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ABSTRACT. The purpose of this study isto develop an appropriate numericalmodel for full encirclement, sleeve repair

welding of in-service gas pipelines and toinvestigate the effects of in-service weld-ing conditions. An axisymmetric finite el-

ement model was used to calculate thetemperature distribution, maximum HAZhardness, and the distribution of residualstress and plastic strain during multipasssleeve fillet welding of in-service API 5LX65 pipelines of 14.3 mm thickness. Ex-periments were also conducted for sleeverepair welding on pipe with internal pres-sure. The calculated geometry of the fu-sion zone and HAZ was in good agree-ment with the macrostructures of sleeverepair fillet welds. The calculated maxi-mum HAZ hardness was in good agree-ment with the measured value. The effectof gas flow rate on the maximum HAZhardness and the effects of internal pres-sure on residual stress and plastic straindistribution were investigated. Risk ofmelt-through and susceptibility to coldcracking were also estimated using a sim-plified analysis.

Introduction

In buried natural gas pipelines, defectscan occur as a result of construction faults,corrosion, third-party interference, andground movement. When a segment of apipeline is found to be defective, one of

the repair methods is to vent the gas withinthe pipeline and cut out the defective seg-ment after shutting down the pipeline.However, the cost is extremely high interms of venting and stopping the gas sup-ply. Therefore, most pipeline companieshave developed repair methods withoutstopping flow through the pipe. These in-service repair methods are widely used

throughout the natural gas, petroleumand petrochemical industries (Refs. 14).Welding onto a gas pipeline in operation,known as in-service welding, is a fre-quently employed repair technique. Thedirect deposition of weld metal, sleeve re-

pair welding, and hot-tap welding are typ-ical examples of in-service welding.There are two important concerns with

welding on in-service pipelines. The firstconcern is the possibility of melt-throughdue to localized heating, leading to loss ofmaterial strength on the inner surface ofpipe during the welding process. The pipe

wall can burst under internal pressure ifthe loss of the strength is large.

The second concern is the high coolingrate of the weld as a result of flowing gasquickly removing heat from the pipe wall.The high cooling rate can promote the for-mation of heat-affected zone (HAZ) mi-crostructure with high hardness, makingthese weldments susceptible to cold crack-ing and sulfide stress cracking in sour ser-

vice. The rapid cooling can be compen-sated by increasing heat input, but theincreased heat input can promote weldpenetration and the possibility of melt-through. Thus, suitable welding proce-dures must ensure optimal HAZ hard-ness, no melt-through, and proper heatinput.

Over the 12-year period between 1978and 1990, a series of hot-tap welding re-search programs was carried out at Bat-

telle and Edison Welding Institute (EWI)

(Refs. 57). A thermal analysis model,based on the two-dimensional finite dif-ference method, was developed to simu-late sleeve and direct-branch in-service

welding. The cooling rate of the HAZ andthe maximum inside surface temperature

were predicted for a given pipe geometryand a set of welding parameters. The HAZhardness was estimated from the coolingrate at 540C (1000F) and the carbonequivalent of the material. Taking Vickershardness 350 HV as a low cracking poten-tial, the maximum inside surface temper-ature of 982C (1800F) was regarded asthe limiting maximum temperature to pre-

vent melt-through for low-hydrogen-elec-trode welding. However, the estimation ofHAZ cracking was highly simplified andnot applicable to todays high-strengthpipeline steels. Also, melt-through limits

were predicted only by the maximum in-side surface temperature and the effectsof internal pressure and stresses were notconsidered.

The finite element method offers acomputational tool for simulation andanalysis of in-service welding of gaspipelines (Refs. 810). In addition to in-

vestigating the transient thermal field andcooling rates by thermal analysis, stressfield and melt-through have been ana-lyzed using a thermoelastic-plastic modelor thermoelastic-viscoplastic model by thefinite element method. Sabapathy et al.(Ref. 10) predicted melt-through using a

thermoelastic-plastic model and also de-veloped an alternative and convenientmethod to predict the bursting pressureduring in-service fillet welding.

Presently, repair welding on an in-ser-vice natural gas pipeline is a primary con-cern of Korea Gas Corporation(KOGAS). Therefore, a systematic study

was undertaken for the repair welding ofAPI 5L X65 main pipeline operatingunder the internal pressure of 6.9 MPa (70kgf/cm

2), with a diameter of 762 mm and athickness of 17.5 or 14.3 mm. Following aprevious study of bead-on-plate welding

and repair welding by direct deposition of

I.-W. BANG, Y.-P. SON, and K. H. OH arewith the School of Materials Science andEngineering, College of Engineering, SeoulNational University, Seoul, Korea. Y.-P. KIMand W.-S. KIM are with R & D Center, Korea

Gas Corp., Ansan, Korea.

