the prediction of burn-through during in-service welding of gas pipelines

The prediction of burn-through during in-service welding of gas pipelines

P.N. Sabapathya, M.A. Wahaba,* , M.J. Painterb

aMechanical Engineering Department, The University of Adelaide, Adelaide, 5005 AustraliabCSIRO Manufacturing Science and Technology, Adelaide, 5012 Australia

Received 21 April 2000; revised 22 August 2000; accepted 8 September 2000

Abstract

Numerical methods have a useful role in the assessment of welding conditions for the safe in-service welding of high-pressure gaspipelines. These have been used for the prediction of thermal cycles leading to an estimate of heat-affected-zone (HAZ) hardness andpossible cracking. In earlier work burn-through limits have only been considered indirectly, i.e. based on the maximum temperature of thepipe wall.

In this paper, two significant research aspects of the numerical simulation of in-service welding have been addressed as follows:

1. A new mathematical description of a heat-source has been formed to represent the common in-service welding process, i.e. vertical-up andvertical-down manual metal arc (MMA) welding with hydrogen controlled electrodes. Empirical relationships between welding processinputs, weld bead size and weld bead shape define the weldment geometry and control the heat source co-ordinates. Finite element modelsusing this heat-source have given good correlation with experimental and field welds. With this approach adequate agreement betweenpredicted weld penetration, weld cooling times and HAZ hardness, has been made.

2. The prediction of burn-through has been achieved using a full thermo-elastic–plastic model, but this leads to lengthy calculations. Here, anew approximate method of predicting burn-through has been developed and shown to give industrially useful results. This is based ontranslating the temperature field into an effective cavity in the pipe-wall thickness, and using this information to calculate a safe workingpressure during in-service welding.

q 2001 Elsevier Science Ltd. All rights reserved.

Keywords: Numerical modelling; In-service welding; Pipe-wall failure; Hot-tapping; Burn-through

1. Introduction

Welding onto a gas pipeline in active operation is a tech-nique that is frequently employed in the repair, modificationor extension of gas pipelines. This ‘in-service welding’ hassignificant economic advantages for the gas transmission orgas distribution industry since it avoids the costs of disrupt-ing pipeline operation and secures continuity of supply tothe customer. If in-service welding were not possible,sections of the pipeline would have to be sealed and de-gassed before any welding operation, and then purgedprior to reinstatement. These are costly, wasteful and alsoenvironmentally damaging actions since methane is a‘greenhouse gas’. In-service welding is an essential part ofhot-tapping, a technique which allows the establishment of abranch connection to a live pipeline. It is also important forpipeline maintenance such as the installation of sleeves

around damaged sections. Direct deposition of weld-metalonto an active pipe has also been suggested as a way ofreplacing wall thickness lost through corrosion or localdamage.

Two factors make welding onto in-service pipelines diffi-cult. Firstly, the flowing gas creates a large heat loss throughthe pipe-wall, resulting in accelerated cooling of the weld.Higher carbon equivalent steels are sensitive to such rapidcooling rates, which increases hardness and the possibilityof HAZ cracking. The second problem concerns localisedheating and loss of material strength during the weldingprocess, since the pipe-wall may burst under internal pres-sure if this reduction in strength is too great. Fast coolingcan be compensated for by increased heat input, but thispromotes weld penetration and possible burn-through.Suitable weld procedures must be a balance between thosethat ensure the HAZ hardness is not high enough to cause acracking problem, whilst heat input and penetration are notso high that the integrity of the pipe-wall is compromised.The pipeline industry clearly requires systems to establish

International Journal of Pressure Vessels and Piping 77 (2000) 669–677

0308-0161/00/$ - see front matterq 2001 Elsevier Science Ltd. All rights reserved.PII: S0308-0161(00)00056-9

www.elsevier.com/locate/ijpvp

* Corresponding author. Tel.:161-8-8303-5439; fax:161-8-8303-4367.E-mail address:[email protected] (M.A. Wahab).

such safe welding procedures. Alternatively, safe limits forpipeline flow and pressure have to be established for whichwelding is possible. In such cases, pipeline operation maybe disrupted slightly, but the integrity of the pipelines is notaffected.

In Australia, there is a significant trend towards usinghigh yield strength steels for pipeline construction. Asspecified by design codes (e.g. AS2885) the minimumallowable pipe-wall thickness is inversely proportional tothe material yield strength (s y) so using high-strengthsteel has a direct economic benefit since it allows the useof thinner pipe. Using X80�sy � 551 MPa� in place of X60�sy � 413 MPa�; e.g. allows a reduction of wall thicknessof 25% and leads directly to a reduction in steel volume for agiven pipeline.

