rsc - in service welding on gas pipelines - part 1 - michael painter final report 01 jun 2000

8/13/2019 RSC - In Service Welding on Gas Pipelines - Part 1 - Michael Painter Final Report 01 Jun 2000

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CRC WS Pipeline Program

Project ReportMarch 2000

In-ServiceWelding of

Gas Pipelines


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CRCWS Project 96:34, In-Service Welding on Gas Pipelines: Final Project Report Contents page

In-Service Welding Of Gas PipelinesPipeline Program Report

September 1996- September 1999

Executive Summary

1. Introduction

2. Research Strategy & Outline3. Literature Review: Experimental Studies of In-Service Welding

4. Literature Review: Simulation of Fusion Welding

5. Appraisal of Battelle In-Service Welding Software

6. Welding on In-Service Pipelines: Industry Survey

7. An Experimental Appraisal of Hydrogen Controlled Electrodes

8. Numerical Modelling of In-Service Welding

9. Burn-through Prediction

10. Model Validation Using the Gladstone Flow Loop

11. A Proposed Method of Presenting Results for Industry Use

Mike Painter, CSIRO Manufacturing Science and TechnologyPrakash Sabapathy, The University of Adelaide


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In-ServiceWelding of Gas

Pipelines CRCWS Project 96:34 Final Report

Final Project ReportJune 2000

M.J.Painter, CSIRO Manufacturing Science and TechnologyP. Sabapathy, The University of Adelaide


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CRCWS Project 96:34, In-Service Welding on Gas Pipelines: Final Project Report 1

In-Service Welding ofGas Pipelines

September 1996 September 1999

A collaborative project of the Cooperative Research Centre for Materials Welding & Joining, underthe sponsorship and guidance of the Pipeline Program Management Committee.

Collaborative PartnersCSIRO Manufacturing Science & TechnologyThe University of AdelaideBHPEpic Energy

AGL Gas

The Pipeline Program Management CommitteeResearchersMike Painter, CSIRO Manufacturing Science & TechnologyM. A. Wahab, The University of AdelaideBing Feng, BHPPrakash Sabapathy, The University of Adelaide

Industrial mentorsPaul Grace, WTIA / ex AGL Gas SydneyHans Borek, Epic Energy

Project ObjectiveTo develop recommended weld procedures for the safe and effective in-service welding of

thin-wall, high-strength steel, high-pressure gas pipelines.

Executive Summary

BackgroundThe process of welding onto a live high-pressure pipe is frequently employed for the repair,modification or extension of gas pipelines. This in-service welding has significant economicadvantages for the gas transmission and gas distribution industries, since it avoids the costs ofdisrupting pipeline operation, and it maintains continuity of supply to customers. In-service welding

is an essential part of hot-tapping, a technique which allows the creation of a branch connectionto a live pipeline. In-service welding is also important for pipeline maintenance, such as theinstallation of sleeves around damaged sections of pipe. Direct deposition of welds onto livepipes has been suggested as a way of replacing wall thickness lost through corrosion or localdamage. If in-service welding is not possible then sections of a pipeline have to be sealed and de-gassed prior to welding, and then re-purged prior to reinstatement. These are costly, wasteful, andenvironmentally damaging actions, since there are large gas losses and methane is a green-housegas. TransCanada Pipelines Ltd. has estimated that relative to a cold connection a hot-tap avoidsgross revenue losses of approximately 1M$Canadian per hot- tap.

Welding Onto a Gas PipelineTwo factors make it difficult to weld onto a live pipeline. Firstly, the flowing gas creates a largeheat loss through the wall of the pipe, resulting in accelerated cooling of the weld. High carbonequivalent steels are sensitive to such rapid cooling rates, which increase hardness, and increasethe possibility of heat affected zone (HAZ) cracking. The second factor concerns the local


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heating, and the reduction in pipe-wall-strength during the welding process. If this reduction instrength is too great the pipe wall can burst under the pipes internal pressure. This hazardousevent is termed burn-through. Increasing the welding heat input can reduce fast cooling, but thispromotes weld penetration and increases the risk of burn-through. Suitable weld procedures mustensure the HAZ hardness is not high enough to cause cracking, whilst heat input and penetrationare not so high that the integrity of the pipe wall is jeopardised.

In Australia, there is a significant trend towards the use of high yield strength steels for pipelineconstruction. Future pipelines using X70 and X80 steels could have wall thickness as low as3mm. Unfortunately, in-service welding is made much more difficult with such thin-walled pipes.Thin pipe walls increase the risk of burn-through during welding, and are more easily cooled by theflowing gas. High strength steels can also be susceptible to the generation of excessive hardnessfor a given cooling rate. If the economic advantages of in-service welding are to be maintainedthen technology to support the safe and effective welding of thin-walled high-strength pipelinesmust be established.

In-Service Welding: Current TechnologyMuch of the technology associated with in-service welding was generated in the USA. This hasconsisted of two approaches:

1. An experimental method of measuring the cooling capacity of the pipeline.This data is used to establish the required heat input and weld cooling rates from laboratorypipe welds under the same simulated cooling conditions. This method, generally known as theEWI Test, was developed by Edison Welding Institute (EWI).

2. Numerical simulation of in-service welding.From 1980 to 1990 researchers at Battelle Memorial Institute developed a 2D finite differenceapproach to simulate sleeve and direct-branch in-service welds. For a given pipe geometryand a set of welding parameters the weld cooling times, t 8/5 (the time to cool from 800 to500 C) and the maximum inside wall temperature could be calculated. HAZ hardness wasestimated from this predicted cooling rate and the carbon equivalent of the steel. Hardnessbelow 350 HVN was considered to have a low potential for cracking. Pipes were consideredsafe from burn-through if the maximum inside wall temperature was below 980 C. Theselimits were determined from a comparison of experimental in collaboration with EWI. Thesetests were mainly on thick walled ( 6.35 mm), lower strength (X52) materials. Recommendedlimits for the valid use of the Battelle approach are pipe wall thickness within the range 3.2-9.6mm, for steel grades, up-to and including X52.

In-Service Welding: The Current Project A survey of the Australian pipeline industry established that there was a substantial conservatismin the in-service welding procedures currently used. Procedures generally involved a reduction ofpressure and gas flow before welding. The potential value to the pipeline industry of a researchprogram, which would allow the safe relaxation of such constraints, was estimated to be $2-4 Mover the period 1998-2002.

In recognition of: the differences between Australian conditions and those previously researched, namely,

reduced pipe wall thickness and use of higher strength steels, the lack of quantified information on burn-through limits for thin walled pipes, the importance of being able to carry out in-service welding on new , thin-walled pipelines, the indicated economic benefits that improved in-service welding would achieve,

the CRCMWJ Pipeline Program initiated a research project with the following aim:

Project ObjectiveTo develop recommended weld procedures for the safe and effective in-service welding of thin-

walled, high-strength steel, high-pressure gas pipelines.


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Project StructureThe important features of the project were:

1. An experimental study of circumferential manual metal arc (MMA) welding in the vertical-upand vertical-down position, with low hydrogen electrodes (E8018G and E7016, 2.5 mm & 3.2mm diameter). These tests used a water-cooled simulation of in-service welding and a rangeof typical Australian pipe grades (X42-X80).

2. The extensive development of numerical finite element simulations of in-service welding.These numerical simulations covered 2D and 3D models, and unlike previous research workaimed to develop numerical simulations of burn-through.

3. The transfer of technology to the pipeline industry.

Brief Project OutcomesAppraisal of Current KnowledgeIt has been suggested that the Battelle program over-estimates the t 8/5 cooling times for thin-walledpipe and under-estimates t 8/5 for thick walled pipe. This leads to the possibility of choosing non-

conservative heat input values for thin pipe ( 4.8 mm). The producers of the Battelle softwareconsidered it should be restricted to a wall thickness range of 3.2-12 mm. The calculation ofhardness relies on a relatively simple estimate of a carbon equivalent, which does not necessarilyapply to modern high-strength steel compositions, and should be restricted to grades up to X52.

It is generally accepted that in-service welds on pipes with wall thickness greater than 6.35 mmhave no significant risk of burn-through with good welding practice. The literature review confirmedthe sporadic nature of information relating to burn-through limits on pipes less than 6.35mm thick.

Recommended weld procedures have the following characteristics:

Severe restriction on welding conditions for pipes less than 5mm wall thickness. Somerecommendations that in-service welding should be restricted to pipes thicker than 4 mm.

Use of low hydrogen electrodes of small diameter 2-2.4 mm (restricting arc current), in thevertical-down welding position.

Limitations on heat input, typically 0.5 kJ/mm for 3.2 mm wall thickness, 1.4kJ/mm for5mm wall thickness.

