cold cracking susceptibility of x100 pipeline steel
TRANSCRIPT
Cold cracking susceptibility of X100 pipeline steel
Yan Chunyan1, Jiang Xinyi1, Yuan Yuan1,2, Ji Xiulin1, Zhang Kezhao1
严春妍,姜心怡,元媛,纪秀林,张可召
1. College of Mechanical and Electrical Engineering, Hohai University, Changzhou 213022, China;2. CRRC Zhuzhou Locomotive Co., Ltd., Zhuzhou 412001, China
Received 7 June 2019; accepted 12 July 2019
Abstract The y-groove Tekken test has been performed to evaluate the cold cracking susceptibility of X100 pipeline steel. The impact ofpreheating state on the microstructure, distribution of hardness, and the stress-strain state in the welded joint was analyzed. The results showthat X100 pipeline steel reveals a low susceptibility to cold cracking with cracking ratios below 20%. It is found that elevated preheatingtemperature leads to longer cooling time in the welded specimen and ultimately results in a lower cold cracking susceptibility. Preheatingtemperatures of up to 100 ℃ are favorable in decreasing the cold cracking susceptibility due to a relative fine microstructure and low M-Aconstituent amount in coarse grained heat affected zone, a low hardenability, and low-level residual stress and strain. However, excessivepreheating temperatures of 150 ℃ and 200 ℃ lead to grain coarsening, higher M-A constituent amount, higher residual stress level and in-creasing strain level in the Tekken specimens. Preheating temperature above 150 ℃ is not favorable for decreasing the cold cracking suscept-ibility of X100 steel.
Key words numerical simulation, temperature field, residual stress, strain, M-A constituent
0 Introduction
With soaring consumption of petroleum and natural gasin different application fields, the increasing transportationefficiency for long-distance oil-gas transportation pipelinesnecessitates the use of higher strength grade pipeline steel.In China, prevailing majority of the existed pipelines weremanufactured using steels below X80 grade[1 – 2]. Steel X100is required to serve for operation of more efficient pipelinesdue to its higher strength. The X100 pipeline steel is com-mercially produced via thermal mechanically controlledprocessing (TMCP) with higher strength level compared toX80 grade steel. Along with the elevated strength level inmodern high strength steels, however, the heat-affectedzones (HAZ) become comparatively sensitive to cold crack-ing. The cold cracking will shorten the service life span ofX100 steel welded joint and lead to catastrophic failure.
Cold cracking, also referred to as hydrogen-induced
cracking (HIC), usually occurs at relative low temperaturesbelow 150 ℃. Among the welding cracking issues for relat-ive higher strength grade pipeline steel, cold cracking is per-haps the most significant and may happen abruptly, causingdisastrous accidents if not well disposed. The occurrence ofcold cracking is determined by the interaction of threefactors: the hardenability of the steel, hydrogen content anddistribution, and the local restraint stress and strain in thewelded joint. To assess the cold-cracking sensitivity of thesteels, y-groove Tekken test is frequently adopted due to itseasy implementation. Węglowski et al.[3] suggested thecold-cracking tendency was decided by heat input and pre-heating temperature. Yi et al.[4] investigated the cold crack-ing in weld metal, and concluded that acicular ferrite is be-neficial to prevent cold cracking and is more influential thanthe other factors. Chen et al. found that cold cracking couldoccur in pipeline steel welding when celluloid electrode wasused[5]. Tomić et al.[6] studied the susceptibility towards cold
Foundation item: Project was supported by the National Natural Science Foundation of China (Grant No. 51804097) and Fundamental Re-search Funds for the Central Universities (Grant No. 2017B17614).Corresponding author: Yan Chunyan, (1982 – ), PhD, Associate Professor. Mainly engaged in hybrid laser arc welding, weldability of fer-rous metals and numerical simulation for prediction of welding residual stress. E-mail: [email protected]: 10.12073/j.cw.20190607001
Cold cracking susceptibility of X100 pipeline steel 25
cracking for API 5L X80 steel, and found that differentwelding materials showed different sensitivity and coldcracks could form in both the HAZ and the weld metal(WM) zone. According the above review, it is concludedthat there is still lack of available information on cold sus-ceptibility of X100 steel.