Existing technology and current experience has beenmainly concerned with relatively thick walled (.6 mmthick) low-strength steel pipes; X70 and X80 pipelineswill have wall thicknesses of 3–5 mm. Unfortunately, in-service welding is made more difficult with thin, high-strength pipes. The reduced wall thickness is more sensitiveto strength loss and increases the possibility of burn-throughduring welding. Thin walls are more easily cooled by theflowing gas and high-strength steels can be more susceptibleto the generation of excessive hardness for a given coolingrate. If the economic advantages of in-service welding are tobe maintained then procedures for the safe and effectivewelding of such thin walled, high strength pipelines mustbe established.

The useful role of numerical simulation of in-servicewelding has already been demonstrated by work at EdisonWelding Institute (EWI) and Battelle Memorial Institute(BMI) [1,2,3]. During the 1970s, these researchers devel-oped a 2D finite difference approach to simulate sleeve anddirect-branch in-service welds. These models calculated aneffective heat transfer coefficient at the inner pipe wall basedon pressure and flow conditions within the pipe. Then, forthe given pipe geometry, and a set of welding parameters theweld cooling time, e.g.T8/5 (time to cool from 800–5008C),and the maximum inside wall temperatures were calculated.The hardness was then estimated from empirical equationsconnecting, the steel’s composition, the cooling rate, and theresulting hardness. Hardness below 350HVN10 was consid-ered to have a low cracking potential. In collaboration withEWI these models were extensively tested against experi-mental data mostly derived from tests with relatively thick(6.35 mm) low strength pipes.

In 1992, Goldak et al. [4] applied a more general 3D finiteelement method to calculate the thermal fields for circum-ferential fillet and direct-branch welds. They found that theassumed weld bead size and shape of the fillet weld had asignificant influence on the calculated penetration andtemperature profile around the weld pool.

It is evident that numerical methods have a useful role inthe assessment of welding conditions for the safe in-servicewelding of high-pressure gas pipelines. However a direct

prediction of the conditions likely to cause burn-throughhas not yet been made. Bruce [5] has shown that burn-through is strongly dependent on welding heat input andon the extent of penetration into the pipe wall. These factorsbeing more significant than the internal pipe pressure. In thework at BMI [1,2] burn-through limits were not directlycalculated but were based on applying a limit to the maxi-mum inner wall temperature of 9808C

In this work we have attempted to increase the accuracyof thermal prediction by using empirical relationshipsbetween welding process inputs, weld bead size and shapeto define the weldment geometry and control the heat sourceco-ordinates. In addition, a new approach to the problem ofburn-through has been developed based on the calculationof an effective cavity in the pipe wall thickness. Thisapproach has been verified by comparison with thethermo-elastic–plastic models directly simulating pipebursting during welding.

2. Thermal field prediction of hot-tap welding

To model hot-tap welding we need to consider the accu-racy with which the welding process is represented. Goldak[4] has already pointed out some of the inaccuracies that canresult if the welding heat input is not accurately determinedor the heat source is not accurately positioned. In addition,the numerical model must account for the convective heatloss into the pressurised gas flow.

2.1. A numerical representation of the common in-servicewelding Process

The welding process that is predominantly used by indus-try for hot-tap welding is MMA welding using low-hydro-gen electrodes and a vertical-down welding technique.Representing the MMA welding process in a heat transfercalculation is challenging. Due to the non-linear, complexnature of welding, a strategy commonly employed has beento represent the welding process as an arbitrarily definedmathematical distribution of heat-flux or ‘a heat-source.’The convective heat transfer mechanism due to the swirlingmolten weld pool is commonly approximated. Manyresearchers [4,6,7] compensate for the weld pool convectiveheat transfer by using an artificially high conductivity forthe metal above melting. Compensation for these effects isalso built into the formulation of the heat-source. The heat-source description is therefore unique to each weldingprocess, and an accurate or suitable manual metal arc‘heat source’ has not been described in the literature.

An example of a heat-source distribution is the ‘DoubleEllipsoidal Heat Source’ (DEHS) developed by Goldak etal. [7]; which defines the heat fluxQ (kJ/mm3) for a pointwithin the heat source volume. This heat source is describedby six parameters:h the arc efficiency, (such thatQ� hVI;where I and V are the arc current and voltage); ellipsoidwidth a; ellipsoid depthb; front lengthcf; rear lengthcb.