Limits on pressure, typically < 6 MPa for thin pipes.

Careful control of arc current, heat input and welding practice.

Experimental studies have generally used water flow to simulate the cooling effect of the gas, butthere is a general recognition that water gives much higher quench rates than gas flow. It thereforegenerates conservative welding conditions for a required hardness. That is heat input determinedwith water-flow will give slower t 8/5 with gas flow. The corollary of this however, is that using water-flow simulations may give non-conservative heat inputs for burn-through.

Outcomes from the Experimental Study of Hydrogen Controlled Electrodes .The analysis of welding conditions and welds produced using pipe cooled with a water jacket togenerate a rapid quench, gave the following results;

It was possible to generate a weld on 7.8 mm thick X80 pipe with a maximum hardness of 325HVN with a weld t 8/5 cooling time of 3.8 seconds.

For the range of pipe grades examined (X42-X80), the Yurioka-1 carbon equivalent (CE)

relationship provided the best correlation between composition, hardness and t 8/5 cooling time.This correlation gave an absolute error of 5.7% for E5548-G (E8018-G) electrodes, and a


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3.4% error with E7016. Based on this relationship the X80 grade of pipe used in this workwould only require a t 8/5 of 0.8 seconds in order to achieve a hardness of >350 HVN.

A small beneficial tempering effect was measured from multi-pass [3 passes] welds. This wasmost noticeable for the X70 grade steel, which with a CE of 0.288 gave a hardness of 381HVN after a single root pass. After three passes, the maximum hardness was reduced to an

acceptable 321 HVN.

Only the two X60 steels with CE of 0.38 and 0.41 gave multi-pass hardness >350 HVN.

A minor difference in the incidence of weld defects was observed between the two electrodetypes tested. There was a greater incidence of HAZ cracking with E8018-G in keeping with itsreduced heat input. Some HAZ cracks were detected with X80 grade although the maximumhardness was 299-310 HVN.

There was no systematic variation in heat input as the weld progressed around the pipe.However, whilst total weld energy remained reasonably constant the natural variation inwelding speed, since this is a manually skilled process, caused variation in the heat input. Thiswas significant and amounted to approximately a 20% variation on a nominal value.

Penetration into the run-pipe was generally greater when using the E7016 electrode in thevertical-up position rather than with E8018-G in the vertical-down.

Penetration into the run-pipe slightly increased with increasing heat input although this effectwas largely swamped by a significant variability at a given heat input. At a nominal heat inputof about 1 kJ/mm this variability was approximately, 0.2-0.8 mm with E8018-G and 0.5-1.0mm with E7016.

These experiments also provided,

empirical data to relate the deposited weld bead volume, and weld bead shape to heat inputfor both E8018-G and E7016 electrodes, and

preliminary data on measured t 8/5 cooling times and measured weld penetrations in order tovalidate and refine numerical models of MMA in-service welds.

Development of Numerical SimulationsUtilising commercial finite element software, a wide range of 2D and 3D models of in-service weldshas been developed. The capacity to generate stable, accurate models of all in-service joint formshas been demonstrated.

To represent the heat loss due to the flowing gas the numerical models follow the Battelle modeland utilise the Sieder & Tate non-dimensional approach. This determines an effective heat transfer

coefficient at the pipe wall, based on the pipe diameter, gas pressure, and flow speed. Byincorporating empirical data describing weld bead volume and joint form the representation of thewelding heat input has been tailored for both vertical-up and vertical-down welding with low-hydrogen electrodes. For circumferential fillet welds, predictions of both t 8/5 cooling times and weldpenetrations have been made with acceptable accuracy.

Model development has concentrated on a 3D quasi-steady-state analysis of circumferential filletwelds. In practice, this is the most popular joint form. It also provides a numerical analysis whichallows an acceptable, short CPU time. Pre-processing software has been produced to efficientlyconstruct and mesh models of varied geometry. Post-processing software has been established tocalculate the distribution of t 8/5 cooling times throughout the weld zone. This can be further used tocalculate a distribution of hardness based on an empirical relationship with composition and t 8/5.Isotherms give an estimate of weld penetration and size and extent of the HAZ.


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Models have mainly considered single root pass welds but some multi-pass welds have beensimulated.

Various methods of assessing the risk of burn-through have been developed. These were basedon: the maximum temperature at the inside surface of the pipe, following Battelle, a thermo-elastic plastic stress analysis, using the thermal field in the pipe wall to calculate the reduction in wall strength.

Thermo-elastic-plastic models of circumferential welds have showed similar deformation patternsto those observed during burn-through, namely a localised bulge in the pipe wall near the weld.Failure can be specified as a bulge that exceeds a limiting height. Plots of internal pressureversus bulge height have shown an effective yield pressure, which can also be used as a failureindex. Only a limited numbers of these models were studied because they were verycomputationally demanding.

Estimating the reduction in pipe wall strength in the weld zone has created a novel alternativemethod of assessing burn-through risk. This method determines the reduction in material strengtharound the weld based on the predicted temperature field, and the known relationship betweenmaterial yield strength and temperature. This reduced strength is regarded as equivalent to a localreduction in the thickness of the pipe-wall at constant ambient temperature. Hence, thetemperature field around the in-service weld effectively converts to a cavity in the pipe wall. Anumber of alternate strategies can be considered. The limiting pressure for safe welding can bebased on the remaining wall thickness, or based on the effective reduction in cross-sectional area.The risk of burn-through is equivalent to the possibility of this cavity causing rupture at the currentoperating pressure. This assessment can also be easily carried out utilising the approachspecified for the evaluation of corrosion cavities in Australian Standard AS2885. This method hasproduced excellent results. Although limited by lack of data, comparison between predicted safewelding pressures and published values measured on 5 mm thick pipes has been good. Theapproach provided a way of assessing burn-through potential which is in agreement with reportedbehaviour. Longitudinal welds are more prone to burn-through than circumferential ones of thesame heat input. Pressure has a significant effect. The width or size of the weld is important aswell as penetration. It provides an efficient approach to in-service weld simulation since it does notrequire a stress analysis and uses only thermal predictions. Unlike Battelles maximum walltemperature approach, it is more realistic since it accounts for weld orientation and internal pipepressure.

Model ValidationThe above numerical simulations were validated by comparing predicted values with:

published result for t 8/5 cooling times measured on pipes of 4.8 mm wall thickness, measured HAZ and fusion zone geometries from a hot-tap coupon, data from welds carried out on an uncooled, empty pipe, measured values of t 8/5 and HAZ hardness obtained from a number of test welds carried outon a flow loop at Duke Energys Gladstone Gate facility.

This last extensive validation used simulated circumferential fillet welds on three materials, 4.8 mmand 5.2 mm thick X70, and 6.4 mm thick Ultrapipe X42, under a range of gas pressures and flows.The t 8/5 cooling times were measured, welds were metallographically sectioned and HAZ hardnessdetermined.

Predicted values of fusion zone depth, HAZ depth, t 8/5 cooling time, and HAZ hardness, comparedfavourably with measured values. This was an aggressive test of the models validity andaccuracy.

Transferring Project Knowledge to Industry . Although finite element software is readily available, developing and using thermal models tosimulate weld processes is a specialised activity. It would be difficult for industry to attempt this


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analysis without significant investment in software and the development of personnel withappropriate expertise.

Although access to the modeling capabilities developed within this project will remain, it was feltthat results could be put in a more accessible form. One possibility has been developed to aprototype stage. This consisted of using the finite element models to develop a database of

predicted values for a range of heat inputs, pipe wall thickness, and heat transfer conditions at thepipe wall.

Concentrating on the circumferential fillet weld, it has been established that the t 8/5 cooling time isalmost independent of pipe diameter provided the heat transfer coefficient at the pipe wall isconstant. That is, the heat transfer coefficient determines the weld cooling rate. This effectivelymeans that a single model can provide results for any combination of gas pressure, flow rate andpipe diameter, which gives a constant heat transfer coefficient.

There is a smooth variation in calculated t 8/5 cooling time, weld penetration, effective cavity sizesetc. as heat input, and pipe wall thickness varies. Hence, it is feasible to interpolate values from adata set with a reasonable degree of accuracy. Using this approach it is possible to develop avery fast computer program which produces estimates of HAZ hardness, and burn-through risk by

simply interpolating an established database.


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1. In-Service Welding of Gas Pipelines:

1.1 IntroductionMetal welding processes are used for the fabrication of structures ranging from, the large andcomplex to, the small and simple. The significance of welding may not be directly noticeable,but it has an important role in the manufacture of many tools, consumer objects, and in almostall industrial structures. Fusion welding is a significant engineering process because of theunparalleled advantages it has over other joining methods, and it is used extensively in theconstruction of Australia's gas pipeline network.