The cold-cracking sensitivity of X100 steel based onTekken test results and relative analysis of welding residualstress-strain distribution in weldments were investigated inthis research. Effects of preheating temperature on coldcracks, microstructure in the weldments, hardness distribu-tion were analyzed. Three dimensional (3D) finite elementmodels were developed to evaluate stress distribution and
investigate the effect of preheating temperature on its mag-nitude and distribution in the weldments. The residual straindistribution in the specimens was also analyzed.
1 Experimental
The base metal was provided as 18.4 mm thick plate andthe element contents are provided in Table 1. Cellulosicwelding consumables of grade E10018 with 4 mm core wirediameter were used. The scanning electron microscope(SEM) morphology of the etched X100 steel specimen isdisplayed in Fig. 1. The microstructure of the X100 steel re-veals a mainly ferritic/bainitic structure.
The y-groove Tekken test was carried out with refer-ence to AWS B4-2007. The Tekken test specimens weremachined to the configuration illustrated in Fig. 2. Shieldedmetal arc welding was adopted to conduct the welding pro-cess. The ambient temperature was 20 ℃ and the relativehumidity was 45%. The welding conditions consist of 175 Awelding current, 24 V average arc voltage, and 16 cm/mintravel speed. Five different preheating temperatures (20 ℃,60 ℃, 100 ℃, 150 ℃, and 200 ℃) were selected. Follow-ing welding, the specimens were left at ambient temperat-ure for a minimum period of 48 h before examination forcracks. The test weld area of the assembly was examined forsurface cracks and cross-sectional cracks, and surface crack-ing ratio and cross-sectional ratio were determined and re-corded.
Transverse sections of the test welds under different pre-heating temperatures were etched with a 2 % nital reagent toreveal the microstructure using an optical microscope. Forrevealing the martensite-austenite (M-A) constituents in dif-
ferent regions of the investigated welded specimens, LePera’s reagent was adopted[7]. The morphology and distribu-tion of the M-A islands was studied using field emissionSEM. The statistical analysis of M-A constituent area frac-tion was performed using the MATLAB software based onquantification from at least ten images for each investigatedzone. Composition analysis of the M-A particles and the vi-cinal matrix metal was conducted using energy dispersiveX-ray spectroscopy (EDS). Vickers hardness test was per-formed for all the welded specimens using 500 g load and15 seconds loading time.
2 Finite element model
In this work, the SYSWELD finite element programwas utilized to perform both the thermal and mechanicalanalysis. Due to the asymmetry of the welded joint acrossthe width direction, it is required to establish the whole FEmodel of the joint. In general, it is best to adopt a fullyidentical 3D model in the numerical simulation, but thesemodels are time consuming and costly. Therefore, simpli-fied 3D models were adopted for welding related simula-tion due to the high numerical effort. Fig. 3 shows the 3Dmesh for the simplified y-groove Tekken test model. To de-crease CPU time and storage capacity, the test weld areawas modeled.
The sequentially coupled thermal-mechanical analyticalmethod was employed to save computation time. The weld-ing temperature field was solved first and its history was re-corded. The residual stress field was solved later using thethermal data. In view of calculating precision and effi-ciency, a fine mesh density was selected for the deposit met-al and the HAZ. The nonlinearity of material properties
Table 1 Composition of investigated X100 steel(wt%)C Si Mn P S Nb Ni Mo Cu Fe
0.053 0.260 1.950 0.011 0.003 0.048 0.360 0.240 0.200 Balance
5 μm
Fig. 1 X100 steel microstructure
26 CHINA WELDING Vol. 28 No. 3 September 2019
used in the numerical simulation was considered. In thethermal analysis, convection and radiation boundary condi-tions was also set and optimized.
The heat transfer analysis depends on the solution of ap-propriate heat equation with suitable boundary conditions.The heat input was represented by a double ellipsoid heatsource to simulate the practical welding process. For repres-enting the influence of heat input, all the parameters of theheat source were tried and confirmed based on the weldconfiguration and the thermal cycles. The power densitydistribution of the double ellipsoid heat source was determ-ined by the following equations:
q f =6√
3ηQ f f
π√πabc f
e−3( x2
a2 +y2
b2 +z2
c f2 ) (1)
qr =6√
3ηQ frπ√πabcr
e−3( x2
a2 +y2
b2 +z2
cr 2 ) (2)
where Q is the heat input; x, y, z account for coordinates; ηis the heat source efficiency; qf and qr account for heat dis-tribution in the front and rear heat source respectively; ff and
fr are distribution fractions satisfying ff + fr =2; a, b, cf andcr are shape parameters of the ellipsoids.
