Engineering Fracture Mechanics 148 (2015) 110–121

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Engineering Fracture Mechanics

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A comparison between fracture toughness at different locationsof longitudinal submerged arc welded and spiral submerged arcwelded joints of API X80 pipeline steels

http://dx.doi.org/10.1016/j.engfracmech.2015.09.0030013-7944/� 2015 Elsevier Ltd. All rights reserved.

⇑ Corresponding author at: Department of Mechanical and Aerospace Engineering, Carleton University, Ottawa, Ontario K1S 5B6, Canada. Tel.: +12600x8308.

E-mail address: xin.wang@carleton.ca (X. Wang).

Xu Chen a, Hongsheng Lu a, Gang Chen a, Xin Wang a,b,⇑a School of Chemical Engineering & Technology, Tianjin University, Tianjin 300072, ChinabDepartment of Mechanical and Aerospace Engineering, Carleton University, Ottawa, Ontario K1S 5B6, Canada

a r t i c l e i n f o

Article history:Received 29 May 2015Received in revised form 30 August 2015Accepted 2 September 2015Available online 11 September 2015

Keywords:Fracture toughnessX80 pipeline steelWelded jointsSingle specimen techniquePrecipitateSingle-edge-notched bending specimen

a b s t r a c t

In this paper, extensively experimental testing of fracture toughness for two differentwelded joints of API X80 pipeline steels after service were carried out. The microstructuresand CTOD resistance curves of longitudinal submerged arc welded (LSAW) and spiralsubmerged arc welded (SSAW) joints were investigated at different locations of the basemetal (BM), weld metal (WM) and heat affected zone (HAZ). The critical fracture toughnessvalues were compared with each other. The optical microscope (OM), scanning electronmicroscope (SEM) and energy dispersive spectroscopy (EDS) were used to analyze thefracture mechanism and precipitates composition. This investigation will provide animportant reference to engineering practice and evaluation of pipeline structures.

� 2015 Elsevier Ltd. All rights reserved.

1. Introduction

In order to provide adequate safety and improve transportation efficiency under high pressure conditions, pipeline steelsused to transport crude oil or natural gas over a long distance essentially not only require high strength and toughness, butalso demand thicker thickness and larger diameter [1,2]. Economic studies have been carried out and demonstrated that longdistance oil and gas transportation mostly select high-grade pipeline steels such as X80 and X100, which are able to sustainsignificant inner pressure generated inside the pipe [3]. For X80 pipeline steels, which are used more and more widely, themain failure mode is leakage or fracture induced by cracks and manufacturing defects in the pipes [4]. Therefore, thetoughness value is the required parameter for characterizing material resistance to fracture. Due to the inherentcharacteristics of the welding process, the toughness of weld metal and heat affected zone are harder to control thanbase metal and these regions have therefore become a weak link in pipelines [5,6]. For this reason, it is necessary toinvestigate the toughness difference between different regions of welded joints. The most commonly used forms of weldedpipeline are either with longitudinal submerged arc welded (LSAW) joints or spiral submerged arc welded (SSAW) joints. TheLSAW is the pipe with a longitudinal weld, and pressurized after welding to ensure roundness, while the SSAW is that with aspiral weld. As they have their own advantages and disadvantages based on the manufacturing cost and applicability, acomparison between toughness at different regions of LSAW and SSAW welded joints is necessary [7].

613 520

Nomenclature

a crack size (length)a0 original crack size (length)ai instantaneous crack size (length)Da crack extensionB specimen thicknessBe effective thickness of specimenBM base metalBN specimen net thicknessC CMOD compliance, the ratio of CMOD increment to load incrementCi measured CMOD compliance at the ith unloadingCGHAZ coarse grain heat affected zoneE modulus of elasticityF loadFGHAZ fine grain heat affected zoneFi load at the ith unloadingFZ fusion zoneICHAZ intercritically reheated heat affected zoneLSAW longitudinal submerged arc weldedR rotation radiusRp0.2 0.2% offset yield strengthRm ultimate tensile strengthS span for SENB specimenSSAW spiral submerged arc weldedV CMODVpl plastic part of CMODW specimen widthWM weld metalZ the distance between the CMOD measurement position and the specimen surfaced crack tip opening displacement, CTODd0.2BL the non-size-sensitive fracture resistance corresponding to 0.2 mm offset blunting linet Poisson’s ratioDV CMOD incrementDP load increment