Unfortunately undesirable changes to material properties can occur during welding and thesehave the capacity for generating structural weakness or premature failure. Therefore a largeamount of research has been carried out on pipeline welding to avoid failures and suchdetrimental economic and environmental results.

A weld procedure defines all the weld process parameters which must be used in order toachieve a weld with the required service properties, e.g. type of welding process, (gas metalarc (GMA), manual metal arc (MMA)), electrode type, voltage, arc current range, weldingspeed etc. The present work concerns the development of MMA welding procedures used forthe maintenance and repair of gas pipelines. These in-service welding procedures are carriedout whilst those pipelines are 'live', or in continuous service. Due to the unique conditions in-service welds and welding procedures have particularly demanding requirements.

1.1.2 Industrial significance of In-service weldingIn-service welding may be used as part of a pipeline construction technique called "hot-tapping". This technique enables the connection of a branch pipe to a pipeline without stoppingor significantly disrupting the gas flow. The major advantage of this operation is that it avoids

the need to decommission the pipeline. That would be costly to the pipeline operator both interms of wasted gas and in un-serviced customers. McElligott et al (1) of TransCanada Ltdhave estimated that relative to using a cold connection, a single hot-tap can reduce grosslosses by $1 millon. In a simplified hot-tap, a pipe sleeve is initially welded to the live pipe, anda slide valve is attached to this fitting, see Figure 1.1(a) . The hot-tap drill is fitted to this valve,see Figure 1.1(b) . Next, the drill is used to cut a hole in the wall of the pipe, see Figure 1.1(c) .

As the drill is extracted it carries with it the cut-out or coupon, see Figure 1.1(d) . Finally thevalve is closed allowing the drill assembly to be removed, see Figure 1.1(e) . The success ofthe operation depends on the ability to weld the valve assembly or sleeve fitting onto the 'live'pipeline.

Because of the difficulties associated with this welding operation many hot-taps are currentlycarried out under conservative conditions which are achieved by reducing gas pressures and

flows. This can have significant impact on normal pipeline operation. Vented gas andcurtailment costs associated with such planned hot-taps in Australia from 1998-2002 havebeen estimated by Venton (2) to be 4 M$. As methane is a green-house gas, purging andventing of pipelines is potentially hazardous to the environment. TransCanada Pipelines Ltd.estimates that their use of hot-tapping will avoid an annual emission of 603 kTonne of carbondioxide equivalent in 1999 and 2000(3). That represents 18% of the total emissions reducedby their green house gas management program.

In-service welding can also be used as a technique for pipeline maintenance, to weldcircumferential sleeves at points of pipeline damage. It has also been suggested, Bruce(4),that weld deposits made directly on to a pipeline could be used to replace pipe wall thicknesslost by corrosion.

The success of such operations depends on safe and effective in-service welding procedures.


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Figure 1(b)Hot-tap drillattached to Slide

Figure 1(a) Sleeve welded in place andSlide valve attached.

valve

Figure 1(c) Drill used to cut holein pipe under pressure

Figure 1(d) Drill and cutcoupon removed and Slidevalve closed

Figure 1(e) Drill removed, branch readyfor connection

Figure 1.1 An illustration of the hot-tapping process taken from IPSCOs animation of hot-tapping on

http:/www.hottap.com.us


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1.2 In-service welding problemsThere are two significant problems associated with in-service welding. Firstly the high gas flowwithin the pipe (up to 15 m/sec) causes the weld to cool rapidly due to the convective transferof heat from the pipe-wall to the flowing gas. The result of increased weld cooling rates isgreater hardness levels within the weld and in the surrounding heat affected zone (HAZ). Withthe increased hardness of the microstructure in the HAZ there is an increased possibility ofhydrogen assisted cracking. The conditions needed for hydrogen assisted cracking include,hydrogen present to a sufficient degree, tensile stresses acting on the weld, and a susceptible,hard, HAZ microstructure.

The second problem concerns the risk of bursting the pipe wall during welding. Pressurisednatural gas (up to 15 MPa) imposes a significant stress on the pipe wall, and since the strengthof the pipe is decreased due to the localised heating during welding this can result in failure ofthe pipe wall. The result can vary from a small localised bulging of the pipe wall, up to burstingof the pipe. This is termed burn-through and occurs when the region around the weld pool hasinsufficient strength to withstand the internal gas pressure, see Figure 1.2 .

1.3 Australian Conditions

Figure 1.2Schematic illustration of burn-through, caused by localised

heating and internal gas pressure.

The recent development of high yield-strength, control-rolled, micro-alloyed steels has allowedthinner steel pipes to have the same load capacity as earlier, low strength, thicker pipes. The

Australian Pipeline Standard AS 2885, designates that the maximum pressure allowed forpipeline design is one giving a hoop stress equal to 72% of the yield strength. That is:

ywt

DP .72.0.2. =

whereP

is the internal pressure, D

is the pipe diameter, tw

is the pipe wall thickness, and y

is the minimum specified yield strength.

Hence for a given diameter pipe and gas pressure the tonnage of pipe required for a givendistance can be reduced as the materials yield strength is increased. Alternatively using highyield strength pipe permits the transmission of natural gas at higher pressures and flow rates.

The Australian pipeline industry recognises the economic advantages of using high strengthsteels. Unfortunately the use of thin walled, high strength steel pipelines increases thedifficulties associated with in-service welding. With the combination of enhanced gastransmission and diminished wall thickness the weld cooling rate for a given weld procedureincreases. Such high strength steels have a greater sensitivity to strength reduction duringwelding and together with the decreased wall thickness are more prone to bulging or burn-

through.


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1.4 SummaryWeld procedure development is particularly difficult for in-service welding. For safety andpracticality, experimental test welds can not simply be carried out on live pipelines. Hence,external means of establishing weld procedures have to be used. Traditionally two approacheshave been developed. Laboratory simulation of pipe flow conditions (Edison Welding Institute,1980-1990 (5,6)) or through the use of simple numerical calculations (Battelle MemorialInstitute, 1985(7,8)). Approximations in such approaches may lead to inaccuracies and excessconservatism in the choice of weld parameters. Such difficulties will be increased by the

Australian pipeline industrys use of higher strength, thin walled pipelines because the existingtechnology may not apply to such new materials .

1.5 Aim of Current ResearchThe aim of this current research is:

To develop recommended welding procedures for thesafe and effective in-service welding of thin-wall, high-strength steel,

high pressure, gas pipelines.

The technical challenge is to develop methods of establishing welding procedures whichproduce welds that are free from the risk of cracking, and do not risk bursting the pipe wallduring welding: and to confirm their application for the thin walled high strength materials thatwill be used in future pipeline construction.

References1. McElligott J. A., Delanty J., & Delanty B. Full Flow High-Pressure Hot Taps: The New

Technology and Why Its Indispensable to Industry, Paper Presented at International

Pipeline Conference pub. ASME v2, 1988, pp813-820.2. Venton P., Report Prepared for Pipeline Program of Cooperative Research Centre for

Materials Welding & Joining 1996.

3. TransCanada Pipeline, http://www.transcanada.com

4. Bruce W. A., Holdren R.L., Mohr W. C., Kiefner J.F. & Swatzel J.F., Repair of Pipelines byWeld Metal Deposition, Paper presented at PRCI 9 th Symposium on Pipeline Research,Houston Texas, September 1996.

5. Bruce W. A. & Threadgill P. L. Welding Onto In-Service Pipelines Welding Design &Fabrication Feb 1991, pp19-24.

6. Cola M.J. & Threadgill P.L., Final Report on Criteria for Hot Tap Welding, American Gas Association, Edison Welding Institute Project J7038, March 1988

7. Kiefner J.F & Fischer R. D. Models Aid Pipeline Repair Welding ProcedureOil & Gas Journal March 1988, pp41-47.

8. Fischer R.D., Kiefner J.F. & Whitacre G.R., User Manual for Model1 & Model 2 ComputerPrograms for the Predicting Critical Cooling Rates and Temperatures During Repair andHot Tap Welding on Pressurised Pipelines, Battelle Memorial Institute Report, June 1981.


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2. Research Strategy & Outline

2.1 BackgroundThe development of a weld procedure is essentially a trial-and-error process. For example ahypothetical structure may require a weld with predefined properties (e.g. penetration).Through experimental welding trials, involving welding on a replica or similar structure, manydifferent weld procedures are tested and the one with the closest properties to those desired ischosen, and replicated in the field. Often weld procedures are set by standards based on pastwelding trials and as a result may lack a scientific footing. The cost of establishing a weldprocedure can be large.