3 Results
3.1 Cold cracking ratioThe cracking ratio obtained is given in Table 2. It is
found that the surface cracking ratio is zero for all the testedspecimens. Cold cracks were inspected only on the crosssections. The cracking ratio is below 20% even without pre-heating and decreases with increasing preheating temperat-ure. With a relative low preheating temperature of 60 ℃,the cracking ratio decreases to zero. Generally, if the crack-ing ratio of y-groove Tekken test for low-alloy steels is lessthan 20%, the welded structures will be safe from coldcracks[8].The results indicate adequate cold cracking resist-ance of the X100 steel.
3.2 MicrostructureIt is commonly known that mechanical properties are
mainly determined by its microstructure, and the worsttoughness has been found associated to the coarse grainHAZ (CGHAZ). The microstructure of CGHAZ under dif-ferent preheating temperatures is shown in Fig. 4. TheCGHAZ of the specimens exhibits a typical bainitic struc-
Table 2 Cold cracking ratio result.
No.Preheating
temperature T0 /℃Surface crackingratio Cf (%)
Cross-sectionalcracking ratio Cs(%)
1 20(No preheating) 0 22.5
2 60 0 15.0
3 100 0 0
4 150 0 8.3
5 200 0 16.0
8 in (203 mm) APPROXSECTION A-A
SECTION B-B
60°
60°
t/2
t/2
t/2
t
t/2
t/2
t
2.0
±0.2
(2.0±0.2)
TEST WELD
AREARESTRAINING
WELDS
RESTRAINING
WELDS
A B
BA3-3/16 in
(81 mm)
2-3/8 in
(60 mm)
2-3/8 in
(60 mm)
6 i
n
(152 m
m)
Fig. 2 Shape and dimensions of the Tekken test specimen
zx
y
Fig. 3 FE mesh used for the analysis of Tekken test
Cold cracking susceptibility of X100 pipeline steel 27
ture. Lath bainite and granular bainite appeared under nopreheating condition and relatively low preheating temper-ature. With 150 ℃ and 200 ℃ preheating, less lath bainiteand more granular bainite are observed. Moreover, the in-crease of preheating temperature from 150 ℃ to 200 ℃coarsens the microstructures of CGHAZ in comparison tothat observed in specimen with no preheating and 60~100 ℃ preheating. Besides, higher preheating results inlonger cooling time t8/3 and t8/5, which means slower cool-ing rate. It is believed that the slower cooling rate is re-quired to eliminate cold cracking, which is proved by thelow cracking ratio of 60 ℃ and 100 ℃ preheating.
It was reported by previous researchers that M-A con-stituents in the microstructure seriously impaired toughnessof the high-strength welded joint[9 – 10]. Firstly, the M-A con-stituent amounts were calculated for CGHAZ, fusion zone(FZ), and weld metal under all preheating temperatures re-spectively. Then the carbon contents of M-A constituentswere measured using EDS, and compared with that of the
matrix. Representative morphology of M-A constituents inCGHAZ is exhibited in Fig. 5. Both blocky and elongatedM-A constituents can be recognized. The blocky M-A is-lands mostly formed at the grain boundaries, with the widthof 0.5−2 μm and the length of 1−3 μm. The slender M-Aconstituents are mostly distributed between two ferrite laths,with the width of 0.2−0.5 μm and the length of 1~6 μm. It isreported blocky M-A constituents which are mostly locatedat the prior austenite grain boundaries dominantly promoteimpact toughness deterioration.