KEY WORDS

Sleeve Repair WeldingFinite Element MethodMelt-ThroughCold CrackingResidual StressesMaximum HAZ Hardness

Numerical Simulation of Sleeve RepairWelding of In-Service Gas Pipelines

A model was developed to predict conditions for sucessful sleeve repairof in-service gas pipelines

BY I.-W. BANG, Y.-P. SON, K. H. OH, Y.-P. KIM, AND W.-S. KIM

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weld metal (Ref. 11), a numerical and ex-perimental study was conducted on in-ser-

vice sleeve repair welding. The objectiveof this work was to develop an appropriatenumerical model for sleeve repair weldingof in-service gas pipelines and to investi-

gate the effects of in-service welding con-ditions.An axisymmetric finite element model

was developed to simulate multipasssleeve fillet welding on in-service API 5LX65 pipelines of 14.3 mm thickness. Themodel was used to predict the tempera-ture distribution, maximum HAZ hard-ness, and the distribution of residual stressand plastic strain. The experimental study

was also conducted on sleeve repair weld-ing with internal pressure applied. Thecalculated maximum HAZ hardness wasused to predict the occurrence of coldcracking, and the effect of gas flow rate onthe maximum HAZ hardness was investi-gated. A detailed evolution of residualstress and plastic strain and the effects ofinternal pressure on their distribution

were investigated. An analysis of single-pass fillet welding was also carried out toassess the allowable heat input for sleevefillet welding.

Experimental Procedure

Full encirclement, sleeve repair weld-ing on in-service gas pipelines is a repair

welding method for relatively large de-

Fig. 1 Schematic of sleeve repair welding on pipe.

Fig. 2 Configuration of pressurized sleeve repair welding test equip-ment at the R and D center of KOGAS Corp.

Table 1 The Chemical Compositions (wt-%) of the Pipe, Sleeve, and Weld Metal

Elements C Mn P S Si Nb V Ti Ni Cr Mo CIIW(a)

Pipe (sleeve) 0.06 1.69 0.020 0.001 0.25 0.034 0.062 0.022 0.018 0.027 0.001 0.36Weld metal 0.04 1.41 0.011 0.003 0.59 0.004 0.008 0.021 0.612 0.043 0.276 0.38

(a) CIIW= C + Mn/6 + (Ni + Cu)/15 + (Cr + Mo + V)/5.

Table 2 The Welding Conditions for Sleeve Fillet Welding

Pass Process Electrode Current Voltage Welding Speed Heat InputNumber Diameter (A) (V) (cm/min) (kJ/mm)

(mm)

1 GTAW 2.4 194 23 16 1.67

2 SMAW 2.6 96 29 5 3.343 SMAW 2.6 96 30 9 1.924 SMAW 3.2 120 30 5 4.325 SMAW 3.2 117 30 10 2.116 SMAW 2.6 100 30 7 2.577 SMAW 3.2 122 31 5 4.548 SMAW 3.2 118 30 9 2.36

Note: An internal pressure of 4.4 MPa was applied with no gas flow condition.

Table 3 The Welding Conditions for Sleeve Butt-Joint Welding

Pass Process Electrode Current Voltage Welding Speed Heat Input

Number Diameter (A) (V) (cm/min) (kJ/mm)(mm)

1 GTAW 2.4 226 25 13 2.612 SMAW 3.2 153 36 12 2.753 SMAW 3.2 152 35 11 2.904 SMAW 3.2 147 34 15 2.005 SMAW 3.2 148 34 13 2.016 SMAW 3.2 146 34 17 1.757 SMAW 3.2 145 35 17 1.798 SMAW 3.2 147 34 18 1.679 SMAW 3.2 143 33 17 1.6710 SMAW 3.2 137 32 17 1.5511 SMAW 3.2 136 32 18 1.45

Note: An internal pressure of 4.4 MPa was applied with no gas flow condition.

Fillet Weld

Fillet Weld

Pipe

Sleeve

Butt-Joint Weld

Butt-Joint Weld Pipe

Sleeve

Sleeve

Pipe

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fects. The schematic illustration of thissleeve repair welding method is shown inFig. 1. Two sleeves are attached to the pipe

around damaged sections and then cir-cumferential fillet welding and longitudi-nal butt-joint welding are performed.