X. Chen et al. / Engineering Fracture Mechanics 148 (2015) 110–121 111

The toughness has always been a focus of research interest for high strength pipeline steels [8–11]. Since the toughness isdifferent through welded joints, as a result of welding thermal cycle, the grain coarsening will deteriorate toughness in theHAZ [12,13]. In the past, the charpy V-notch (CVN) impact test and the drop weight tear test (DWTT) were considered to bethe indicator method showing resistance against unstable ductile fracture [9,14–17]. These methods are designed forevaluating pipeline fracture behavior under impact loading. To quantify the fracture characteristics under static servicecondition, quasi-static testing method is required. Recently, because of the clear physical meaning of the fracture mechanicsmethod, it has been adopted as the main approach to characterize the fracture toughness of pipeline steels [18–20].

Fracture mechanics parameters such as the stress intensity factor K, J-integral and crack tip opening displacement(CTOD), or d have been widely used as fracture criteria to characterize the onset and continuation of crack growth[18,19]. Due to the thin thickness of high-grade pipeline, it is difficult to satisfy the plane strain size conditions for Kcriterion. But with respect to CTOD fracture criterion, the requirements on size are not that strict. The CTOD concept wasproposed in 1963 by Wells to serve as an engineering fracture parameter, and can be equivalently used as K andJ-integral in fracture toughness characterization [19]. The CTOD concept bridges the elastic and plastic fracture conditions.According to Hertzberg, the CTOD was related to the extent of plastic strain in the plastic zone, which is analogous to thetoughness measurement from the area under the stress–strain curve in an uniaxial tensile test [21]. The CTOD value wasusually associated with the onset of cleavage fracture under plan strain conditions. Such result obtained directly from areasnear the blunting crack tip is inconvenient and subjective. Later, a plastic hinge model was developed by Hollstein and Blauelto determine CTOD by assuming that two arms of the specimen rotated rigidly about a plastic hinge point in the uncrackedligament [19]. This was considered to be an effective method for calculating CTOD value. In order to apply the plastic hingemodel to both elastic and elastic–plastic conditions, the total d was separated into elastic and plastic components, just likethe J separation. The equations for CTOD calculation from GB/T 21143 [22] are adopted in this paper.

In this study, macroscopic morphology and metallography of different regions of X80 pipelines were observed. The basicmechanical properties were obtained through tensile test. A procedure and equations for measurement of CTOD-resistance(d–R) curves using a single-specimen technique for three-point bending specimens recommended in GB/T 21143 [22] were

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adopted for toughness testing. The d–R curve tests of BM, WM and HAZ were carried out in LSAW and SSAW joints,respectively. The critical fracture toughness of different regions was compared between LSAW and SSAW pipeline steels.It is important to note the welds being characterized here are the seam welds that make up the original pipe, and not thegirth welds between the pipes that would be made during installation. The fracture surfaces of tensile, fatigue pre-crackand fracture toughness tested X80 specimens were documented using a scanning electron microscope (SEM) at differentregions to identify any changes with location changes. The energy dispersive spectroscopy (EDS) analysis was performedon the precipitates of welded joints fractography to identify the chemical composition.

2. Experiment procedure

2.1. Material

API X80 pipeline steels and welded joints provided by Baoshan Iron and Steel Co. Ltd. were used in this study, which werein service for five years. In particular, they could be divided into two categories according to the welding procedure, i.e. theLSAW and SSAW pipes. They were cut from API X80 linepipe with the diameter of 1219 mm and thickness of 22 mm and18.4 mm, respectively. The properties of seam welds that make up the original pipe are characterized here.

2.2. Metallographic observation of the welded joints

In order to investigate the properties of different regions in welded joints, the macroscopic morphology of the weldedjoints was determined. So two specimens that were symmetrical based on weld centerline were cut from the pipes includingwelded joints. Then the specimen were polished and etched for about 10 s using 3% nital solution [23]. As shown in Fig. 1, thedifferent regions of welded joints can be distinguished easily, identified as base metal (BM), weld metal (WM), fusion zone(FZ), and three heat affected zones, i.e., coarse grain (CG), fine grain (FG), and intercritically reheated (IC), respectively. Thenthe metallurgical of different regions were observed in the optical microscope OLYMPUS GX51. Fig. 2 shows the opticalmicrographs in the location of BM, WM, FZ, CGHAZ, FGHAZ and ICHAZ of LSAW and SSAW joints, respectively. The BM ofthe two X80 were typical acicular ferrite steel embedded with large amount of granular bainite (GB), a bit of pearlite (P),martensite/austenite (M/A) and some amount of acicular ferrite (AF). The WM was composed of acicular ferrite (AF) andgranular bainite (GB) together with some microphases. While the density of M/A in HAZ is much higher than that of BM.The M/A constituent, which is a hard-brittle phase, will deteriorate the toughness of welded joints. The FZ was a mixtureof WM and HAZ. Since the acicular ferrite had finer grain and higher dislocation density, it generally resulted in higherstrength and toughness [24]. Due to the different size of grains, it is expected that this would lead to differences in strengthand toughness. This will also be demonstrated in later experiments.