The effect of preheating temperature on M-A constitu-ent amount in CGHAZ, WM and FZ is shown in Fig. 6. Asthe preheating temperature increases, the area fractions ofM-A constituent in CGHAZ, WM and FZ tend to decreaseuntil it reaches a minimum with 100 ºC preheating.However, the area fractions of M-A constituent increasewith preheating temperature in the range of 100 ºC to200 ºC. This result indicates that larger amount of M-A con-stituent generates due to excessive preheating. Besides, over
(d) (e)
20 μm 20 μm
(a) (b) (c)
20 μm 20 μm 20 μm
Fig. 4 CGHAZ microstructures in y-groove Tekken test specimens (a) No preheating (b) 60 ℃ preheating (c) 100 ℃ pre-heating (d) 150 ℃ preheating (e) 200 ℃ preheating
(a)
5 μm
(b)
1
2
1 μm
Fig. 5 SEM observation of M-A constituents in CGHAZ (a) Morphology at 2 000× (b) Morphology at 10 000×
28 CHINA WELDING Vol. 28 No. 3 September 2019
preheating can lead to incidence of more blocky M-A con-stituents in CGHAZ which in turn contributes to worse im-pact toughness and higher sensitivity to cold cracking. Fur-thermore, the microstructure coarseness can be enhancedwith high preheating temperature, which further impairs thetoughness.
EDS spectra of the M-A constituent and the adjacentmatrix metal marked with arrow 1 and arrow 2 in Fig. 5b is
displayed in Fig. 7. The comparison of Fig. 7a and Fig. 7bindicates that the compositions of the matrix metal and theM-A constituents are similar. Carbon content analysis basedon EDS shows that M-A constituents are more enrichedwith carbon compared with the matrix metal. The differ-ence of the two is estimated to be about 1.55 wt %.
3.3 HardnessTraditionally, the maximum hardness value in HAZ is
often used to evaluate the cold cracking sensitivity in HAZ.It is suggested that 350 HV is the maximum tolerable HAZhardness for avoiding welding cold cracks. Hardness distri-bution across the welded joints under different preheatingtemperatures is presented in Fig. 8.
It is obvious from Fig. 8 that HVmax value of unpre-heated specimen is 270 HV. The HVmax values under allfive conditions are below 350 HV, which means relativelow hardenability of the HAZs. However, preheated speci-mens exhibit lower HVmax values. Under preheating condi-tion, hardness values in both HAZ and the weld metalslightly decrease. Under preheating temperatures of 150 ℃and 200 ℃, the hardness profiles are relatively flat com-pared to that of the unpreheated specimen. For the preheat-ing temperature of 200 ℃, the HVmax value (240 HV) isslightly below the hardness of the base metal.
0 50 100 150 200 2500
2
4
6
8
10
12
14
16
Fra
ctio
n o
f M
-A c
onst
ituen
t (%
)
Preheating temperature/°C
CGHAZWMFZ
Fig. 6 Effect of preheating temperature on fraction of M-Aconstituent
cps/eV
100
80
60
40
20
0
cps/eV
100
(a)
(b)
80
60SCOMn
Fe Si S Mn Fe
SC
O
MnFe
Si S Mn Fe
40
20
0
0 2 4 6 8 10
Energy/keV
0 2 4 6 8 10
Energy/keV
Fig. 7 EDS spectra analysis result (a) Matrix (b) M-A constituent
Cold cracking susceptibility of X100 pipeline steel 29
3.4 Thermal analysisInvestigation on the temperature distribution and the
thermal cycles in the welding process is of great import-ance, as the thermal cycle data is the basis of many otheranalyses like the prediction of microstructures in weld met-al and HAZ, prediction of residual welding distortion andstress, and susceptibility of the weld joint for cracking, etc.The transient temperature distribution in the weldmentswith preheating of 100 ℃ is illustrated in Fig. 9. The tem-perature distribution is unsteady at the initial stage of weld-ing (t=5.3 s), and gradually becomes steady (t=15 s). Due toa preheating of 100 ℃, the peak temperature in the welding
process can reach 2 488 ℃, which is much higher than thepeak temperature of 1 888 ℃ with no preheating. It is alsoobvious that the temperature distribution in the weld pieceis nonsymmetrical due to the asymmetrical Y joint shape.
Fig. 10a and Fig. 10b show the thermal cycles at the loc-ations which are 3.0 mm from fusion-line in 1/2 weld thick-ness on both sides of the weld, respectively. The overalltemperature is high at a high preheating temperature.