In relation to hot-tap welding, the majority of research has involved welding trials, either using aflow-loop (a diversion adjacent to an operating pipeline which allows welding trials to be carriedout on a test section of pipe without disruption to the existing pipeline) or using a laboratorysimulation. For in-service welds this process is made difficult since the heat loss due to gasflow cannot be reliably simulated in the laboratory. The details of the normal experimentalapproach will be discussed later, but it often consists of using water flow to generate highcooling rates. Water is a more efficient coolant than gas and hence there is a basic limitation induplicating the heat losses due to high gas flow. In addition, with this approach only part of theproblem is addressed. Such tests examine the required penetration, bead shape and HAZhardness but fail to consider burn-through, since the system is not pressurised. Thedetermination of burn-through limits requires further testing. If carried out at all, these testswould use pressurised, non-flowing gas, and again this represents an approximation to the realpipeline conditions. Because of the expense and time-consuming nature of this experimentalapproach there is a tendency to only examine a limited range of parameters. This can result ina lack of understanding of the sensitivity of the process to slight variation in weldingparameters, and it may not establish how close a particular weld is to the failure limit. Lack of

knowledge of process sensitivity can cause the welding trial to be both unreliable and unstable.The economic cost of performing in-service welding trials to this degree of accuracy would beprohibitive.

Because of the inherent difficulties with an experimental approach some research has usednumerical methods to calculate temperatures and cooling rates. Numerical simulations of in-service welding offer significant advantages. Firstly this approach avoids or minimises the useof time-consuming experimentation. The wide range of fittings and pipe geometries that areused in hot-tapping do not represent a significant problem to computer models. Similarly, thevariation of gas flows and pressures can be economically dealt with. Numerical models alsoallow the alternative of determining safe pressures and appropriate gas flows for a givenwelding process and this information can facilitate the management of hot-tapping and in-service welding procedures.

To-date only 2D models have been applied in the pipeline industry. These have only treatedburn-through control in a simple fashion by suggesting that it be signified by a limit to the pipewall temperature.

2.2 Project StrategyThis research programme has developed 3D numerical models of in-service welding using theFinite Element Method. 3D numerical models of the welding process, calculate the thermalfield during welding. This provides estimates of the cooling rate, and by linking this data toappropriate empirical equations relating hardness and cooling rate the HAZ hardness is alsopredicted. Burn-through has been directly examined by combining the thermal analysis with anelastic-plastic stress analysis. The numerical analysis of burn-through has been a novelfeature of this research programme. In particular this research has developed a new method ofassessing burn-through risk which takes into account the major factors affecting burn-through,namely, thermal field and its distribution, weld orientation, wall thickness and gas pressure.


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Different joint configurations are used by the pipeline industry which require individual analysis. Australian hot-tap fittings can be broadly classified under three types,

the full encirclement, circumferential sleeve fitting, see Figure 2.1(a), the direct-branch with reinforcement-saddle, see Figure 2.1(b), and the direct-branch-to-pipe weld, see Figure 2.1(c).

As the pipe thickness decreases the full encirclement sleeve provides the best structuralsupport to both the pipe and attachments. Therefore it is the most common joint configuration.

Although longitudinal welds are used to secure the sleeve around the pipe these do not directlycontact the pipeline and therefore are not critical. This research program has concentrated onnumerical simulations of circumferential sleeve welds and branch connections withreinforcement sleeves.

Figure 2.1Common in-service welding pipe configurations:(a) full encirclement fitting, longitudinal weld to join sleeves and a circumferential fillet to the

run pipe.(b) Reinforcing saddle around branch pipe.(c) Directly welding the branch pipe on to the run pipe.

Numerical models of fusion welding processes always include some empirical factors to ensurethat the resulting calculated values agree with those found in practice. This means that suchmodels can not be created without significant experimental input, and their accuracy must bevalidated.

The welding process commonly used for in-service welding in Australia is MMA welding usinghydrogen controlled electrodes. Unlike other welding processes, MMA welding requiresrelatively little equipment (power supply + stick electrode) and is the traditional process for in-field pipeline welding. An experimental assessment of the performance of two typical electrodetypes in common use has been carried out using a range of pipe material grades.

Although there is a large body of work on numerical modelling of fusion welding there is littlespecifically addressing MMA welding. This research has addressed that deficiency. Inparticular it has developed appropriate modelling strategies, for vertical-up or vertical-downMMA welding positions.

The current numerical simulations have been validated by comparing predicted values with:

published results for t 8/5 cooling times measured on pipes of 4.8 mm wall thickness, with measured HAZ and fusion zone geometries from a hot-tap coupon, with data from welds carried out on an uncooled, empty pipe,

(a) (b) (c)


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measured values of t 8/5 and HAZ hardness obtained from a number of test welds carriedout on a flow loop at Duke Energies Gladstone Gate facility.

- The results of thermal analysis have been directly compared with micrographs of test

welds to ensure the correct calculation of weld penetration and HAZ geometry.

- The measured cooling rates of test welds have been compared with predicted values.

- HAZ hardness has been measured for a range of materials and test welds and hasallowed further comparisons between predicted hardness and measured values.

This last extensive validation was an aggressive test of the models validity and accuracy underoperational conditions on a live pipeline.

Although finite element software is readily available, developing and using thermal models tosimulate weld processes is a specialised activity. It would be difficult for industry to adopt thisapproach without significant investment in software and the development of personnel with

appropriate expertise.

Although access to the modelling capabilities developed within this project will remain, it wasfelt that results could be put in a more accessible form. One possibility has been developed toa prototype stage. This consists of using the finite element models to develop a database ofpredicted values for a range of heat inputs, pipe wall thickness, and heat transfer conditions atthe pipe wall. Using this approach it is possible to develop a very fast program which producesestimates of HAZ hardness, and burn-through risk by simply interpolating an establisheddatabase.

It is anticipated that through the development of improved numerical simulations of in-servicewelding, and the establishment of their accuracy and scientific credibility, more efficient weldprocedure development will result. Validated numerical models will also allow a safecombination of welding procedure, gas pressures and gas flows to be determined for a givenpipe geometry, and this in itself will facilitate more efficient management and control of in-service procedures.

2.3 OutlineSection 3 will discuss the experimental work that has examined in-service welding. The areasof post weld hardness and the possibility of pipe wall failure during in-service welding will be itsfoci. Section 4 will introduce the concepts related to the numerical simulation of weldingprocesses. Sections 5 & 8 will concentrate on the computer simulation of in-service welding,and on work related to the development of a numerical approach to the prediction of safewelding procedures.


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3. Literature Review:Experimental Studies of In-Service Welding

3.1 IntroductionThis section reviews the past research on in-service welding. In particular it will identify the need foradditional research and the reasons for the current research activities.

The research work into in-service welding can be grouped into three areas, namely: experimentally determined weld procedures, experimental studies of burn-through and investigations of weld repair, and, the development and application of numerical simulation.

This list does not imply that there is a large amount of information available, for published researchwork and data related to in-service welding is rare. Most work has been carried out by the AmericanGas Associations Pipeline Research Committee, in sponsored work at the Edison Welding Institute(EWI), by Cola & Threadgill (1) and Bruce & Threadgill (2,3), and at the Battelle Memorial Institute

(BMI), by Kiefner & Fischer (4) and Fischer et al (5). The results of that work form the basis of thecommon methods used to establish appropriate in-service welding procedures.

EWI (2) developed a method for experimentally establishing in-service welding procedures, followingsimilar work by British Gas. At BMI, Fischer, Kiefner & Whitacre (5) developed a numerical approach,and produced commercial software to predict weld cooling rates and the possibility of the pipe bursting(burn-through) during in-service welding. In collaboration with the EWI, their numerical approach wasvalidated and connections between the two approaches established (6). Other literature is sporadic,generally relating to limited experimental studies of particular hot-tap welding conditions. Phelps et al(7), Wade (8,9,10) and Bruce et al (11,12), have produced data on burn-through limits.

3.2 Experimental Studies of In-Service Welding

- Factors controlling crackingThe older generation of steel pipes has compositions (high carbon equivalent) which are susceptible tohydrogen assisted cracking. Much experimental work on in-service welding was therefore concernedwith determining which of the process variables controlled post weld hardness, since this is commonlythe property used to assess crack susceptibility. Figure 3.1 broadly summarises the relationshipsbetween post weld hardness and the weld process variables.

Baileys work (13), reported by Graville & Read (14), identified that with manual metal arc (MMA)welding using low hydrogen or rutile electrodes, cracking did not occur below a critical HAZ hardnessof 350 HVN. Graville & Read considered that in Baileys work even the low-hydrogen weld exhibitedsignificant hydrogen implying some conservatism in the critical hardness specified by Bailey.