Figure 2 shows the configuration of thepressurized sleeve repair welding testequipment. Both ends of the pipe were

welded using the cap made of WPHY 65,and a pressure gauge and a fitting for mea-suring the gas temperature were installed.In order to simulate the in-service condi-tion, an internal pressure of 4.4 MPa wasapplied using a nitrogen gas.

The diameter and the thickness of thepipe are 762 and 14.3 mm, and the repair

sleeves were made by expanding radially a762-mm-diameter pipe with a 17.5-mmthickness. The chemical compositions ofthe API 5L X65 pipe, sleeve, and weldmetal are listed in Table 1.

Figures 3A and B show the weld jointshape and welding sequence of sleeve fil-let welds and groove welds, respectively.Manual shielded metal arc welding(SMAW) and gas tungsten arc welding(GTAW) processes were applied to sleeverepair welds. The welding conditions forfillet welding and butt-joint welding areshown in Tables 2 and 3, respectively.

In the butt-joint welds, melt-through is

a matter of little concern because the weldmetal is not in direct contact with the pipe.In this study, fillet welds are in direct con-

tact with a pipe with an internal pressureof 4.4 MPa, but melt-through can beavoided by controlling heat input. Small-diameter electrodes are recommended for

welds in contact with the pipe becauseheat input can be reduced with them (Ref.12). Small-diameter electrodes were usedin this experiment. A 2.4-mm, ER70S-Gelectrode was used for pass 1, and a 2.6-mm low-hydrogen E9016-G electrode wasused for passes 2, 3, and 6. A 3.2-mm low-hydrogen E9016-G electrode was used forthe layers that were not in direct contact

with the pressure-containing pipe. Pre-

heating was not applied considering thefield welding conditions.The weld joints were sectioned and

polished to observe the geometry of theweld metal and the HAZ. The penetrationof weld metal and depth of HAZ weremeasured. Microhardness measurements

with a 500-g load were made in base metal,coarse-grained HAZ (CGHAZ), fine-grained HAZ (FGHAZ), and weld metalof pipe and sleeve.

Computational Procedure

The commercial finite element code

ABAQUS (Ref. 13) was used for the ther-mal and mechanical analysis. A sequen-tially coupled analysis of thermal and me-

chanical analyses was performed. Anaxisymmetric model was used to calculatethe distributions of temperature, residualstresses, and plastic strain during the mul-tipass sleeve fillet welding of in-service

API 5L X65 pipeline. The welding condi-tions listed in Table 2 were used. An analy-sis of a single-pass fillet weld applied withSMAW was also carried out to assess theallowable heat input.

Finite Element Mesh

The thermal and mechanical response

of a weldment is a three-dimensionalproblem that requires a considerableamount of computational time. The com-putational time required to simulate mul-tipass welding increases in proportion tothe number of weld passes. Therefore, it isnecessary to develop cost-effective proce-dures to reduce computational time whilepreserving the accuracy of the solution.

An assumption common for most analysisof circumferential multipass welding is theassumption of axisymmetry, that is, weld-ing heat is assumed to be deposited at thesame time around the circumference. This

assumption, which strongly reduces the

Fig. 3 Joint shape and pass sequence of sleeve

repair welds. A Fillet welds; B butt-joint

welds.

B

A

Fig. 4 Axisymmetric finite element mesh for sleeve fillet welding.

(a) multipass welding (b) 1-pass welding

Symmetry line

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size of the FE model and the CPU timeneeded, can be justified by observation inexperiments that residual stresses in cir-cumferential groove welds are reasonablyaxisymmetric (Refs. 14, 15). So, a two-dimensional axisymmetric model was usedin this study.

Based on the geometry and weld passsequence shown in Figs. 2 and 3, an ax-isymmetric finite element mesh for sleevefillet welding was made as shown in Fig. 4.In Fig. 4, the r-coordinate and z-coordi-nate correspond to the radial and axial di-rection of the pipe, respectively, and the

welding direction is hoop direction. The x-coordinate was defined along the innersurface of pipe in order to describe the dis-tributions of residual stresses and strain.The origin of the axis corresponds to theposition at the inner surface of the pipebelow the weld root.

