pre-cracking of the specimens was carried out under four-point bending to produce sharp crack front as shown in Fig. 3.
2. Material properties and specimen details 2.1. Pipe specimen The seamless pipe specimens were fabricated from 316LN SS pipe of 80 NB, Sch. 40 category (OD 88.9 mm � 5.49 mm thick). The circum ferential through-wall notch was introduced at mid span of length. The girth welded pipe specimens are prepared using 16-8-2 filler wire for root pass and modified E316-15 electrode for subsequent passes with weld configuartion as shown in Figure 1. The root passes are made by gas tungsten arc welding (GTAW) procedure and subsequent passes by shielded metal arc welding (SMAW) procedure. X-ray radiography of the welds has been carried out to ensure absence of weld defects. Circumferential U-notch of root radius 1.5 mm has been introduced in
Fig. 1. Details of pipe weld configuration.
Table 3 Experimental details of pipe specimens. Specimen ID
SP4-90TWC-M1 SP4-120TWCM2 SP4-150TWCM3 SP4-60TWC-M4 SP4-60TWCWM5 SP4-90TWCWM6 SP4-150TWCWM7 SP4-120TWCWM8
Crack location
Inner Span (mm)
Outer Span (mm)
Fatigue pre-cracking load (kN)
Frequency (Hz)
Max. (Pmax)
Min. (Pmin)
10 7.5
1 0.75
1.0–2.0 1.0–2.0
5
0.5
12.5 15
Weld
Number of cycles
Pre-cracked crack length (mm)
Notch angle (pre� cracked (2θ ))
Tip “A”
Tip “B”
5,05,000 5,12,880
2.45 1.75
2.80 2.20
96.89 124.83
1.5–2.5
9,70,682
3.05
2.50
159.08
1.25 1.5
2.5–5.0 5.0–7.5
9,66,652 10,86,367
2.65 3.00
2.80 2.85
68.82 77.06
15
1.5
5.0–10.0
17,03,671
3.65
2.30
108.09
Weld
10
1
5.0–10.0
20,23,897
2.90
2.55
161.52
Weld
14
1.4
10
6,44,846
3.10
2.95
126.24
Base Base
400
1000
Base Base Weld
350
900
2
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International Journal of Pressure Vessels and Piping 178 (2019) 104008
Fig. 2. (a) C(T) specimen sketch. (b) Weld configuration of C(T) specimen.
Fig. 3. Geometry and loading of pipe specimen.
Pre-cracking was done under load control, using a 50 kN capacity servo-hydraulic actuator. The fatigue pre-cracking details for pipe is provided in Table 3. The pre-cracking was carried out under constant amplitude sinusoidal cyclic loading. The maximum load during fatigue pre-cracking was approximately 20% of the analytical limit load of the pipe specimen. The limit load was calculated using following equation.
PL ¼
� 16tσf R2m θ cos 2 Z L
1 sin θ 2
� (1)
where Rm is the mean radius of the pipe, t is the thickness of the pipe, σ y is the yield stress, σu is the ultimate tensile stress, σf ( ¼ (σ y þσu)/2) is the flow stress of the material, Z and L are outer and inner spans 3
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International Journal of Pressure Vessels and Piping 178 (2019) 104008
respectively, and θ is the half crack angle. The specimens were fatigue pre-cracked till the crack growth in circumferential direction reached nearly 2 mm at both the notch tips. The number of load cycles required for introducing 2 mm crack extension at each notch tip for selected R ratio and load range is given in Table 3. 3.2. Fracture test for pipes Fig. 4(a) and (b) shows the fracture test on a straight pipe. It consists of a servo hydraulic loading system and various instruments for the measurement of data while testing. A 500 kN capacity universal testing machine (UTM) was used for loading the specimens. The specimen has been supported using pedestals with a roller support on one side and a hinge support on the other side; the loading was applied at two points through rollers. This arrangement ensures that the specimen is subjected to pure four point bending. The specimens are subjected to monotonic loading after fatigue pre-cracking. The test is carried out under displacement control. The UTM consists of an in-built linear variable displacement transducer (LVDT) for measuring load-line displacement and a load cell for measuring the applied load. The load, load-line displacement (LLD) and crack mouth opening displacement (CMOD) data are acquired. Typical load vs. displacement and load vs. crack growth data for base and weld specimens with nominal circumferential crack angle (2θ) are shown in Figs. 5 and 6 [12, 13]. Cameras are used to capture the crack growth at both tips simul taneously (Fig. 7). From the recorded images and the grids marked along the crack growth direction, crack extension has been estimated. The periodic unloading is carried out to estimate the compliance at various crack depths. This is used to calculate the plastic area required for
Fig. 4. (a) Fatigue pre-cracking on a straight pipe. (b) shows a close-up view of instrumentation during fracture test.