For in-service welding, as in any fusion weld, the cooling rate and the chemical composition of thesteel are the main factors influencing post weld hardness. The weld heat input clearly influences thecooling rate, but for in-service welding the complicating factor is the high heat loss generated by thegas flowing in the pipe. This accelerated cooling, which in-turn depends on the pressure, the flow rateof the gas and the pipe geometry, has a dominating influence on the process.

Experiments on operational pipelines and under simulated conditions have clearly shown that gas flowstrongly increases the cooling rate of the weld, particularly in thin walled pipes (1,3), see Figure 3.2 forexample. This observation is of particular relevance to Australian conditions, where there is asignificant trend towards the use of thin-walled pipes. Phelps et al(7) reported that the HAZ of in-service welds could become brittle, giving rise to cracking immediately after welding, or at a later stage


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through hydrogen embrittlement. They quoted the case of a 200mm diameter 8mm thick, steel pipewith a carbon equivalent (CE) of 0.48. Welds on this pipe gave a HAZ hardness of 415 VHN for a gasflow of 0.518 scm/h and 285 VHN with no gas flow. Whilst their main concern was with the pipes HAZhardness they also indicated that the composition and cooling rate within the fitting should not beignored.

Cola et al(1) reported on the cracking propensity of welds made with basic and cellulosic electrodes.They considered that with cellulosic electrodes a limiting t 8/5 (time to cool from 800 C to 500 C) couldbe established independent of pipe composition. This conclusion was based on data generated fromsimulated in-service welds using an E6010 electrode at 1kJ/mm heat input. For a range of steels witha CE of 0.3-0.5 cracking was only found for t 8/5 cooling times below 5 seconds. When using basicelectrodes they recognised that the pipe composition was an important factor, and determined thatcracking was only a concern for an HAZ hardness >400 VHN. They also reported that, provided basicelectrodes and a sound low-hydrogen welding practice were used, the risk of cracking was notsignificant for steel having a CE < 0.5. The significance of this relates to the age of the pipe sincemodern steel compositions generally give a CE < 0.4-0.45. The use of basic electrodes is not justrelated to their cracking propensity. Whilst they clearly reduce the risk of hydrogen embrittlement, asPhelps et al(7) showed, they also generate significantly lower penetration for a given heat input andhence reduce the risk of burn-through (see Section 9). Welding vertically-down or welding with theelectrode DCEN were also reported (7) to give a lower penetration than welding vertically-up, or usingelectrode DCEP.

Boran(15) found that post-weld hardness was influenced by the electrode polarity of the MMA weldingprocess. The polarity influenced the apportionment of heat between the electrode and the workpiece,.DCEN gives the greater fraction of heat in the weld region, which reduces the weld cooling rate andgives the least post weld hardness for a given heat input.

Preheating the joint before welding is an obvious, traditional, way of controlling and reducing thecooling rate of during welding. For example, DeHertogh & Illeghems (16) preheated a 323.5 mmdiameter 4.4 mm thick, X60 pipe with a gas flow of 25,000 m 3/hr at 4.8 MPa. They found that thehardness of a weld decreased considerably from 367 to 317 when using an 80 C preheat. Usinginductive heating they examined preheat levels from 50-200 C and reported that minimum hardnesswas achieved at a 100 C preheat, although no explanation of the minimum value was proposed.

Cassie et al (17) listed the recommended features of a satisfactory in-service welding procedure as;use of basic low hydrogen electrodes, a preheat of 100 C for material with a CE < 0.4 and 150 C formaterial with CE>0.4, and the use of a stringer bead technique.

Preheating has an economic penalty of course, but for in-service welding there are other significantdifficulties. These relate to the extremely high heat loss generated by the flowing gas. Under suchconditions achieving a consistent preheat is difficult. A number of methods were tried, including directgas flame heating, electrical resistance and inductive heating. Phelps et al (7) reported that none ofthese was entirely satisfactory. The method they recommended was direct flame heating using ahand-held propane torch. This was used to heat the region ahead of the weld to a maximumtemperature of 250 C. This preheating step was followed by welding for a short time, until thetemperature of the region fell below the desired preheat level. This cyclic process was then repeated.Cooling rates were extremely high so it followed that welding-runs would be short. Since consistencyof heat input relies on manual skill of the welder, such an intermittent process can only contribute tothe variability in welding speed and in heat input.

In-service welds are multi-pass, so the sequence of welding can be controlled to minimise HAZhardness. Using a stringer bead technique, Figure 3.3(c), in which the current weld tempers thehardening created by the previous one, is recommended by Cassie (17) and Bruce & Threadgill (2).Bruce (18) describes the sequence recommended by British Gas which uses a buttering layer and atemperbead sequence, also shown in Figure 3.3(b) . Rietjens (19) made reference to a desirable weld


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preparation for split circumferential sleeves, as shown in Figure 3.3(a) and identified that buttering andtemperbead sequences were useful in minimising crack susceptibility. He also advocated using weldmetal with low yield strength, in order to relieve residual stress.

Cassie et al (17) examined the application of post-weld heating as a means of reducing the hardnessat the weld toe. They found that using a gas tungsten arc was effective but they did not consider it aviable field technique, other methods were not reported in detail but were considered ineffective.

Variability in the manual welding process was identified as a concern by a number of researchers.Cola et al (1) referred to the inherent variability in the manual process. Cassie (17) identified variationsin hardness due to different welders using different welding speeds and hence a varied heat input.Bruce et al (3) recommended the use of controlled deposition rates in order to minimise suchvariations. There are some reports of a systematic variation in weld properties around circumferentialwelds. DeHertogh et al(16) indicated an increase in the depth of HAZ as the weld progressed from thetop to the bottom-dead-centre of a circumferential MMA weld. They considered that this was aconsequence of the general heating of the pipe as welding progressed. For 4.8 mm thick pipe andvertical-down MMAW Phelps et al (7) also found that penetration was significantly greater at the 6oclock position (1.3-1.4 mm) than at the 3 oclock position (1.0 mm).

3.3 Experimental Studies of In-service Welding- Factors Controlling on Burn-throughConsidering the safety implications, there have been few attempts to determine the conditionsnecessary to avoid pipe-wall failure during in-service welding. Experimental work has generally useda small number of test welds under widely varied experimental conditions, so conclusions tend to begeneral directions rather than quantified limits. The relevant factors are shown in Figure 3.4. Clearly,the risk of burn-through is related to the loss of pipe wall strength in the weld zone, and its inability toresist local stress. The reduction in wall strength depends on the elevated temperature around theweld, and on the depth of weld penetration relative to the original wall thickness. Observations ofburn-through generally show significant local plastic distortion of the pipe wall, and a fracture along theweld pool axis (11,12).

3.3.1 The influence of pipe wall thickness on burn-through In-service welds on thin pipe walls have a high risk of burn-through. Weld penetration is largelyinfluenced by the welding heat input, so the same heat input on a thinner walled pipe causes a greaterrelative reduction in the wall strength. Research efforts have generally sought to determine the lowerlimit of pipe wall thickness that could be safely welded. For longitudinal welds on X60 using 3.2 mmand 4 mm diameter electrodes, Wade(8) found that successful welds could be made on 6 mm thickpipe at normal operational pressures with a heat input of up to 1.8 kJ/mm. For 5 mm thick pipe heconsidered that welding could still take place but recommended a considerable restriction ofpressure (below 3 MPa) and heat input (less than 1.4 kJ/mm). He established that burn-through wasprobable for pipe with 3 mm wall thickness, even at low pressure. Specifying a minimum pipe-wallthickness below which in-service welding should not take place has been a convenient way ofdesignating a safe procedure. It is generally accepted that burn-through risk is minimal for pipeswhich have walls thicker than 6.35 mm(18). Many operational standards have restricted in-servicewelding by specifying a minimum wall thickness, with restrictions in the range 4-5 mm beingcommon(20).

In 1983 Hicks(21) listed a recommended approach for avoiding burn-through. This indicated that atthe maximum allowable operating pressure (MAOP), {equivalent to the pressure giving a hoop stressof 72% of the yield strength} burn-through was possible with a 4 mm wall thickness. Hisrecommendations included limiting in-service welding for longitudinal welds to pipes greater than4.8mm thick, and for circumferential welds to greater than 4mm wall thickness. He also recommendedthat pressure should be restricted to that specified by the ASME Gas Piping Standard,


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Where p is pressure, y is minimum yield strength, D is pipe diameter, t is pipe wall thickness, and,c is an assumed thickness reduction, to account for reduced wall strength equivalent to 2.38 mm.