P.N. Sabapathy et al. / International Journal of Pressure Vessels and Piping 77 (2000) 669–677670

In addition, an apportionment of heat to the front or rearsection of the source is made by factorsr f andrb, whererf 1rb � 2: The double ellipsoidal heat source gives the heatflux at a point (x,y,z) within the ellipsoid in front of thearc centre by the following equation:

q�x; y; z� � hVI·rfp

��pp 6

��3p

abcf

× exp 23xcf

� �2

23ya

� �2

23zb

� �2" #

�1�

An identical equation except with the substitution ofcb andrb applies to points within the rear section of the ellipsoid.

The DEHS has been often used to approximate commonnon-autogenous welding processes. However these are

usually simple welds carried out in the ‘down hand flatposition’, i.e. welding horizontally in a straight line on ahorizontal flat plate, with the electrode perpendicular tothe plate. A number of modifications to the formulation ofthe DEHS need to be made to accurately model the out-of-position, low hydrogen, MMA welding process.

The MMA welding process used for in-service welding isapplied using a ‘weave’ technique. This technique and thecharacteristics of the low-hydrogen electrode often gener-ates shallow penetrations, which suggests a heat distributionwhich is flatter and more evenly distributed than Gaussian.This can be achieved by changing the exponential terms inthe DEHS. The new Eq. (2) is:

q�x; y; z� � Qf exp 23uxucf

� �n1

23uyua

� �n2

23uzub

� �n3� �

�2�

The constant23 is chosen in the same manner as Goldak etal. [7], by requiring that the value of heat flux at the bound-ary is 0.05Qmax. In this case the boundary is not ellipsoidalbut corresponds such that given by:

xc

� �n1

1ya

� �n2

1zb

� �n3� 1

Now the value ofQf has to be determined by numericalintegration as follows in Eq. (3):

hVI � Qf

Xvolume-of-source

× exp 23uxucf

� �n1

23uyua

� �n2

23uzub

� �n3� �

�3�

A visual comparison of heat-flux between a Gaussian form�n1 � n2 � 2� and a non-Gaussian�n12;n2 � 2� can be seenin Fig. 1a and b. Increasing the power of exponentn2 above2, spreads the distribution in they-direction, and simulates‘weave’ across the weld. Appropriate values have beendetermined as�n1 � n2 � 2� and n� 10; giving a ‘disk-like’ source with low penetration. A comparison betweenthe heat flux distribution from the DEHS and that from themodified distribution used in this work is also shown inFig. 2.

In-service welding often requires attaching either a full-encirclement (circumferential) sleeve or a branch pipedirectly onto the carrier pipe. Fitting an encirclement sleeveinvolves welding a fillet circumferentially between thesleeve and pipe. Direct branch-on-pipe in-service weldsrequire welding along the ‘saddle-like’ intersection createdbetween the carrier and branch-pipe fitting. The positionalvariation of the end of the electrode against time describesthe motion of the heat-source. The transient effects of weld-ing are then modelled using a moving heat-source. Themoving heat-source has a local co-ordinate system whilethe stationary finite element mesh has a general co-ordinate

P.N. Sabapathy et al. / International Journal of Pressure Vessels and Piping 77 (2000) 669–677 671

Fig. 1. Heat flux distributions at the planez� 0 from: (a) Goldak’s ‘doubleellipsoidal heat source’; and (b) and from (b) a modification to representmanual metal arc welding.

system. Applying co-ordinate transformation due to themotion of the heat-source, as shown in Fig. 3a and b, theheat-flux can be calculated at any point along the weld path.

The modelling strategy employed considers the transienteffects during welding, but any possible effects due to arc-initiation are not modelled. This approach is based on theassumption that the unknown effects of arc-initiation onlyinfluence the first few seconds of welding and any welddeposited thereafter is not affected.

An alternative strategy that may be used for modelling afull circumferential fillet weld is to consider a quasi-steady-state approximation. This considers the heat source to be

stationary, with the joint moving past it at a constant speed.The steady state thermal field is calculated. Where it can beapplied, this quasi-steady-state modelling approach hassome advantages in computational speed compared to thetransient models.

2.2. Numerical modelling of flowing pressurised gas

The natural gas flowing within a pipeline increases thecooling rate of in-service welds. The heat transfer mechan-ism present within in-service welding is assumed to beforced convection. Calculations of the Reynolds numberfor natural gas pipe flow found typically in Australian hot-tap operations gaveRe. 100; 000: The combination oflarge �Re. 100; 000� Reynolds number along with theoften remote-locations of hot-taps (i.e. hundreds or thou-sands of diameters long) firmly place the flow regime as‘fully developed turbulent flow’.