Actual circumferential sleeve filletwelding is performed sequentially or si-multaneously at both left and right sides of

a sleeve. Assuming the welding is per-formed simultaneously at both left andright sides of a sleeve, half of the geome-try was modeled by applying the symmetrycondition at the centerline of the axial di-rection. The finite element model of Fig.4A consists of 750 quadrilateral elementsand 919 nodes. A refined finite elementmesh was used in and near the weld re-gion.

Modeling of one-pass welding is an ef-ficient way to investigate the effects ofheat input on penetration, melt-through,and cold cracking. The finite elementmesh for single-pass welding near the weldregion is shown in Fig. 4B. The geometryof the pipe and sleeve is equal to that ofthe finite element mesh of Fig. 4A andonly the mesh of weld metal was changedin the simple triangular geometry.

The area of weld metal (weld bead rein-forcement) was assumed to be proportionalto heat input. The SMAW process was usedfor modeling single-pass welding.

Thermal Analysis

To simulate arc heating effects effi-ciently during multipass welding, theequivalent heat input can be assumed asthe combination of both surface and bodyheat flux components (Ref. 14). The totalheat input can be given as follows:

Q = Qs

+ Qb

= E I (1)

where QS and QB are the heat input due tosurface flux and body flux, respectively, is the arc efficiency,E is voltage, andIiscurrent. The ratio of QB/QS can be ad-

justed to achieve an accurate representa-tion of the fusion zone. In this study, thetotal heat input was assumed to be 20% ofsurface flux and 80% of body flux from thecomparison between the experimentaland the calculated size of the fusion zone.The arc efficiencies used in the analysisare 0.75 for SMAW and 0.40 for GTAW.

The surface fluxqS and body fluxqB are

Fig. 5 Temperature-dependent material properties of API X65 pipe. A Thermal properties; B mechanical properties.

Fig. 6 Comparison of (A) calculated peak temperature distributions and (B) experimental macrostructures at the welding conditions listed in Table 2.

A

A

B

B

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generally represented in the form of aGaussian distribution as follows (Ref. 16):

(2)

(3)wherea, b, andc are the semicharacteris-tic arc dimensions inx, y, andz direction.This heat source model has often beenused to approximate simple weldingprocesses carried out in the flat position,

i.e., welding horizontally in a straight lineon a horizontal flat plate with the elec-trode perpendicular to the plate (Ref. 10).

A characteristic of a low-hydrogen elec-trode is often a shallow penetration, whichsuggests a heat distribution flatter andmore evenly distributed than Gaussian.Sabapathy et al. (Ref. 10) modified theGaussian heat source model by changingthe exponential terms and simulated in-service welding. In this study, the heatfluxes of Equations 2 and 3 were modifiedby assuming the uniform distribution ofheat fluxes in width and thickness direc-

tion in order to simulate the shallow pen-

etration. The surface and body fluxes canbe given as follows:

(4)

(5)whereA is the cross-sectional area of thefusion zone. The values ofa and c werechosen as the half width of the fusion zone.The fixed z-coordinate is related to themoving coordinate as follows:

z = ( t) (6)

where is the welding speed and is a lagfactor to define the position of the heatsource at time t = 0.

The element birth technique was usedto model the multipass weld metal deposi-tion effects (Ref. 13). The elements ofeach weld pass were meshed separatelyand then connected to adjacent passes andthe base metal mesh with contact surfaces.The weld metal finite elements and con-tact surfaces were inactive at the begin-ning of the analysis, and then activated atthe specified time to simulate the deposi-

qQ

Ac

z

cB

B=

3 3 2

2exp

qz

cs

Q

ac

s=

32

2

3

exp

qQ

abc

x

a

y

b

z

cB

B=

6 3 3 3 32

2

2

2

2

2

exp

qQ

ac

x

a

z

cs

s=

3 3 32

2

2

2exp

Table 4 Thermophysical Data of Natural

Gas (Ref. 7)

Property Value

Specific heat 2245.36 J/kg-K of gas, CpgDensity of gas, g 36.9 kg/m3

Thermal Conductivity 3.5354 X 102

of gaskg w/m-KViscosity of gas, g 1.11 X 105 Pa-s

Location Hardness (HV)

Weld metal SMAW (4, 5) 219.0GTAW (11) 211.0

CGHAZ Pipe (3, 10) 235.0Sleeve (6, 12) 213.0

FGHAZ Pipe (2, 9) 202.5sleeve (7, 13) 196.5

Base metal Pipe (1) 178.0Sleeve (8) 179.0

Fig. 7 Variation of temperature with time at the position where the peaktemperature is maximum on the inner surface of the pipe at the weldingconditions listed in Table 2.