( ) D

ct p y

=

.44.1

3.3.2 Influence of heat input on burn-through

As the welding heat input is increased the weld penetration, and size of the heated region increases,with a consequent increase in the possibility of burn-through. Cassie(17) investigated the processparameters controlling weld penetration using various out-of-position welds on 6, 9 and 12 mm thickplate. From these results, he determined that welds in a vertical-down position using basic electrodesprovided the lowest penetration, and therefore would be the most appropriate for in-service welding.Cassie(17) used simulated in-service welds on pressurised cylinders of 450 mm diameter X52 steelwith wall thickness of 3.2, 4.8 and 6.4 mm. He defined a safe limit by specifying the maximum arccurrent allowable for a given pipe thickness, see Figure 3.5 . Safe welding currents could bedetermined for all pipes with wall thickness above 3.2 mm. However, the final recommendation wasthat welding should not take place on pipes of less than 4 mm thickness at an internal pressuregreater than 7 MPa.

Bruce et al(11) also referred to a restriction on welding current for safe in-service welding on pipe walls 3 mm thick. They identified that at the same heat input, welding with a smaller diameter electrode(equivalent to a reduced current) reduced the burn-through risk. The limits suggested are showndiagrammatically in Figure 3.6 . This shows an interesting difference in behaviour between 3.2 mmand 4 mm thick pipes. For a 3.2 mm pipe wall the division between safe and unsafe welding wasstrongly dependent on electrode diameter (arc current), with 2 mm diameter (50 A) electrodes safe,and 2.4 mm diameter (80 A) electrode borderline. With a 4 mm thick pipe-wall the conditions are moredependent on heat input. This apparent difference is unexplained. Bruce et al(11, 12), determinedrecommendations for safe in-service weld repair as: A maximum internal pressure of 6.7 MPa during welding on a minimum remaining wall thickness

of 3.2 mm, Electrode type to be a hydrogen controlled, E7018, with electrode size to be restricted to 2.4 mm

diameter, A heat input of 0.51 kJ/mm for the first weld runs.

3.3.3 Influence of welding technique on burn-throughManual metal arc welding (MMAW) is a process requiring considerable skill and hence the control ofpenetration and heat input is welder dependent. Wade(8) recommended that electrode types shouldbe specified for their capacity to run smoothly and produce uniform penetration at all points in theweld. He also recommended that welds should have equal leg length and that weaving should beminimised and a planned sequence of short welds should be used to minimise heat build up ahead ofthe welding arc. Phelps et al(7) observed that penetration was greater at the 6 o-clock position, duringvertically-down welding.

3.3.4 Influence of preheat on burn-throughCassie(17) considered that preheat did not influence burn-through limits. Wade(9) found a smalleffect, noting an increased tendency for local bulging with increased preheat. However, for typicalpreheat levels of around 100 C the effect was not large.

3.3.5 The influence of pipe diameter on burn-through Wade(10) considered that the diameter was not important. He argued that internal pressure acting onthe weakened pipe wall was the main driving factor, not the hoop stress. Bout & Gretskii(22) however


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considered burn-through limits in terms of hoop stress, implying that the pipe diameter has aninfluence.

3.3.6 The influence of welding direction on burn-through Longitudinal welds are more prone to burn-through than circumferential ones. Bout & Gretski(22)quote limiting heat inputs for welds in both directions indicating that welds in a circumferential direction

can tolerate higher heat input before burn-through, see Table 3.1 . This observation appears to pointtowards the importance of the applied hoop stress on burn-through, and it is difficult to rationalise thiswith Wades(10) conclusion that the pipe diameter is not significant.

3.3.7 Influence of internal pressure on burn-throughInternal pressure is recognised as a factor influencing burn-through, although Bruce et al(11)considered it secondary to heat input. Wade(8) carried out longitudinal fillet welds on 250mmdiameter cylinders pressurised with nitrogen. For safety, a mechanised welding system was used inthe down-hand position. He observed that significant plastic deformation occurred within the weldzone prior to bursting. Therefore he measured the local pipe wall deformation at the weld site anddetermined the onset of burn-through to be a critical local bulge height of 1 mm. He produced graphsrelating pressure, bulge height and heat input, and determined a critical pressure/heat input line forburn-through (see Figure 3.7) . His work lead to a diagram, as shown in Figure 3.8 , which specified

acceptable working zones (pressure, versus heat input) for in-service welding on pipes of different wallthicknesses.

3.3.8 The influence of pipe grade on burn-throughSteel undergoes a dramatic reduction in strength as its temperature is increased, such that attemperatures over 800 C its yield strength is 4-10% of its room temperature value. The strength attemperature of a higher grade steel such as X70 is not significantly higher than that of a low strengthgrade. So although a high room temperature strength allows pipes to be thinner, this increasedstrength is not present in the weld pool region during welding. This non-proportional reduction instrength is an additional factor that increases the burn-through risk with thin X70 or X80 pipe materials.

Because many of the above observations have been made on statically pressurised pipes, and thegas flow within a normal operational pipe is an aggressive coolant, it is thought that many of theseburn-through conditions are conservative. It is interesting that Hicks(21) recommended that aminimum gas flow of 0.4 m/sec be maintained during welding since this acts to reduce pipe walltemperatures. It is recognised that the unsafe region for MMA welding is on pressurised pipes withwall thickness around 4 mm, however no clearly defined, quantitative limits exist.

3.3.9 A Burn-through Avoidance StrategyPipe failures at defects such as corrosion cavities or notches are referred to as pressure controlledfailure. Bursting strength is dependent on defect dimensions and yield strength. Bout & Gretskii(22)were the first to use this approach to estimate the load carrying capacity of the pipe during in-servicewelding. They considered that above 700 C the metals strength was effectively zero, therefore, the700 C isotherm around the weld pool could represent a surface defect, of a given depth and length.

They determined that the following formulae could be used to estimate the maximum hoop stress thatcould be sustained during welding.

=

M r t

r t 85.0

/85.02.0

Where,


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t is the pipe wall thickness Rt

L M

4.01 +=

L is the axial length of the weld pool perpendicular to the hoop stressr is the maximum depth of the defect

R is outside radius of the pipe

Using these formulae Bout & Gretskii(22) then estimated the critical length of heated zone ( Lcritical )before the pipe would not be able to operate at its maximum load carrying capacity.

1935.91.0/85.0.212.1

=

r t

Rt Lcritical

This was a potentially valuable approach to assessing the impact of a weld on the load carryingcapacity of the pipe but Bout & Gretskii(22) did not continue with this or integrate this method with anumerical thermal model. Instead, they determined that the permissible sizes of heated zone werevery restrictive and suggested a way of overcoming this.

They considered that by increasing the support of the pipe wall during welding the freedom for thepipe wall to plastically deform would be reduced, and burn-through would be prevented. Theyachieved this by surrounding the weld with reinforcing rings or bands as shown in Figure 3.9 . Thegap between the rings could be adjusted to give varied support to pipes of different wall thickness orpressure. The relationship between allowable pressure, effective thermal penetration and gap width(a o) between the reinforcing rings was given as follows.

( ) 227002.0 /4 oC at ar t p = For a 3 mm wall thickness, 320 mm diameter pipe with an internal pressure of 4 MPa, burn-throughoccurred at a heat input of 0.475 kJ/mm with a normal sleeve. With a constraining band the criticalheat input was 0.915 kJ/mm.

This represents an interesting novel approach. The only concerning feature is that the practical fieldapplication quoted are for 9mm thick pipes. Also no consideration was given to the increased coolingrate that would be generated within this joint configuration.

3.4 The Determination of Safe In-Service Welding Procedures

The major thrust of research work on in-service welding has been to generate sound welding practicesand to determine processes whereby weld procedures can be easily and efficiently established for agiven hot-tap. Because the gas flow, gas pressure and pipe geometry strongly influence weldingoutcomes this latter aspect is particularly relevant, since a weld procedure must be established foreach individual hot-tap. As Bruce et al(3) indicates, many codes require that a qualified weldingprocedure must be determined for a given in-service weld. AS 2885(23) requires that, welds shall bemade following a qualified welding procedure which takes into account pressure and cooling effectsfrom the flow of fluid within the pipe and simulates site condition.

Although such techniques seem essential, not all pipeline companies use them. It has beenreported(3) that of eleven member companies of the American Pipeline Research Committee, onlythree used specific procedure development. The others either weld on decommissioned pipelines, orweld using reduced pressures, flows and preheat to overcome high cooling rates.