2.3. Uniaxial tensile test

The geometry of uniaxial tensile specimens, which was machined with CNC machines, were shown in Fig. 3a, and thedrawn way from each section was shown in Fig. 3b. Tensile test was conducted by an electro-hydraulic servo fatigue testmachine using strain-controlled mode with the strain rate 0.0001/s in room temperature.

2.4. Fracture toughness test

The single-edge-notched bending (SENB) and compact tension (CT) were the common specimens for resistance curve test[25,26]. Considering all aspects, we choose the former one for our tests. Since the results of multiple specimens method were

Fig. 1. Macrograph of X80 welded joints (a) LSAW and (b) SSAW.

Fig. 2. The optical micrographs in the location of (a) BM, (b) WM, (c) FZ, (d) CGHAZ, (e) FGHAZ, (f) ICHAZ of LSAW and (g) BM, (h) WM, (i) FZ, (j) CGHAZ, (k)FGHAZ, and (l) ICHAZ of SSAW.

Fig. 3. Schematic illustrations of (a) tensile specimen and (b) the locations where the specimens were prepared. All the units are in mm.

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scattered and hence require a significant amount of test material, we adopt the single-specimen method [27]. The testtemperature is room temperature. The SENB specimen with the length of 120 mm, width of 24 mm and thickness of12 mm, each with a 10 mm deep V-notched machined with 0.12 mm molybdenum wire was adopted. The specimen sizeand drawn way were shown in Fig. 4. The crack was positioned at six different regions for comparison. They were BM,WM, FZ, CGHAZ, FGHAZ and ICHAZ, respectively. Due to the conventional specified bending orientations of the SENB spec-imen is relative to the pipeline steel bending formation orientations, the drawn direction is perpendicular to the weldingdirection in the two situations. In this way we can make sure that the crack propagation all in the range to be tested, sothe results were reliable. The resistance curve test was conducted according to the newest fracture toughness test standardGB/T21143 [22]. Note GB/T21143 is the counterpart of ASTM-E1820 [25] in Chinese standards. At first, a fatigue pre-crack

Fig. 4. Schematic illustrations of (a) three-point bending specimen and the locations where the specimens were prepared in (b) LSAW pipes and (c) SSAWpipes. All the units are in mm.

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about 1.5 mm on each specimen was needed. The load-controlled fatigue pre-cracking was performed using aservo-hydraulic test machine with a loading frequency of 10 Hz. The mean load used for fatigue pre-cracking was 4 KN,and the load amplitude was 3.8 KN. After the fatigue pre-cracking, the fracture toughness test was then performed on aservo-hydraulic test machine with a maximum capacity of 100 KN. During test of each SENB specimen, a load–CMOD curvewith 12–20 unload/reload sequences was measured. The loading procedure was displacement-control with the loading rateof 0.02 mm/s, while the unloading procedure was load-control with the load rate of 0.4 KN/s. This was to prevent excessiveunloading to exceed the provisions of the standard. The CMOD was measured with a COGNEX Insight-Micro 1050 instead ofclip gauge. Two pairs of round markers were set on the specimen surface, which was used to recognize by the vision system.The pixels between a pair of marker could be recorded and transformed into CMOD. The advantage of this method was toprevent the destruction of the specimen and can obtain two CMOD values in one time. It can also be used to measure theCMOD for sample which is inaccessible or inconvenient to have direct contact. The test system was shown in Fig. 5.

After finishing fracture toughness test, the specimens were loaded by fatigue again to distinguish the leading edge ofunload/reload test. After that, the specimens were cooled with liquid nitrogen and broken open to expose the crack. Thenine-point optical technique described in ASME E1820 was used to measure the crack size for comparison with that fromunloading compliance [25]. The Hitachi S-4800 field emission scanning electron microscope (SEM) was used to identifythe distinct fractographic and microstructural features of different specimens. And the spherical precipitates observed onfracture surface of some tested specimens were characterized by SEM and EDS.