Fig. 5. (a–d) Load vs. displacement record of base and weld pipes with nominal crack angle, 2θ� ¼ 60 & 120. 4
S.A. Krishnan et al.
International Journal of Pressure Vessels and Piping 178 (2019) 104008
Fig. 6. (a–d) Load vs. crack extension record of base and weld pipes with nominal crack angle, 2θ� ¼ 60 & 120.
Fig. 7. Photograph of crack growth of crack tip-A of fracture experiment no. (a) SP4-60TWC-M4, (b) SP4-120TWC-M2, (c) SP4-60TWCW-M5 and (d) SP4120TWCW-M8. 5
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International Journal of Pressure Vessels and Piping 178 (2019) 104008
Fig. 8. (a) J-R curves for base material. (b) J-R curves for weld material.
Fig. 9. (a) J-T curves for base material. (b) J-T curves for weld material.
evaluating the J-R curve. Table 4 Values of fitting constants.
3.3. C(T) specimen testing The CT specimens are fatigue pre-cracked using resonant type precracker machine and side grooved (10% net depth/2 mm on either side) along the crack propagation path. The tests are conducted con forming to ASTM E1820. Load-line displacement (LLD) was monitored using high resolution crack opening displacement (COD) gage. Online crack extension (Δa) measurement was carried out using a direct current potential drop (DCPD) device. Crack extension is determined from change in voltage drop manifested due to growing crack. The tested specimen is heat-tinted and pulled open to identify the crack extension regime for the fracture test. The P, LLD and Δa record is used to construct J-Δa curve as per ASTM procedure. 4. Determination of fracture resistance curve and tearing modulus
α
β
γ
SP4-90TWC-M1 SP4-120-TWC-M2 SP4-150TWC-M3 SP4-60-TWC-M4 SP4-60TWCW-M5 SP4-90TWCW-M6 SP4-120TWCW-M8 CT specimen (Base) CT specimen (Weld)
4041.26 2708.24 2603.81 3709.68 10213.12 6672.05 3891.82 2411.09 1299.05
0.015 0.043 0.031 0.066 0.005 0.018 0.086 0.228 0.437
0.265 0.454 0.317 0.469 0.398 0.476 0.633 0.692 0.738
point bending is given by � Jel ¼ K 2 E’
(3)
pffiffiffiffiffiffiffiffiffi K ¼ σ :ðFG Þ ðπaÞ;
(4)
where
The load-displacement and load-crack extension data has been used to generate J-R curve for pipe based on method proposed by Zahoor and Kanninen [3]. The J-integral for through-wall cracked pipe under four-point bending is given by the following expressions. J ¼ Jel þ Jpl
Specimen ID
ðFG Þ2 ¼ 0:7631
1:7602x þ 1:3511x2
0:3822x3
�� ð1
xÞ3 ;
x is the ratio of cracked area to the cross-sectional area of the pipe,
(2)
a is half of the circumferential crack length, and
The elastic solution for through-wall cracked pipe subjected to four6
S.A. Krishnan et al.
International Journal of Pressure Vessels and Piping 178 (2019) 104008
Table 5 Experimental and analytical limit moment. Specimen ID
Crack Location
Expt. (Mmax) expt./ 4R2m tσf
Predicted Eqn. (16)/4R2m tσf σf ¼
Eqn. (17)/4R2m tσf σf ¼
σy þ σu
σy þ σu
2 (% diff.)