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An important feature of procedure development concerns selecting the minimum suitable heat input.This should be selected to achieve a weld with HAZ hardness below the level likely to cause cracking,but not be so high that the pipe wall may burn-through during welding. The most comprehensiveapproach is to carry out a conventional weld procedure development using test welds on a special flowloop or pipe by-pass. This provides a controllable segment of pipe under identical conditions to thosein-service. However such facilities are not common, and the normal approach is to physically simulate

e thermal characteristics of the operational pipeline by using water, air or oil flow through a section of

the pipeline. This value is often referred tos the EWI cooling time, T . The experiment is repeated on other spots upstream from the first, and

ringelding as compared with that generated during EWI testing. Tests have been reported in which

7018 and E8018 electrodes.he cooling capacity was measured using the EWI process and t values were determined from

the extrapolations have

enerally been chosen conservatively. With this approach, the required heat input to achieve a

thtest pipe.

Following on from early work by British Gas, EWI developed an experimental method of determiningthe cooling capacity of a working pipeline(2). This is achieved by a simple procedure suitable for usein the field. An oxy-propane torch is used to heat a 50mm diameter area on the exposed pipe wall.Heating is stopped when the temperature reaches approximately 325 C. Using a stopwatch and athermometer, the temperature drop is monitored as the pipe wall cools. The time taken to coolbetween 250 and 100 C is taken as the cooling capacity of a ewia final cooling time is arrived at by averaging six readings.

The philosophy behind the test is to measure a parameter which reflects the cooling capacity of theworking pipeline, and to do this in a safe, experimentally simple and practical way. An experimentaltest bed can then be set up by duplicating the measured T ewi for this test pipe. To use this method forestablishing a physical simulation it is not necessary to adhere to the EWI procedure. Simplyduplicating the experimental procedure, heat source, spot size, temperature range etc, in the field andin the workshop should be sufficient. The general philosophy is satisfied provided the heat transferbehaviour of the gas or the test fluid is not significantly altered by the wall temperature achieved duwwater, water-mist sprays, compressed air or oils are used to achieve the desired cooling capacity.

EWIs work also generated much experimental data relating the weld heat input and cooling rate for in-service welds. The data were measured from flow loop tests and from test welds carried out undersimulated conditions. The t 8/5 weld cooling times were measured for heat inputs between 1 2 kJ/mmusing pipe materials ranging in thickness from 4.8-8.0 mm and E6010, ET 8/5temperatures measured by thermocouples harpooned into the weld pool.

For a given pipe thickness and gas flow an approximately linear relationship between t 8/5 and heatinput were often found. In addition, EWI established that there was a proportional relationship linkingthe cooling capacity T ewi , and the t 8/5 value obtained for a given heat input. Experimental data wereused to form graphs such as Figure 3.10 for a 4.8 mm thick pipe. To use this graph to derive asuitable heat input requires a measurement of the cooling capacity of the operational pipeline. Takethe value of 20 seconds for example. This represents a particular line between t 8/5 and weld heat inputon Figure 3.10 . For a given CE, the IIW relationship between hardness and t 8/5 can be used toestimate the minimum t 8/5 cooling time required to give a hardness level of 350 HVN, (6 seconds inthis example). The point on the relevant cooling capacity line at this t 8/5 value gives the requiredminimum heat input (1.1 kJ/mm). The empirical data used to establish these lines are also shown inFigure 3.10 , so it can be seen that the data points are relatively sparse and

gdesired hardness can be estimated directly from the measured heat capacity.

To make use of the existing EWI data linking t 8/5 , and EWI cooling time it is necessary to have moreconcern about using the same procedure as EWI for determining the cooling capacity. The EWI test isnot rigorously defined, and there is some concern about possible variability due to the changes in thesize of the heated region and the heating rate used (see Appendix 1). EWI carried out some numericalanalysis of the test process and concluded that the cooling capacity was not overly sensitive to thechosen spot size or heating rate(1). There are other difficulties here however, since the relationshipsbetween heat input and t 8/5 are empirical and may vary with electrode type, welding position, joint type


25/38


etc. Similarly, the EWI approach makes no reference to fluid type, joint position, weld preparation,electrode type or preheat. It simply relies on the measured heat capacity and the developed empiricaldata set. It was pointed out by Cola & Threadgill(1) that the correlation between heat sink capacity andthe cooling rate of the weld had only been established over a limited range of field conditions.

ertainly the data are sparse and requires considerable extrapolation in some instances. Data are not

r estimated, thus being sure to obtainardness below the chosen limit. The critical level of hardness is chosen at 350 HVN, which is also

e any distinction between the saddle or sleeve configuration, and does not consider theffect of multi-pass welds or tempering effects. It also provides no direct information about burn-

gas flow could be achieved by using a low flow of water,


26/38


The simulated flow is usually unpressurised, so these tests do not rigorously establish burn-throughconditions. If burn-through limits are required, these are commonly determined with a staticpressurised test using compressed inert gas. A test vessel is fabricated using a sample length of pipe.This should include an internal cylinder to reduce the volume of compressed gas and increase safety.The static gas has a lower cooling capacity than the gas flow in an operational pipeline so burn-through limits established on such test are also conservative. That is, for a given heat input walltemperature

s during the test are likely to be higher than those achieved on the working pipeline.ence, under test conditions burn-through is likely to take place at a lower heat input than it would do

ckness greater than 6 mm. However as pipeall thickness is reduced it is clear that experience and information about in-service welding becomes

rate. If the economicdvantages of in-service welding are to be maintained then technology to support the safe and

ll temperature. Theegree of conservatism in the established limits is unknown. Clear information on the heat input and

pipes is desirable.

enerallyquired. Again because such tests rarely use flowing gas the heat input limits will be conservative.

in walled pipes. On thin pipes the heatput to avoid burn-through is very restricted, and it is necessary to recognise the danger of the

put in a manual process.

possibility of using an automated GMA

Hin practice.

3.5 Summary Of Section 3 and Required Research

With care and systematic planning it is clear that systems and technology are in place to allow in-service welding to be carried out on pipes with wall thiwless and the process becomes inherently more difficult.

3.5.1 Increased use of thin walled pipes burn-through limitsIn Australia, there is a significant trend towards the use of high yield-strength steels for pipelineconstruction. Future pipelines using X70 and X80 steels could have wall thickness as low as 3 mm.The reduced wall thickness is more sensitive to strength loss and increases the risk of burn-throughduring welding. Thin walls are more easily cooled by the flowing gas and high strength steels can besusceptible to the generation of excessive hardness for a given coolingaeffective welding of thin-walled high-strength pipelines must be established.

This inevitably raises a question about the wall thickness at which in-service welding becomes unsafe.The consensus is that the current lower limit is approximately 4-5 mm, although welds have beenmade on pipes of 3.2 mm thickness. Current burn-through limits largely rely on experimental data thathave been determined under conservative conditions, or use a restricted pipe wadpressure limits to avoid burn-through particularly on thin walled

3.5.2 Experimental weld procedure development Weld procedure development is particularly difficult for in-service welding. Clearly, for safety andpracticality, experimental test welds can not be carried out on live pipelines. Hence other means ofestablishing weld procedures have to be used. Simulation of accelerated cooling generally usesystems that give greater cooling rates during the test than will be experienced on the live pipe. Thisgenerates a weld procedure with a conservative heat input with respect to the generation of excesshardness. However this may be non-conservative with respect to burn-through. In such cases, saywith pipe < 6.4 mm thick, then additional experiments on pressurised segments of pipe are greSuch conservative outcomes may unnecessarily restrict in-service welding to thicker pipes.

Approximations in such systems may lead to inaccuracies and excess conservatism in the choice ofweld parameters. Such difficulties will be enhanced with thinpotential loss of control over heat in

3.5.3 Heat Input VariationManual processes are inherently variable in welding speed, and hence in heat input. This is a clearproblem for a welding procedure that depends on meeting a critical working range of heat input.Interestingly the possibility of addressing this through the introduction of automated systems wasraised by Cassie(17) some 20 years ago. He suggested the


27/38


welding system to accurately control heat input. In addition, he speculated on the possibility or

burn-through conditions . Secondly, HAZ hardnessay not be the sole criteria for the assessment of hydrogen assisted cold cracking and t 8/5 may not be

cooling.

thickness, that our conservative

The technical challenge is to address in-service welding on the thin walled high strength materials that

ed for modern pipe compositions,

es, and

addressing improved control of heat input .

. Cola M. J. & Threadgill P. L., Final Report on Criteria for Hot Tap Welding, American Gas

2. ruce W. A. & Threadgill P. L., Welding Onto In-Service Pipelines,

. Bruce W. A. & Threadgill P. L., Effect of Procedure Qualification Variables for Welding Onto In-

4. iefner J.F & Fischer R. D., Models Aid Pipeline Repair Welding Procedure,

el 2 ComputerPrograms for the Predicting Critical Cooling Rates and Temperatures During Repair and Hot Tap

implified Weld Cooling

n VA 22209, 1993, pp31.1-31.22.

incorporating pre-heat or post heating within such a system.