Fig. 5. Set-up for the three-point bending test of the X80 pipeline steel.

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2.5. Equations used in the test

When a single-specimen technique was used, the crack size must be evaluated during the test. The instantaneous cracksize was conveniently measured using the crack mouth opening displacement compliance, as recommended in ASTM E1820[25]. The crack length was given as follows:

Table 1Tensile

Posit

WeldHeatBase

a The

ai=W ¼ 0:999748� 3:9504uþ 2:9821u2 � 3:21408u3 þ 51:51564u4 � 113:031u5� � ð1Þ

where

u ¼ 1BeWECiS=4

h i1=2þ 1

ð2Þ

Ci ¼ DVm=DPð Þ on an unloading/reloading sequence, Vm ¼ crack opening displacement at notched edge, Be ¼ B� ðB� BNÞ2=B.The CMOD at the surface of the specimen were estimated from the CMOD measurements, V1 and V2, by using geometric

triangulation and was calculated with the following formula:

CMOD ¼ V1 � Z1

Z2 � Z1V2 � V1ð Þ ð3Þ

The format of the equations for CTOD d evaluation in the resistance curve test method used in GB/T 21143 [22] for SE(B)specimen was adopted. d is calculated at point i at current values of crack size ai, plastic CMOD Vpi, rotation radius Ri and loadPi as follows:

d ¼ SW

� �F i

ðBBNWÞ0:5� g1

aiW

� �" #2 ð1� t2Þ2Rp0:2E

� þ Ri � ai � Zð ÞVpi

Rið4Þ

where the equations for evaluating rotation radius Ri were described by the following expression:

Ri=W ¼ ai=Wð Þ21� ai=W

g ai=Wð Þ ð5Þ

where

g ai=Wð Þ ¼ 10:5108� 18:23566 ai=Wð Þ � 92:96373 ai=Wð Þ2 þ 406:1855 ai=Wð Þ3 � 648:1842 ai=Wð Þ4

þ 482:2612 ai=Wð Þ5 � 139:8167 ai=Wð Þ6 ð6Þ

The value of g1ðai=WÞ are listed in Annex B of GB/T 21143.Z = the distance between the CMOD measurement position and the specimen surface. In test the value of Z is �2 and �6,

respectively, the negative sign stands for the inner part of specimen.Vpi = plastic component of Vi.When the d–Da resistance curve was determined, we need to determine a blunting line in accordance with the following

equation:

d ¼ 1:87ðRm=Rp0:2Þ � Da ð7Þ

After that, the equations of left limit line and 0.2 mm offset line were determined.

3. Results and discussion

3.1. Uniaxial tensile test

Tables 1 and 2 showed the typical tensile properties of the BM, WM and HAZ of the two tested welded joints, respectively.All specimens showed continuous yielding behavior. The studied welded joints were overmatched in that WM had higherstrength than the BM due to its high alloy content.

test data for LSAW.

ion Elasticity modulus (GPa) Yield strength (MPa) Ultimate strength (MPa) Yield to tensile ratio (%) Total elongationa (%)

metal 180 668 715 93.4 15affected zone 202 598 669 89.4 22metal 190 570 637 89.5 19

gauge length is 20 mm.

Table 2Tensile test data for SSAW.

Position Elasticity modulus (GPa) Yield strength (MPa) Ultimate strength (MPa) Yield to tensile ratio (%) Total elongationa (%)

Weld metal 191 672 718 93.6 27Heat affected zone 193 677 740 91.5 26Base metal 195 590 710 83.1 25

a The gauge length is 20 mm.