The cooling times t8/5 and t8/3 of a node in HAZ wereplotted against preheating temperature, as shown in Fig. 11.With higher preheating temperature, the cooling rates inboth HAZ and WM decrease, changing the HAZ micro-structure accordingly. Besides, a high preheating temperat-ure enlarges the austenite formation and may lead to ingrain coarsening in the HAZ.
3.5 Mechanical analysisThe residual stress distribution in the weldments under
different preheating conditions is provided in Fig. 12 toFig. 14. It is apparent that the residual stress distribution inthe entire joint is clearly uneven due to the asymmetrical Y
−10 −5 0 5 1050
100
150
200
250
300
no preheating
60°C preheating100°C preheating150°C preheating200°C preheating
Distance from the weld centerline/mm
Har
dnes
s (H
V10
)
Fig. 8 Hardness distribution in the welded joints
1 939.665(a)
(b)
1 816.8091 693.9521 571.0951 448.2391 325.3821 202.5261 079.669956.812833.956711.099588.242465.386342.529219.67396.816
1 882.4001 763.2181 644.0371 524.8551 405.6741 286.4921 167.3111 048.129928.947809.766690.584571.403452.221333.040213.85894.677
Tra
nsi
ent
tem
per
ature
/°C
Tra
nsi
ent
tem
per
ature
/°C
Fig. 9 Transient temperature distribution (a) 5.3 s (b) 15.0 s
0 50 100 150 200
Time/s
250 300 350 400
0
200
400
600
800
1000
1200
1400
1600
1800no preheating
60°C preheating
100°C preheating
150°C preheating
200°C preheating
no preheating
60°C preheating
100°C preheating
150°C preheating
200°C preheating
Tem
per
ature
/°C
0 50 100 150 200
Time/s
250 300 350 400
0
200
400
600
800
1000
1200
1400
1600
Tem
per
ature
/°C
(a)
(b)
Fig. 10 Simulated thermal histories in the weld (a) non-symmetrical side (b) symmetrical side
30 CHINA WELDING Vol. 28 No. 3 September 2019
joint shape. Tensile stress occurs in both WM and HAZclose to fusion boundary while compressive stress occurs inbase metal remote from the weld center. It can be seen fromthe figures that the stress levels of the stop positions ofwelding are higher than that of the start positions.
The maximum magnitudes of two stress componentsand the equivalent stress were described in Fig. 15. Themagnitude of longitudinal stress is higher than that of thetransverse stress. Orientation of maximum principal stress isclose to the weld direction. It is therefore believed that themain crack propagates along the transverse weld direction.This agrees with the experimental findings of the presentwork. As one can see, Von Mises stress magnitude de-creases with preheating temperature until the preheatingtemperature reaches 100 ℃, and then increases with the pre-heating temperature. The longitudinal residual stress mag-nitude decreases with the preheating temperature until
0 50 100 150 200 2500
10
20
30
40
50
60
70
Cooli
ng t
ime/
s
Preheating temperature/°C
t8/5
t8/3
Fig. 11 Impact of preheating temperature on cooling time
610.042(a)
(b)
(c)
569.781529.519489.258448.997408.736368.475328.214287.952247.691207.430167.169126.90886.64646.3856.124
582.430544.464506.499468.533430.567392.601354.635316.670278.704240.738202.772164.807126.84188.87550.90912.944
557.653521.398485.143448.889412.634376.379340.124303.869267.614231.359195.104158.849122.59486.33950.08413.829
Von M
ises
str
ess/
MP
aV
on M
ises
str
ess/
MP
aV
on M
ises
str
ess/
MP
a
Fig. 12 Distribution of Von Mises stress with different pre-heating conditions (a) No preheating (b) 100 ℃ preheating(c) 150 ℃ preheating
651.035(a)
(b)
(c)
589.426527.817466.207404.598342.988281.379219.770158.16096.55134.942−26.688−88.277−149.886−211.496−273.105
613.977554.502495.027435.551376.076316.601257.126197.651138.17678.70119.225−40.250−99.725−159.200−218.675−278.150
617.974558.585499.196439.807380.418321.029261.640202.251142.86283.47324.084−35.305−94.694−154.083−213.472−272.861
Longit
udin
al r
esid
ual
str
ess/
MP
aL
ongit
udin
al r