The calculation of a heat transfer coefficienthc at theinside wall of the pipe was determined by a non-dimen-sional approach using Seider and Tate’s empirical equation[8] for heat transfer under fully developed turbulent flow, as


Fig. 2. Comparison of a typical heat distribution from the DEHS, with thedistribution developed in this work to represent manual metal arc welding.This figure shows a typical contour map of the heat flux distribution at thecentre section, for (a) the DEHS; (b) and for the modified distributiondeveloped to represent manual metal arc welding with low hydrogenelectrodes.

Fig. 3. Diagrammatic representation of the orientation (welding angle andtrailing angle) used to position the heat source through co-ordinate trans-formations at all points throughout the welding run in the manual metal arcprocess.

follows:

Nud � 0:027Re0:8d Pr1=3 m

mw

� �0:14

�4�

where Nud is the Nusselt Number� hc=rk; Red theReynolds Number based on the diameter;Pr the PrandtlNumber;mb, mw the viscosities at bulk and wall tempera-

tures, respectively;hc the effective heat transfer coefficient;andr andk are the density and thermal conductivity of thegas.

2.3. Transient thermal modelling assumptions

The 3D finite element model was created in theNisa [9]software, using a mesh of 8 noded-linear, brick elements.


Fig. 4. Thermal field calculated from a 3D transient heat transfer model for a single pass circumferential fillet weld. Temperature contours set to the meltingpoint (15008C) and heat-affected-zone (HAZ) limit (7208C) to show the extent of melting and the size of the HAZ.

The deposition of the weld bead was not modelled based onthe assumption that the welding speed is fast enough tominimise the level of thermal diffusion in front of theweld. Temperature dependent properties were consideredand the convective boundary conditions on the outsidesurface of the pipe usedhc � 0:12× 1024 �W=m2K� forconvection to ambient air. Radiation was assumed to beinsignificant and was ignored.

2.4. Comparison between predicted and experimentalvalues

EWI has measured theT8/5 cooling time of numeroussimulated hot-tap circumferential fillet welds on a test pipe-line by harpooning a thermocouple into the molten weldpool. This data allowed a test of the modelling approachby comparing measured and predicted values.

Both 3D transient and 3D quasi-steady-state models wereconstructed using the welding parameters (arc current,voltage, welding speed and electrode diameter) defined byEWI’s field welding trials [10].

Numerical models predict the thermal field as it variesduring welding, as illustrated in Fig. 4. From this dataT8/5

cooling times were extracted and compared with themeasured values. The results show reasonable good agree-ment with the minimumT8/5 values having an average of12% error compared to EWI’s field welding trials, as indi-cated in Fig. 5. A 12% error is considered acceptable parti-cularly in view of the variation in heat-input, which canarise with a MMA welding processes.

3. Numerical predictions of burn-through

3.1. Thermo-elastic–plastic model

The modelling approach to burn-through prediction canbe demonstrated by considering an extreme case of weldingdirectly onto the wall of a pressurised pipe. This approachuses a full thermo-elastic–plastic analysis, but some approx-imations are implemented in order to reduce the size of theproblem and achieve solution in a reasonable computationaltime.

The pipe wall would be at its weakest for a fully devel-oped (steady state) temperature field, and burn-throughwould be most critical at that time. The high thermal fieldaround the weld is relatively small. Hence the reduction ofstrength of the pipe wall is restricted to a small area, there-fore only a segment of the pipe has been considered.

Fig. 6 shows the steady state temperature field for a1.0 kJ/mm weld. The stress calculation determines thevariation of geometry during the incremental applicationof internal pressure; and Fig. 7 illustrates this by showingthe radial deflection of the weld pool region. Fig. 7 showsthat as the internal pressure increases a point of non-linear-ity occurs between the radial deformation in the pipe walland the pressure. The gas pressure at this effective local‘yield point’ can then be used as a limit to signify burn-through.

3.2. Alternative model for the prediction of burn-through

As long as the above models could be used to predictburn-through, the approach was computationally expensiveand so, a faster approximate method has been developed.Burn-through clearly depends on the localised high


Fig. 5. An indication of the accuracy of numerical models by comparingpredictedT8/5 (T85C) cooling times with EWI field data (the closer the datapoints are to the median line the better the agreement between predicted andexperimental data).