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3.2. Resistance curve tests

During each resistance curve test, a load–CMOD curve was measured. Fig. 6 showed P–CMOD curve for the base metal ofSSAW pipes. The specimen surface CMOD can be obtained from Fig. 6 through Eq. (3). The P–CMOD curves of differentregions are slightly different. The unloading compliance which basically continually increased with test was obtained fromthese curves. Then it was used to evaluate the crack size from Eq. (1). The crack size and the load–plastic CMOD curves werethen used to evaluate the CTOD d value from Eq. (4). The d–R resistance curves were shown in Fig. 7. Fig. 7a–c shows d–Rcurves of BM, HAZ and WM of LSAW pipes. For comparison, four resistance curves: one of FZ and three (FG, CG and IC)for HAZ, were included in Fig. 7b. The curve for FGHAZ specimen was slightly above those other three, although thedifference is small. The reason why FGHAZ had better toughness was the smaller and more uniform grain size than the otherthree because of the welding heat input [12,28]. So the reduction of toughness was smaller in comparison with those otherthree. The toughness of FZ was lowest in all sections of welded joints. The reason of this phenomenon was the unevencharacteristics of FZ in which one part was the WM, while the other part was HAZ. This results in the FZ has the lowesttoughness. As the BM didn’t affect by the welding heat input, so it has the highest resistance. The later fractographicobservations can also explain the toughness difference between different locations. The resistance curve results of SSAWpipes were shown in Fig. 7d–f. It showed similar trend with LSAW pipes.

From the intersection of 0.2 mm offset line and resistance curve, the critical fracture toughness d0.2BL was obtained. Theeffect of position on the d0.2BL value of LSAW and SSAW was illustrated in Fig. 8. The toughness of the WM and HAZ werelower than that of the BM which was discussed previously. For BM, the toughness of SSAW was lower than the LSAW. FromFig. 8 we can also find that the toughness values of LSAW pipes in WM was smaller in comparison with that of SSAW pipes,which was in contradiction with the BM. There is a small toughness difference between LSAW and SSAW in HAZ.

The initial and final crack sizes were evaluated using CMOD unloading compliance. After the test, each specimen wascooled by liquid nitrogen and was then broken open. The initial and final crack sizes were measured via a readingmicroscope. In Fig. 9, the initial and final crack sizes (solid and open points respectively) estimated using the unloadingcompliance technical were compared with those measured optically. It was found that the excellent agreement was achievedbetween these two techniques. Hence, the unloading compliance technique was reliable in the test.

3.3. Fractographic observations

3.3.1. Uniaxial tensile test fractographic observationAdditional indirect measurements of the micro-structural changes involving the density and particles were achieved

through SEM observations of fracture surfaces obtained from the tensile fracture specimens for the base metal, weld metaland HAZ. The fracture surfaces exhibited a ductile-type failure with well-defined micro-void morphology, which was linkedto the mechanism of nucleation, growth and coalescence of micro-cavities, as can be observed in Fig. 10. The micro-cavity

Fig. 6. Load–CMOD curves for BM of SSAW pipes.

Fig. 7. d–R curves in the position of (a) BM, (b) HAZ, (c) WM of LSAW pipe and (d) BM, (e) HAZ, and (f) WM of SSAW pipe.

Fig. 8. Critical CTOD values in different positions of LSAW and SSAW.

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nucleation centers were promoted by the presence of carbides and non-metallic inclusions. From Fig. 10a it was found thatthe dimples in BM were the deepest and biggest, suggesting the highest toughness throughout the welded joints. This wasalso supported by results concluded from former resistance curve test. The dimples of WM and BM were basically circular,while those were ellipse in HAZ presumably because of the uneven welding heat input. Since distances from weldingcenterline were different, as a result a heat gradient was generated, resulting in a stress gradient. This eventually inducedthe ellipse dimples in HAZ tensile specimens.

3.3.2. Fatigue pre-crack test fractographic observationAn SEM fractography was also carried out after fatigue pre-crack and fracture toughness test on SENB specimens on SSAW

pipe. The fatigue pre-crack surfaces of three specimens were shown in Fig. 11. The presence of secondary microcracks andslip band cracks was observed in BM, WM and HAZ specimen. This can be considered as the typical fatigue crack surface of

Fig. 9. Crack size for specimens. a0, af: initial and final crack size.

Fig. 10. SEM fractographs of tensile specimens tested in (a) BM, (b) WM and (c) HAZ of the SSAW pipes.

Fig. 11. SEM fractographs of pre-crack procedure tested in (a) BM, (b) WM and (c) HAZ of the SSAW pipes.

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X80 [24,29]. It can be found in Fig. 11 that the fatigue pre-crack surface of WM and HAZ were different from the BM. Thecrack surface of BM was smooth, which had fewer secondary cracks, while it was rough in WM and HAZ. Due to the effectof welding heat input, more brittle phase of martensite/austenite (M/A) constituent formed in WM and HAZ which wouldpromote the formation of secondary cracks and will reduce the toughness [21,29]. So the toughness of WM and HAZ werelower than BM.