esid
ual
str
ess/
MP
aL
ongit
udin
al r
esid
ual
str
ess/
MP
a
Fig. 13 Distribution of longitudinal residual stress withdifferent preheat conditions (a) No preheating (b) 100 ℃preheating (c) 150 ℃ preheating
Cold cracking susceptibility of X100 pipeline steel 31
150 ℃ and then increases with the preheating temperature.The calculated equivalent strain for non-preheated spe-
cimen is shown in Fig. 16. It can be seen that the higheststrain appeared at the symmetrical side of the weld root, in-dicating high possibility of cracking occurrence in this posi-tion. The variation of strain distribution under different pre-heating conditions is shown in Fig. 17. Under a relative low
394.788(a)
(b)
(c)
345.416296.043246.671197.299147.92698.55449.182−0.191−49.563−98.935−148.307−197.680−247.052−296.424−345.797
450.621395.745340.869285.994231.118176.242121.36666.49111.615−43.261−98.137−153.012−207.888−262.764−317.639−372.515
472.803415.715358.626301.537244.449187.360130.27273.18316.094−40.994−98.083−155.172−212.260−269.349−326.438−383.526
Tra
nsv
erse
res
idual
str
ess/
MP
aT
ransv
erse
res
idual
str
ess/
MP
aT
ransv
erse
res
idual
str
ess/
MP
a
Fig. 14 Distribution of transverse residual stress with dif-ferent preheat conditions (a) No preheating (b) 100 ℃ pre-heating (c) 150 ℃ preheating
0 50 100 150 200 250
200
400
600
800
Res
idual
str
ess/
MP
a
Preheating temperature/°C
Von Mises stressLongitudinal residual stressTransverse residual stress
Fig. 15 Variation of maximum residual stresses with pre-heating temperature
0.0570.0530.0490.0450.0410.0380.0340.0300.0260.0230.0190.0150.0110.0080.0040
Equiv
alen
t st
rain
Fig. 16 Calculated distribution of equivalent strain
(a)
−80 −60 −40 −20 0 20 40 60−0.5
00.51.01.52.02.53.03.54.04.5
No preheating
60°C preheating100°C preheating150°C preheating200°C preheating
No preheating
60°C preheating100°C preheating150°C preheating200°C preheating
Str
ain (
%)
Str
ain (
%)
Distance from the weld centerline/mm
(b)
0 10 20 30 40 50 60 70 80 90−0.5
0
0.5
1.0
1.5
2.0
2.5
3.0
3.5
4.0
Distance from the welding strart position/mm
Fig. 17 Effect of preheating temperature on equivalentstrain distribution (a) Width direction (b) Length direction
32 CHINA WELDING Vol. 28 No. 3 September 2019
preheating temperature of 60 ℃, the distribution of strain inthe welded specimen becomes more even and the residualstrain level is depressed. Yet, under a higher preheatingtemperature of 150 ℃, the residual strain level ascends andstrain localization adjacent to WM is developed. This im-plies that the relationship between the preheating temperat-ure and mechanical properties is not monotonous. Up to100 ℃, preheating is beneficial and a preheating temperat-ure above 150 ℃ is not appropriate for cracking prevention.
4 Conclusions
In the present paper the cold cracking susceptibility ofHAZ in X100 pipeline steel was evaluated using the Tekkentest. Effects of preheating on distribution of microstructures,the hardness distribution, the welding temperature historiesand the stress-strain state were discussed. The useful con-clusions are summarized as below.
(1) Experimental Tekken test results show that the sus-ceptibility of investigated X100 steel to cold cracking is lowand the cracking ratio is significantly decreased with pre-heating.
(2) The microstructures in the welded joints are signific-antly influenced by the preheating temperature. Preheatingat above 150 ℃ leads to grain coarsening and increase infraction of M-A constituents.
(3) Preheating of 60 ℃ and 100 ℃ results in a reduc-tion in the magnitude of residual stress and strain level.
However, preheating above 150 °C is not appropriatefor reducing the cold cracking susceptibility of X100 steeldue to increasing level of residual stress and strain.
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