3.5.4 Hardness limits to assess potential for crackingWith appropriate validation computer models can reliably predict weld cooling behaviour, howeverhardness is conventionally used to assess the potential for cracking. This raises two possible gaps in

our knowledge. Firstly there is some inaccuracy possible in the determination of hardness from therelationships between composition, and cooling rate. There are many such relationships, which mayhave varied relevance for certain ranges of steel. The current solution to this is to adopt theconservative approach, but here again for thin-walled pipe this may lead to the selection of higher heatinputs than necessary, and run into conflict withmthe most relevant factor to assess weld

3.6 Technical ChallengesTechniques for in-service welding on pipes with wall thickness greater than 6 mm are well established.Provided a careful and systematic approach is taken, such welds can be produced with safety. Toapply similar techniques to pipes with wall thickness below 5 mm, or for weld repair where remaining

wall thickness is low is a technical challenge. For it is below thisapproach, and lack of knowledge about suitable HAZ hardness levels, burn-through limits, and theimpact of a poorly controlled manual process, becomes prohibitive.

will be used in future pipeline construction. The major aspects will be:

the determination of quantifiable, validated burn-through limits for thin walled pipes,

improve confidence in the hardness limits that are us

quantifying the role of temperbead techniqu

3.7 References

1 Association, Edison Welding Institute Project J7038, March 1988.

BWelding Design & Fabrication Feb 1991, pp19-24.

3service Pipelines, American Gas Association Report J7141, July 1994.

K

Oil & Gas Journal March 1988, pp41-47.

5. Fischer R. D., Kiefner J. F. & Whitacre G. R., User Manual for Model1 & Mod

Welding on Pressurised Pipelines, Battelle Memorial Institute Report, June 1981.

6. Bruce W. A., Bubenik T. A., Fischer R. D. & Kiefner J.F., Development of SRate Models For In-Service Gas Pipelines, Line Pipe Research Proceedings, 8 th Symposium,September 1993, Paper 31, pub Arlingto


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. Phelps B., Cassie B. A., & Evans N. H., Welding Onto Live Natural Gas Pipelines, Metal

. Wade J. B., Hot Tapping of Pipelines, Australian Welding Research Association Research Report,

. Wade J.B., Effect of Preheat on Hot Tapping Procedures, Australian Welding Research

amage onPerformance and Hot Tapping Techniques, paper presented at Australian Welding Research

ren R.L., Mohr W. C., Kiefner J.F. & Swatzel J.F., Repair of Pipelines by WeldMetal Deposition, Paper presented at PRCI 9 th Symposium on Pipeline Research, Houston Texas,

2. Bruce W. A., Holdren R.L. & Mohr W. C., Repair of Pipelines by Direct Deposition of Weld Metal

3. Bailey N., Welding Procedures for Low Alloy Steels, The Welding Institute Cambridge England

4. Graville B. A. & Read J. A., Optimization of Fillet Weld Sizes, Welding Journal Research1s-167s.

15.The Hot-Tapping of Sub Sea Pipelines

6. DeHertogh J. & Illeghems H., Welding Natural Gas Filled Pipelines, Metal Construction & British

7. Cassie B. A., The Welding of Hot Tap Connections to High Pressure Gas Pipelines, paper

8. Bruce W. A., Welding Onto In-Service Pipelines: A Review, paper presented at Pipeline Welding

9. Rietjens I. P., Safely Weld and Repair In-Service Pipelines, Pipeline Industry, December 1986,

0. Considerations of Welding Methods Adopted on Pipelines During Operation, IIW Document XI-E-

1. Hicks D. J. Guideline for Welding on Pressurised pipe, Pipeline & Gas Journal,

2. Bout V.S. & Gretskii Yu.Ya., Arc Welding Application on Active Pipelines, Pipeline Technology,

85 1987 Clause 7.13.11.

4. Holman J.P., 1976, Heat Transfer, 7 th edition pub. McGraw-Hill.

7Construction, August 1976, pp350-354.

8Snowy Mountains Corporation 1978.

9 Association Research Report, Snowy Mountains Corporation, September 1978.

10. Wade J.B., Description of Experimental Results on the Effects of Pipeline D

Associations conference Pipeline Welding in 80s, Melbourne March 1981, Paper 4a.

11. Bruce W. A., Hold

September 1996.

1Further Studies, Final report Edison Welding Institute, EWI Project J7283, November 1996.

11970.

1Supplement pp16

Boran J

Welding Review, vol6, no 4 Nov 1987, pp283-284

1Welding Journal, March 1972, pp224-227.

1presented Pipeline Industries Guild J. W. Jones Memorial Lecture, October 1974.

198, International Symposium on Pipeline Welding, May 1998.

1pp26-29.

2477-87, 1987.

2

March 1983, pp17-192

Volume 1, R. Denys Ed. R. Denys, pub Elsevier Science BV. pp550-558.

23. Australian Standard Pipelines Gas & Liquid Petroleum AS 28 2


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HAZ Cracking

hydrogen

electrode type

hardness > 350

susceptible microstructure

material composition

weldcooling

rate

gas flow

gas pressuregas thermal properties

Pipe geometry

weld heat input

Figur e 3.1Factors influencing the possibility of hydrogen assisted cracking during in-service welding.

0

1

2

3

4

5

6

7

8

0 0.5 1 1.5 2 2.5 3 3.5 4

Gas Flow (mmscmd)

T 8 5 C C o o

l i n g

T i m e

( s e c s

)

4.8 mm6.4 mm9.3 mm

15.1 mm

Figur e 3.2For a constant heat input of 0.9 kJ/mm this graph shows the reduction of t 8/5 cooling time due to increased

gas flow in pipes of different wall thickness(3).


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10

8 9

Figur e 3.3Recommended weld bead deposition sequences in order to make most benefit of tempering,

(a) general view of recommended joint configuration(b) buttering layers and a temperbead at weld toe

(c) stringer bead arrangement.

7 6 5 1 2 3

1 2 4 3 5

6

2t t

45 deg .

(a) Less than 1 mm

(b) Buttering layers & temperbead

(c) Stringer bead arrangement

4


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Arc current

Arc voltage

Welding speed

Electrode polarity

Electrode type

Electrode diameter

Welder technique/direction

Pipe wall thickness

Preheat temperature

Gas flow rate

Gas temperature

Pipe diameter

Gas pressure

Yield strength at temperature

Weld orientation

Local wall support

Heat Input

Weld Penetration

Local Pipe Wall Strength

Pipe Wall Temperature

BURN THROUGHRISK

Pipe WallCooling

Local

Applied Stress

Figur e 3.4Causal factors involved in burn-through during in-service welding


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Figur e 3.5Burn through limits expressed as limits to the allowable welding current, for both 3.2 mm and 4 mm

diameter electrodes from Cassie(17). Additional points from Bruce et al (11).

Recommended limits for avoiding burn throu h during repair welding from Bruce et al(11)Figur e 3.6

g

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

1.1

50 60 70 80 90 100 110

Welding Current (Amps) / Electrode Diameter

M a x

i m u m

H e a

t I n p u t

( k J / m m

)

3.2 mm wall thickness

4 mm wall thickness

2.0 mm 2.4 mm 3.2 mm

Burn-through Limits

100

120

140

160

180

200

220

3 4 5 6Pipe Wall Thickness (mm)

M a x

i m u m

A l l o w a b

l e W e l

d i n g

C u r r e n t

( A m p s

)

7

3.2 mm diameter electrode

4 mm diameter electrode

safe < 0.87 kJ/mm

burn-through

Not allowed without significantpressure restriction


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Table 3.1 Conditions for burn-through from Bout & Gretskii(22)

Pipe r(mm)

Wall ess(mm)

Internal Pressure(MPa)

Welding Direction Crit forBu h

(kJ/mm)

Diamete Thickn ical Heat Inputrn-throug

320 3.0 4.0 Longitudinal 0.37

320 3.0 4.0 C

320 3.0 3.0 Circumferential 0.51

ircumferential 0.48

320 3.0 3.0 Longitudinal 0.48

Burn-through limits established by Wa a 300 mm diameter, X60 steel pipe of5 mm wall thickness.

0

2

4

6

8

10

12

14

0.5 0.7 0.9 1.1 1.3 1.5 1.7 1.9

Welding Heat Input (kJ/mm)

P i p e

I n t e r n a l P r e s s u r e

( M P a )

Wade results, bulge height < 1.0 mm

Wade results, bulge height > 1.0 mm

Wade results, burst

Figur e 3.7de (8) for welding ont

rsc - in service welding on gas pipelines - part 1 - michael painter final report 01 jun 2000

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