Fig. 8 Effect of gas flow rate on maximum HAZ hardness for the weld-ing conditions in Table 2 with the exception of the gas flow condition.

Table 5 Vickers Hardness of Sleeve Fillet Welds

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tion sequence of weld passes.The temperature-dependent thermal

conductivity and specific heat of pipe ma-terial were obtained from the chemicalcomposition of the pipe material, asshown in Fig. 5A (Ref. 17). The latent heatof fusion 2.1 x 109 J/m3was used to model

the solid-liquid transformation.The inner and outer surfaces of the

pipe and sleeve are subjected to a combi-nation of natural convection, forced con-

vection, and radiation at high tempera-ture. The natural convection and radiationboundary condition were used at the out-side of the sleeve and the outer surface ofthe pipe. The heat transfer mechanism atthe inner surface of the pipe was taken asa natural convection, a radiation boundarycondition at the condition of no gas flow,and as a forced convection boundary atthe condition of gas flow. The convective

heat transfer between pipe and flowing

gas was characterized with a heat transfercoefficient, hg, determined from the fol-lowing dimensionless relation (Ref. 18):

(7)

whereD is the pipe diameter, gis the ve-locity of the gas,

gis the viscosity of the

gas, andhg is the heat transfer coefficientbetween pipe and flowing gas. Thermo-physical data to calculate hg are listed inTable 4 (Ref. 7).

Mechanical Analysis

Temperature histories from the ther-mal analysis were given as inputs for themechanical analysis. Similar to the ther-mal model, the filler metal passes weretied to adjacent passes and the base metal

with contact surfaces. It is also possible to

use the element birth technique in the me-chanical analysis. However, this may giveserious numerical problems in the me-chanical analysis because inactive ele-ments at the boundary between old, al-ready deposited material, and newmaterial deposited in the particular weld

pass may be strongly distorted when beingactivated. Attempts to fit the undeformedfiller material to the deformed geometry

will lead to a stress buildup in this stress-free material and a redistribution of resid-ual stresses from previous passes. More-over, the computation will break downimmediately if the magnitudes of defor-mation of the inactive elements exceed thesize of the elements.

Troive and Johnson (Ref. 19) modifiedthe common element birth technique

where the degrees of freedom for allunique birth nodes are fully constrained

until the time of birth by changing the con-

h D

k

D C

k

g

g

g g

g

pg g

g

=

0 023

0 8 0 4

.

. .

Fig. 9 Contour plots of residual stress distributions. A Axial stress; B hoop stress using the welding conditions listed in Table 2.

Fig. 10 Evolution of residual stress along the inner surface of the pipe. A Axial stress; B hoop stress using the welding conditions listed in Table 2.

A

A

B

B

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ditions for all birth nodes to be totally free,even if not activated. By this technique,the deformation field of the distorted un-activated elements will be smoothed outover the area of unactivated elements, andthereby high magnitudes of element dis-tortions will be avoided. Another ap-proach is to keep the unactivated elementsat a high temperature. These elements willthen have a low stiffness and not affect anydeformation mechanism even though they

were present. The latter method was usedin the present study.

The thermoelastic-plastic constitutivemodel was used to describe the deforma-tion behavior. Figure 5B shows the tem-

perature-dependent mechanical proper-ties, i.e., Youngs modulus E, yield stressy, hardening modulus ET, poissons ratio, and thermal expansion coefficient .The true yield stress of pipe material atroom temperature is 503 MPa. The ther-mal expansion coefficient was set to bezero above melting temperature to pre-

vent stress being applied to the liquid.The pressure loading by internal pres-

sure was applied at the first step of the me-chanical analysis, and then the thermalloading from the temperature histories wasapplied, preserving the pressure loading.

Assessment of Cold Cracking

Cold cracking, or hydrogen-inducedcracking, is one of the most serious prob-lems encountered during welding. De-pending on the location of the cold crack,it is called underbead crack, root crack, ortoe crack. Major factors that contribute tocold cracking are susceptible microstruc-ture of high hardness, hydrogen content,and tensile restraint stresses.

These factors mutually interact, and itis difficult to simply evaluate the effect

they may have on cold cracking suscepti-

bility. Generally, themaximum HAZ hard-ness is regarded as anapproximate index forsusceptibility to coldcracking. The maxi-mum HAZ hardness isoften limited to 350HV in welding fabrica-tion of offshore struc-tures and line-pipes foravoidance of coldcracking (Refs. 20, 21).