3.3.3. Fracture toughness test fractographic observationThe fracture morphology of the SSAW pipes that underwent fracture toughness testing was shown in Fig. 12(a)–(c). The

failure mode was ductile fracture which was due to the ductile nature of X80 steel. Dimples in the fractography can alsobe viewed as evidence supporting this conclusion. Obviously, the tear dimples in Fig. 12 which was ellipse were not the samewith dimples shown in previously tensile tests. In SENB test, as the stress distributed in the vicinity of crack tip wasnonuniform, tear dimples were generated by nonuniform stress in the front of crack tip. Comparing Fig. 12b and c withFig. 12a, it can be found that a lot of abnormal spherical bodies scattered in the dimples of WM and HAZ, while those werescarcely observed in the BM. The abnormal spherical bodies were indicated by arrows in Fig. 12. The spherical bodies

Fig. 12. SEM fractographs of unload/reload procedure tested in (a) BM, (b) WM and (c) HAZ of the SSAW pipes.

Fig. 13. SEM micrographs showing morphology of precipitates on SSAW fracture; (a) and (b) morphology of precipitates at lower magnification in HAZ andWM separately; (c) and (d) morphology of single precipitate at high magnification.

Table 3EDS analysis results of precipitates on the fracture surface in Fig. 13.

Precipitate Element Wt% At%

1 Co 1.27 1.45Ni 1.02 1.17Fe 55.52 66.82Mo 40.90 28.65

2 Co 1.99 2.12Mn 0.80 0.91Fe 67.75 76.43Mo 28.29 18.58

X. Chen et al. / Engineering Fracture Mechanics 148 (2015) 110–121 119

generated in WM and HAZ were precipitates caused by welding heat input. Fig. 13a and b showed the low magnificationimages of abnormal spherical precipitates in HAZ and WM, while Fig. 13c and d were high magnification images. Thechemical composition of area 1 and 2 in Fig. 13a and b were listed in Table 3. From Table 3 it can be concluded that thecomposition of abnormal spherical precipitates were mainly molybdenum compounds. The Mo content of precipitates inHAZ was higher than that in WM. According to the component design codes of X80 pipeline steels, the main function of

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added alloying elements was for precipitation strengthening. While in fact, they aggregated to form micron-sized sphericalproduct and hence losing the original function of precipitation strengthening [30]. The actual original total content ofmolybdenum elements was very low in X80 pipeline steels, but diameter of nearly 10 microns abnormal molybdenumprecipitates appeared in fracture surface which will undoubtedly affect the properties of materials. The main function ofadded molybdenum in X80 pipeline was to promote the formation of acicular ferrite, which will induce microstructurerefinement and achieve high toughness in pipeline steel [31]. However, the formation of abnormal molybdenum sphericalprecipitates in WM and HAZ would serve to hinder the above effect, which would decrease the toughness. Accordingly,the toughness of BM were higher than that of WM and HAZ which in well accordance with results shown in Fig. 8.

4. Conclusions

In the present study, the d–R curves of different locations of LSAW and SSAW joints in API X80 pipeline steels wereobtained through single-edge-notched bending tests with single specimen technique. The toughness of different seamwelded after service was characterized. The critical fracture toughness d0.2BL was compared between different regions ofLSAW and SSAW joints and the fractography of specimens was observed by SEM. It has been observed that ductile fracturebehavior occurs in all these regions. Precipitates on the fracture surface were analyzed by SEM and EDS. The mainconclusions can be summarized as follows:

(1) The fracture toughness of fine grain HAZ (FGHAZ) was higher throughout HAZ of the welded joints which was due tothe finer and more uniform grain size induced by welding heat input.

(2) The fracture toughness of base metal (BM) of LSAW pipes was higher than that of the BM of SSAW pipes.(3) The fracture toughness of weld metal (WM) of LSAW pipes was lower than that of the weld metal of SSAW pipes.(4) The fracture toughness of the base metal (BM) was the highest throughout welded joints in SSAW pipes. As the result

of both the grain coarsening induced by welding heat input in WM and HAZ and of the spherical precipitates ofmolybdenum, there is a decrease in fracture toughness in these locations.

The outcome of this work will provide an important reference to engineering practice and evaluation of pipelinestructures.

Acknowledgments

The authors are grateful for the financial support from the National Natural Science Foundation of China (Nos. 51435012,11172202) and Ph.D. Programs Foundation of Ministry of Education of China (No. 20130032110018). XW also acknowledgessupport from Natural Sciences and Engineering Research Council (NSERC) of Canada.

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