SP4-90TWC-M1 SP4-120TWCM2 SP4-150TWCM3 SP4-60TWC-M4 SP4-60TWCWM5 SP4-90TWCWM6 SP4-150TWCWM7 SP4-120TWCWM8
Jpl ¼ β
0.594 0.460
0.537 (þ9.67) 0.411 (þ10.63)
0.46 (þ23.22) 0.35 (þ24.03)
61.8 45
62.05 48.00
80.61 73.44
Base
0.310
0.277 (þ10.71)
0.24 (þ24.10)
52
32.38
64.79
Base Weld
0.692 0.574
0.672 (þ2.93) 0.633 ( 10.29)
0.57 (þ17.49) 0.54 (þ6.25)
68.4 53.7
72.25 79.26
87.72 123.70
Weld
0.423
0.486 ( 15.03)
0.41 (þ2.23)
41
58.35
112.30
Weld
0.253
0.27 ( 6.61)
0.23 (þ9.38)
26
34.98
92.79
Weld
0.408
0.407 (þ0.29)
0.35 (þ15.24)
43
56.37
105.64
θ
Pdδ þ 2 0
of Δa ¼ 3.927 mm and 4.547 mm for base and weld specimens respec tively which are small compared to crack growth in the case of base and weld pipes. The C(T) specimen J-R curves have been extrapolated for effective comparison with the component J-R curves. Beyond maximum J-resistance value, the J-R curve has been extrapolated in a tangentially linear manner up to a J-resistance value equal to twice the measured maximum J-resistance value on J-resistance vs. tearing modulus plane [15]. The equations of the extrapolated J-T curve and the corresponding J-R curves are as follows; For base C(T) specimen, the equations for the extended J-T curve and corresponding J-R curve are:
(5)
γ:Jpl dθ; θ0
where δ varies from 0 to final δ and θ varies from θ0 to final θ, 2P ¼ total bending load, δ - load line displacement, 2θ - total crack angle β¼
(6)
h’ ð2θÞ = Rthð2θÞ;
� ;
1 2 sinðθÞ
=
� hð2θÞ ¼ cosðθ = 2Þ
J ¼ 2343:77
(7)
γ ¼ h’’ ð2θÞ = h’ ð2θÞ;
h’ ð2θÞ ¼ dhð2θÞ = 2dθ
(8)
J ¼ 2343:77
J ¼ 1294:87
J ¼ 1294:87
(10)
E dJ
σ 2f da
(14)
3:2T
129:78*e
�
4:55 Δa 3:81
(15)
5.2. Limit load analysis
where α, β, γ are the fitting constants and Δa is the crack extension in mm. The tearing modulus is evaluated as per the procedure proposed by Paris et al. [14]. T¼
(13)
The fitting constants for pipe and CT specimens are given in Table 4. Figs. 8 (a–b) and 9 (a–b) shows the J-R and J-T curves for base and weld specimens with through-wall circumferential crack of various sizes. Due to insufficient synchronized experimental data for weld pipe of 150� crack angle, it has not been analyzed. With increase in initial crack size, the pipe J-R curve is found to shift downwards. Large crack size increases the crack tip constraint, leading to decrease in crack resistance. Though there is a uniform initial crack angle interval of ~30� , however, the J-R curve of 60� & 90� are close to each other and similarly 120� & 150� are close to each other in case of base pipes.
The J-R curves obtained from CT specimen and pipe tests are shown in Fig. 8. The J-R curves for both the pipes and CT specimens are fitted to exponential function. The equations are as follows; Þ
601:75*e
�
5.1. The J-R curves
βΔa γ
�
3:93 Δa 3:65
For weld-CT specimen, the equations are:
(9)
5. Results and discussion
e
(12)
3:06T �
The second term in Eq. (5) represents the correction due to crack growth. This term vanishes when the plastic J is computed at the point of crack initiation. For stable crack growth, the first term in Eq. (5) is calculated using area under the load versus LLD record to obtain an approximate value of Jpl .
J ¼ αð1
Analy. limit load (kN)
Base Base
Z
δ
Expt. Max. load (kN)
2:4 (% diff.)
σ is the tensile bending stress in the outer fibre Z
Expt. Crack initiation load (kN)
The experimental maximum moment of pipe is compared with analytical estimate. This will help to identify dominant failure mode i.e. plastic collapse or ductile tearing. The plastic collapse moment (ML) is calculated as follows:
(11)
ML ¼ 4R2m tσ f ½cosðθ = 2Þ
In the above equation, E is the Young’s modulus in N/mm2, σf is the flow stress in N/mm2, J is the J-resistance in kJ/mm2, da is the change in crack extension d (Δa) in mm. The J-R curves for C(T) specimens are obtained for crack extensions
0:5 sinðθÞ�
(16)
where, Rm is the mean radius of the pipe cross-section, t, wall thickness,
σf is flow stress (σf ¼ (σy þ σu)/2) and θ, the semi-crack angle.