A previous studyshowed HAZ mi-crostructures with

hardness of 248 HVand higher are suscep-tible to stress corrosioncracking (SCC) when the welding is doneon in-service pipelines used to transportmildly sour gas (Ref. 22).

The formulas for estimating maximumHAZ hardness can be separated into threedifferent groups (Ref. 23): 1) formulasthat estimate hardness entirely from car-bon equivalents; 2) formulas that estimatehardness from various carbon equivalentsin conjunction with welding parameters,such as cooling time t8/5; and 3) formulas

that estimate hardness by means of vari-ous carbon equivalents in conjunction

with microstructure. For the first group,the application of the formulas must be re-stricted to the circumstances for whichthey were developed. The formulas be-longing to the last group are not as versa-tile as the formulas from groups 1 and 2because detailed information on themartensite content, which only can be ob-tained by detailed metallographic exami-nation, is required in addition to chemicalcomposition. The formula proposed byKasuya et al. has the widest applicable

range among the formulas belonging to

group 2 and was reported to show goodcorrelation between measured and esti-mated maximum HAZ hardness (Ref. 24).

The maximum HAZ hardness in theformula proposed by Kasuya et al. is givenas follows:

HV = (HM+ HB)/2(HM HB).

arctan (X)/2.2 (8)

where HM is the hardness value where

martensite volume fraction reaches 100%in CGHAZ and HB is where martensite

volume fraction becomes almost 0% inCGHAZ. X is defined by the followingequation:

X = 4 log(/M)/log(B/M)2.0 (9)

where is the cooling time between 800and 500C (t8/5), and M and B are thecooling times corresponding to HM andHB, respectively. The four constants (HM,M, HB and B) depend on the chemicalcomposition of steel (wt-%) and are de-

fined from the experimental data of vari-

Fig. 11 Evolution of equivalent plastic strain along the inner surface ofthe pipe using the welding conditions listed in Table 2.

Fig. 12 Effect of internal pressure on residual stress distribution along

the inner surface of the pipe. A Axial stress; B hoop stress using thewelding conditions listed in Table 2.

A

B

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ous steels as follows:

HM= 884C(10.3C2) + 297 (10)

M= exp(10.6CE1 4.8) (11)

where CE1 = Cp + Si/24 + Mn/6 + Cu/15+ Ni/12 + Cr(1 0.16 Cr)/8 + Mo/4 +H; Cp = C(C 0.3), C/6 + 0.25 (C >0.3); H = 0(B 1 ppm), 0.03fN(B = 2ppm), 0.06fN(B = 3 ppm), 0.09fN(B 4ppm); andfN= (0.02 N)/0.02

HB = 145 + 130tanh(2.65CE2 0.69) (12)

where CE2 = C + Si/24 + Mn/5 + Cu/10+ Ni/18 + Cr/5 + Mo/2.5 + V/5 + Nb/3

B = exp(6.2CE3 + 0.74) (13)

where CE3 = Cp + Mn/3.6 + Cu/20 +Ni/9 + Cr/5 + Mo/4.

Equations 813 were used to predictmaximum HAZ hardness from the chem-ical composition and calculate coolingtime from thermal analysis in this study.The occurrence of cold cracking was as-sessed by the comparison between the cal-culated maximum HAZ hardness and thelimiting hardness of 350 HV. This ap-proach to predict maximum HAZ hard-ness and the occurrence of cold cracking

was successfully applied in the analysis ofbead-on-plate welding of API 5L X65plates with the various welding conditions(Ref. 25).

Results and DiscussionTemperature Distributions

The geometry of the fusion zone andHAZ can be predicted from the peak tem-perature distributions. The peak tempera-

ture distributions were

obtained from the cal-culated transient tem-perature field. The fu-sion zone isdetermined by themelting temperatureand the geometry ofthe HAZ can also bedetermined by Ae3temperature. TheHAZ consists of sev-eral subzones, whichare normally definedby the peak tempera-

tures. Lundin et al.(Ref. 26) reported theaverage peak tempera-tures of 1316C and954C are commonlyused to represent CGHAZ and FGHAZ,respectively.

Figure 6A shows the calculated peaktemperature distributions for the weldzone. The isothermal lines of 1515.6C,1316.0C, and 870.8C correspond to melt-ing, CGHAZ, and Ae3 temperatures, re-spectively. Figure 6B shows themacrostructures of fillet welds. It can beseen that the size and shape of the fusionzone and HAZ observed in macrostruc-tures are in good agreement with theisothermal lines of melting temperatureand Ae3 temperatures, respectively. Fromthe agreement between calculated and ex-perimental weld geometries, it is knownthe temperature distributions for multi-pass sleeve fillet welding can be satisfacto-rily calculated from the model.