Fig. 6. (a) Temperature field; and (b) calculated deflections from a thermo-elastic– plastic analysis for bead on pipe.

temperature, which creates a local reduction of pipe-wallstrength in the region of the weld pool. This methodattempts to calculate the effective strength of the wall inthat region. For example, a temperature distribution fromthe inner pipe-wall (at radiusr i) to the weld bead surface (at

radiusrouter) can be seen in Fig. 8. The local yield strength atthese elevated temperatures can then be calculated along asection line A–A. The load to create this yield stress alongA–A, will then be, as follows in Eq. (5):

PA–A �Xrouter

r i

syrdr �5�

wheresyris the yield stress at a given radial location and

temperature.The thickness of the material at ambient temperature that

would support this load can then be calculated as in Eq. (6):

teff � PA–A

sy0

�

Xrouter

r i

syrdr

sy0

�6�

wheresy0is the yield stress at ambient temperature.

The reduction in strength in the weld region is now repre-sented by an effective reduction in thickness of the pipewall, at its original strength. By performing this calculationover a localised area of the pipe wall, a region of reducedthickness is generated, which can be viewed as an effectivecavity. Such a region is shown in Fig. 9.

The maximum allowable pressure of a pipe with a cavity inits wall can be calculated by using a procedure specified in the


Fig. 7. Calculated radial deflection versus pressure for a simulated in-service weld of 0.75 kJ/mm directly onto a 3 mm thick, X70 pipe. This predicts aneffective yield pressure of 7 MPa.

Fig. 8. Section through the calculated temperature field giving the tempera-ture distribution from the pipe wall to the weld bead surface. From this thetemperatures along line A–A can be determined.

Australian Pipe Standard AS2885 [11]. This approach allowsthe calculation of the maximum allowable pressure for a speci-fic cavity defined as a maximum depth and a projected lengthperpendicular to the hoop stress. An alternative, and closelyrelated approach is to calculate the burst pressure using a set ofequations identified as the B31G criterion, as described byKeifner et al. [12]. This approach relates the bursting pressureto the yield strength of the material, and the lost cross sectionalarea of the cavity. There are some restrictions on the use ofthese equations and variations exist, these are described byKiefner et al. [12].

Now using these equations and the cavity described by thecalculated effective thickness (teff), the approximate burstingpressure during in-service welding can be calculated.

3.3. Comparison between predicted failure andexperimental data on burn-through

To assess the suitability of this approach, model predic-tions have been compared against experimental data gener-ated by Wade [13]. Wade carried out welds on a cylinderpressurised by nitrogen, and determined the heat inputslikely to cause burn-through at particular pressures. Burn-through was considered to occur when local bulging in theweld region approached 1 mm high.

Models have been created for the welding conditionsused by Wade and the technique described above used topredict critical bursting pressure for burn-through. As indi-cated by the comparison between Wade’s data and model


Fig. 9. The calculated ‘effective cavity’ in the pipe wall representing the loss of strength during in-service welding. From this cavity the maximum depth andwidth perpendicular to the hoop stress can be determined and used to calculate the pipe’s bursting pressure.

predictions, see Fig. 10, the cavity method gives a reason-ably accurate estimate of burn-through.

4. Discussion and conclusions

This work has shown that it is possible to reliably simu-late the process of in-service welding. Three dimensionalfinite element models of circumferential fillet welds havebeen developed. These models have been used to predictweld zone geometries andT8/5 cooling times. Comparisonsbetween experimental and predicted values ofT8/5 haveshown that it is possible to calculate the thermal fields andweld cooling times to a useful accuracy. Predictions werewithin 12% of measured values. Given that the heat inputfrom a manual welding process can vary significantly, thisaccuracy of 12% is within an acceptable limit.

Thermo-elastic–plastic models were used to examine thepossibility of the pipe bursting during in-service welding.These models showed that failure occurs as a localisedradial bulging of the weld pool region. After a critical pres-sure the rate of radial deformation rapidly increases, this iseffectively a yield pressure, and can be used as an indicatorof the onset of burn-through.

An alternative and new approach to the prediction ofburn-through was also developed. This is based on the

concept of representing the softened weld region as an‘effective cavity’ in the pipe wall. This approach onlyrequires a thermal field calculation and therefore results ina considerable reduction in computer calculation time bycomparison to the thermo-elastic–plastic solution. It istherefore more industrially relevant. In addition, this tech-nique accounts for internal pressure within the pipe whereasthe previous methods only used the maximum temperatureof the pipe wall and did not consider internal pressure.

Predicted burn-through limits have shown good agree-ment with available experimental data.

Acknowledgements

The financial support of the Australian Co-operativeCentre for Welded Structures (CRC-WS) and the in-kindsupport of the CSIRO Manufacturing Science and Technol-ogy (CSIRO-CMST) and The University of Adelaide isgratefully acknowledged.

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Fig. 10. Predicted failure conditions for burn-through compared withexperimental data from Wade [13].

the prediction of burn-through during in-service welding of gas pipelines

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