Moulin and Delliou [5] modified equation (16) to account for the crack propagation at maximum moment and proposed the following
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International Journal of Pressure Vessels and Piping 178 (2019) 104008
equation. Mc ¼ 0:854 �
References R2m t f ½cosðθ = 2Þ
σ
0:5 sinðθÞ�
(17)
[1] V. Sumathi, S. Jalaldeen, P. Selvaraj, S. Murugan, Implications of large scale sodium water reactions in an LMFBR, Nucl. Eng. Des. 337 (2018) 364–377. [2] ASTM International, E1820-18a Standard Test Method for Measurement of Fracture Toughness, ASTM International, West Conshohocken, PA, 2018. [3] A. Zahoor, M.F. Kanninen, A plastic fracture mechanics prediction of fracture instability in a circumferentially cracked pipe in cracked pipe in bending -Part-I: JIntegral analysis, Press, Vessel Technol. Trans. ASME 103 (1981) 352–358. [4] K. Forster, L. Gruter, W. Setz, S. Bhandari, J.P. Debaene, C. Faidy, K.H. Schwalble, Crack resistance of austenitic pipes with circumferential through-wall cracks, Int. J. Press. Vessel. Pip. 65 (1996) 335–342. [5] D. Moulin, P. Delliou, French experimental studies of circumferentially throughwall cracked austenitic pipes under static bending, Int. J. Press. Vessel. Pip. 65 (1996) 343–352. [6] J. Chattopadhyay, B.K. Dutta, H.S. Kushwaha, Experimental and analytical study of three-point bend specimen and through-wall circumferentially straight pipe, Int. J. Press. Vessel. Pip. 77 (2000) 455–471. [7] P.K. Singh, K.K. Vaze, A.K. Ghosh, H.S. Kushwaha, D.M. Pukazhendi, D.S. R. Murthy, Crack resistance of austenitic stainless steel pipe and pipe welds with a circumferential crack under monotonic loading, Fatigue Fract Engg Mater Struct 29 (2005) 901–915. [8] S. Tarafder, S. Shivaprasad, M. Tarafder, P. Prasad, V.R. Raghunath, D. Swapan, Specimen Size and Constraint Effect on J-R Curves of SA333Gr-6 Steel, National Metallurgical Laboratory, Jamshedpur, India, 2000. Technical report. [9] T.V. Pavankumar, J. Chattopadhyay, B.K. Dutta, H.S. Kushwaha, Transferability of specimen J-R curve to straight pipes with throughwall circumferential flaws, Int. J. Press. Vessel. Pip. 79 (2002) 127–134. [10] J. Chattopadhyay, H.S. Kushwaha, E. Roos, Some recent developments on integrity assessment of pipes and elbows. Part II: experimental investigations, Int. J. Solids Struct. 43 (2006) 2932–2958. [11] S. Vishnuvardhan, D.M. Pukazhendhi, M. Saravanan, P. Gandhi, G. Raghava, Experimental Tests on 219mm OD Carbon Steel Pipes Having Circumferential Through-Wall Notch with Internal Pressure, 2009. Fracture test report number3. [12] S. Vishnuvardhan, M. Saravanan, P. Gandhi, G. Raghava, Fracture Studies on AISI Type 316L(N) Stainless Steel Straight Pipes Having Notch in the Base Metal, October 2016. Report No. R&D 03-SSP 17741-SR-01. [13] S. Vishnuvardhan, M. Saravanan, P. Gandhi, G. Raghava, Fracture Studies on AISI Type 316L(N) Stainless Steel Straight Pipes Having Notch in the Weld, March 2017. Report No. R&D 03-SSP 17741-SR-02 & Final. [14] P.C. Paris, H. Tada, A. Zahoor, H. Ernst, The theory of instability of the tearing mode of elastic-plastic crack growth, in: J.D. Landes, J.A. Begley, G.A. Clarke (Eds.), Elastic-plastic Fracture, ASTM STP 668 Philadelphia: American Society for Testing and Materials, 1979, pp. 5–36. [15] Report of the United States Nuclear Regulatory Commission Piping Review Committee, Evaluation of Potential for Pipe Breaks vol. 3, NUREG/CR-1061, 1984.
The above equation is obtained by modifying the flow stress term to
σf ¼ (σy þ σu)/2.4 instead of σf ¼ (σy þ σu)/2 in eq. (16) and is found to be valid for pipes with crack angle greater than 30� ; below which the pipes undergo ovalisation and fail by buckling. eqns. (16) and (17) are adopted to estimate the critical limit moment for 88.9 mm OD pipes. The estimated values and percentage difference with respect to experiment [(experiment-estimated) X 100/experiment values] are shown in Table 5. eqn. (16) is able to provide close prediction for base pipe and reasonable prediction for welds except 900. Eqn. (17) is able to provide close prediction for weld pipes except for 1200. The positive difference indicates the prediction is conservative. 6. Conclusions 1. The fracture resistance of the pipes found to be decreasing with the increase in initial crack size (angle). The effect of initial crack angle on J-R curve in case of weld pipe is negligible, especially in the initial regime of crack growth (up to 15 mm). 2. In case of weld pipes, the crack has deviated into interface/base material region which could be due to extensive plastic deformation in the interface region. 3. Crack initiation occurs much earlier than maximum bending moment. 4. Existing limit load expressions are able to provide reasonable pre diction for base and weld pipes. The expression based on multipli cation factor of 0.85 for flow stress is able to provide better prediction for weld pipe. Appendix A. Supplementary data Supplementary data to this article can be found online at https://doi. org/10.1016/j.ijpvp.2019.104008.
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