Figure 7 shows the variation of tem-perature with time at the position wherethe peak temperature is maximum on theinner surface of the pipe. The temperatureprofile has higher values when the welds

are in contact with the pipe, such as passes2, 3, and 6, rather than when in contact

with the sleeve. The highest peak temper-ature that corresponds to the maximuminside surface temperature was 515C atpass 6.

This value is much lower than 982C,which is the limiting maximum inside sur-face temperature for preventing melt-through (Ref. 7). As shown in Fig. 6B, ex-

cessive deformations or melt-througharound welds were not observed in weld-ing the pipe with an internal pressure of4.4 MPa. The heat input of welding pass 4

was 4.32 kJ/mm due to low welding speed,but melt-through was not found numeri-cally nor experimentally.

Hardness Distributions and Effect of Gas

Flow Rate on Maximum HAZ Hardness

The measured hardness values for basemetal, CGHAZ, FGHAZ, and weld metalare shown in Table 5. The hardness in theCGHAZ of the pipe has the highest value,

Fig. 13 Effect of internal pressure on equivalent plastic strain distribu-tion along the inner surface of the pipe using the welding conditions listedin Table 2.

Fig. 14 Variation of (A) maximum inside surface temperature and (B)maximum HAZ hardness with heat input and gas flow rate for a single-

pass sleeve fillet weld.

A

B

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235 HV, which is much lower than 350 HV.The calculated maximum HAZ hardnessis 242 HV and in good agreement with themeasured value. Cold cracks were not ob-served in the macrostructures of fillet

welds, as shown in Fig. 6B.Natural gas flowing within a pipeline

can increase the cooling rate of in-servicewelds. From the pipeline design standardof Korea Gas Corp., the safe flow rate of

natural gas within a pipeline is generallybelow 18 m/s, and the normal operatingflow rate is around 10 m/s. The effect ofgas flow rate on the maximum HAZ hard-ness was numerically investigated at thegas flow rate from 0 to 20 m/s.

As the gas flow rate increases, the heattransfer coefficient of the inner surface ofthe pipe increases, but the maximum HAZhardness increases only a little, as shownin Fig. 8. The maximum HAZ hardness atthe gas flow rate of 20 m/s is 257 HV, whichis only 6% higher than that with no gasflow. This result is in agreement with the

fact that cooling rates for thicknessesgreater than 0.5 in. (12.7 mm) are little in-fluenced by the fluid inside the pipe (Ref.7). From Fig. 8, it can be suggested thatsleeve repair welding of API 5L X65pipelines of 14.3-mm thickness with flow-ing gas can be performed without coldcracking occuring for the given weldingconditions.

Residual Stresses and Plastic

Strain Distributions

Residual stress and plastic strain are

produced by localized heating and coolingduring welding. The distributions of resid-ual axial and hoop stresses are shown inFig. 9. The axial residual stresses at theinner and outer surfaces of the pipe belowthe weld metal are tensile and compres-sive, respectively. During heating, the

welds are expanded, and the pipe is de-formed toward the outside of the pipe.During cooling, the pipe is bent towardthe inside by the faster cooling weld re-gion. Therefore, tensile axial residualstress is developed at the inner surface andcompressive stress at the outer surface.Hoop stress is tensile at both inner andouter surfaces of the pipe near welds.

Figures 10A and B show the evolutionof residual axial and residual hoop stressesalong the inner surface of the pipe during

welding, respectively. The intermediateresidual stress distributions of the weldpass in contact with the pipe and the finalresidual stress distributions are shown inFig. 10.

The axial stress at the inner surfacegradually increases with each pass se-quence as shown in Fig. 10A. The finalresidual axial stress is tensile within thezone ranging from +30 mm to 20 mm

and becomes compressive outside thezone. The maximum value of tensile resid-ual axial stress is found near the origin ofthe x-coordinate, that is, below the weldroot. The hoop stress distributions on theinner surface also show the tendency ofgradually increasing with pass sequence asshown in Fig. 10B. The tensile residualhoop stress shows a little decrease at thelast pass (pass 8). High tensile hoop stress

near the yield stress at room temperatureis developed at passes 3 and 6, which are

welds in contact with the pipe.The evolution of equivalent plastic

strain along the inner surface is shown inFig. 11. The plastic strain gradually in-creases from the second pass to the lastpass. The equivalent plastic strain is de-

veloped within the zone ranging from +40mm to 10 mm. The maximum plasticstrain is 0.44% at the given welding condi-tions. The equivalent plastic strain, whichcharacterizes permanent deformation,may be used as an indicator of cumulative

damage of the material during the weldingprocess. Because the pipe material hasmore than 30% elongation, which is muchhigher than the maximum plastic strain, ata temperature range from room tempera-ture to near 1000C, it is known the severeloss of material strength or excessive de-formation on the inner surface of the pipe

will not occur by the above plastic strain.From the calculated plastic strain and themaximum inside surface temperature, itcan be suggested the sleeve repair weldingof API 5L X65 pipelines of 14.3-mm thick-ness can be performed without melt-

through at the given welding conditions.

Effect of Internal Pressure on Residual

Stresses and Plastic Strain

The effect of internal pressure on theresidual stresses and plastic strain was in-

vestigated varying the magnitude of inter-nal pressure from 0 MPa to the maximumoperating pressure of 6.9 MPa. As shownin Fig. 12, the distributions of tensile resid-ual axial and hoop stresses on the innersurface are hardly influenced by the varia-tion of internal pressure. The effect of theincrease of hoop stress by internal pres-sure, which is called Barlow stress, is ob-served only in the compressive region.

The residual plastic strain decreases asinternal pressure increases, as shown inFig. 13. This is because the bending defor-mation toward the inside of the pipe dur-ing cooling is reduced by internal pres-sure. In fact, the condition in whichinternal pressure has a significant effecton in-service welding is that the tempera-ture of the inner surface of the pipe is highenough to cause melt-through.

The calculated maximum inside sur-face temperature is much lower than the

temperature of melt-through generation.From Fig. 13, it can be suggested thesleeve repair welding of API 5L X65pipelines of 14.3-mm thickness can be car-ried out without melt-through at the max-imum operating pressure of 6.9 MPa.

Allowable Heat Input for Single-Pass

Sleeve Fillet Welding

An analysis of single-pass fillet weldingwas carried out to assess the allowableheat input for sleeve fillet welding. TheSMAW process was used and the heatinput was varied from 0.5 to 5 kJ/mm. Thegas flow rate was varied from 0 to 20 m/sto investigate the effect of gas flow rate onthe maximum HAZ hardness.

Figure 14 shows the variations of max-imum inside surface temperature andmaximum HAZ hardness with the varia-tions of heat input and gas flow rate. Thecalculated maximum inside surface tem-peratures show lower values than the

melt-through prediction temperature of982C for all conditions, as shown in Fig.14A. The calculated maximum HAZhardness shows lower values than the lim-iting hardness of 350 HV for all condi-tions, as shown in Fig. 14B. From Fig. 14,it is shown melt-through and cold cracking

will not occur for a range of heat inputsand gas flow rates for single-pass sleeve fil-let welding.

Conclusions

An axisymmetric finite element model

was developed to simulate multipasssleeve fillet welding of in-service API 5LX65 pipelines of 14.3-mm thickness. Themodel was used to predict the tempera-ture distribution, maximum HAZ hard-ness, and distributions of residual stressesand plastic strain. The calculated geome-try of the fusion zone and HAZ was ingood agreement with the macrostructuresof the sleeve repair fillet welds. The pre-dicted maximum inside surface tempera-ture was much lower than the limitingmaximum temperature for preventingmelt-through. The calculated maximumHAZ hardness was in good agreement

with the measured value and much lowerthan the maximum allowable HAZ hard-ness for avoiding cold cracking. Coldcracking was not found in either the nu-merical simulation or the experiment.

Tensile axial residual stress is devel-oped at the inner surface and compressivestress at the outer surface. The hoop stressis tensile at both inner and outer surfacesof the pipe near welds.

Melt-through was not predicted fromthe calculated plastic strain distributionand the maximum inside surface tempera-ture. This was confirmed experimentally

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by applying internal pressure. The effectof internal pressure on the residualstresses and plastic strain was small. Theequivalent plastic strain showed little de-crease as internal pressure increased.

From the numerical simulation, it canbe suggested the sleeve repair welding of

API 5L X65 pipelines of 14.3-mm thick-ness can be carried out without melt-through at the maximum operating pres-

sure. Melt-through and cold cracking werenot predicted for a range of heat input andgas flow rates for single-pass sleeve fillet

welding.

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