Cranfield University

Mark Keeler

The steps needed to avoid failure during in-service welding on a live Gas line

School of Applied Sciences

MSc Design project

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Cranfield University

School of Applied Sciences

MSc Design project

Academic year 2009-2011

Mark Keeler

The steps needed to avoid failure during in-service welding on a live Gas line

Supervisor: Dr Paul Colegrove

Academic Year 2009 to 2011

This thesis is submitted in partial fulfilment of the requirements for the degree of Welding Engineering

(NB. This section can be removed if the award of the degree is based solely on examination of the thesis)

© Cranfield University, 2010. All rights reserved. No part of this publication may be reproduced without the written permission of the copyright holder.

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ABSTRACT

In-Service welding is carried out all over the world where modifications need to be made on existing oil or gas pipelines. This has many inherent risks that are not present during normal workshop, or field, welding. The fact that you have a flowing product moving through the pipe which is constantly cooling your weld creates the problem of the weld cooling too rapidly and can lead to HAZ (heat affected zone) cracking.The more obvious problem is the fact that you’re creating a heat source of over a thousand degrees only a few millimetres away from these flowing flammable materials.This paper looks at some of these risks and how to mitigate them to safely complete these in-service welds.

Keywords:In-service welding, HAZ (heat affected zone) cracking, Weld Hardness, Weld simulation software, Weld procedure qualification, Burn-through/blowout, Thin walled high strength pipe.

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ACKNOWLEDGEMENTS

I would like to thank TheWeldnet welding consultancy for providing me with the results of previous in-service weld procedure testing and also Dr Paul Colegrove for reviewing and advising on the writing of this paper.

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TABLE OF CONTENTSAbstract iiiAcknowledgements ivTable of contents vTable of figures viTable of table’s viiTable of graphs vii1 Introduction 12 Literature Review 2

2.1 In-service welding and Burn-through/blowout 22.2 Development of guidelines for repair and hot tap welding on pressurized pipelines. 2

2.2.1 Battelle results 32.3 Increasing risk of burnthrough for thin walled-high strength pipelines 4

2.3.1 3D Modeling and pipewall cavity 52.4 Heat affected zone cracking (HAZ cracking)8

2.4.1 Hardenability and its effect on microstructure 82.4.2 But what’s hydrogen got to do with it? 112.4.3 How does hydrogen cause cracking? 12

2.5 Welding Processes 142.5.1 Temperbead welding 15

3 Methodology 163.1 Qualifying the weld procedure for use on an in-service weld 163.2 How to work out the carbon equivalent on an unknown pipe material 173.3 Calculating cooling rates 173.4 AS 2885.2 Guidelines for running a PQR 183.5 Weld test simulation 19

3.5.1 PQR test piece set up. 203.5.2 Welding the PQR 203.5.3 Testing 21

4 Results 234.1 Previous PQR data 234.2 New PQR modelled data 244.3 Actual results 25

5 Discussion 286 Conclusion 307 References 31

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TABLE OF FIGURES

Figure 1 Type of pipe wall failure (bulging) due to in-service welding, and type of pipe wall failure (burn through) due to in-service welding. 4

Figure 2 Modelled metallographic compared to the actual macro. 6

Figure 3 Casual factors involved in burn-through during in service welding. 7Figure 4 continuous cooling transformation diagrams 9

Figure 5 Microstructure hardness values 10

Figure 6 BS 4515.1 Positions for the hardness reading to be taken for a fillet weld. 10

Figure 7 Factors influencing the possibility of hydrogen assisted cracking during in service welding. 11

Figure 8 Diffusion of hydrogen from weld metal to HAZ during welding. 11

Figure 9 The effect of hydrogen on crack propagation. 12

Figure 10 Hydrogen diffusion in cracks. 12

Figure 11 Classic Hydrogen crack. 13

Figure 12 Effect of hydrogen cracking. 13

Figure 13 Weld bead sequence for temperbead welding. 15

Figure 14 AS 2885 suggested in-Service welding test assembly. 18Figure 15The air/mist cooling equipment. 19

Figure 16 PQR test piece set up. 20Figure 17 PQR weld joint sequence. 21

Figure 18 Completed welds. 21

Figure 19 Macros for weld one. 26

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TABLE OF TABLES

Table 1 Example of a Battelle table.3

Table 2 Welding processes suitability 14

Table 3 Previous PQR data.23

Table 4 Thermal modeling from the worst case service conditions. 24

Table 5 PQR running sheets. 25

Table 6 The hardness survey of weld one. 27

TABLE of GRAPHSGraph 1 CSIRO modeling Graphs. 6

Graph 2 CSIRO effects of hardness 7

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1 Introduction

What are the reasons for carrying out in-service welding? The main reason is economic. The costs involved in shutting down a gas line to carry out the works using conventional welding may mean a full refinery or chemical plant would have to be shut down for considerably longer than the 6 hours or so to complete the welding activity. The plant would have to be run down in a safe and systematic fashion, step by step. The line would have to be isolated; hopefully an isolation point is reasonably close. The line would have to be degassed prior to the welding operation and purged. (Ref 1)In addition there is an increased risk of a breakdown occurring when the plant is brought back online. One of the concerns with in-service welding is the risk of “burn through” or “blowout”. This is where the pipe fails due to the internal pressure and loss of material strength through the localized heating of the welding operation.The second issue is the rapid cooling the pipe caused by the product flowing through the pipe during in-service welding which can lead to Heat affected cracking (HAZ cracking) months or even years after the welding has been completed. In the following pages we will show that this type of welding can be carried out successfully using Industry recognized modeling software backed up by empirical data from successful PQR’s using simulated conditions.

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2. Literature Review

2.1 In-service welding and Burn-through/blowout

Whilst carrying out in service welding the material in the weld pool region must have sufficient strength to maintain the pipewall integrity to contain the internal pressure. The pipe temperature, hoop stress acting on the pipewall, pipe material and thickness must all be considered.

2.2 Development of guidelines for repair and hot tap welding on pressurized pipelines.

In the late seventies and early eighties extensive experimental work was carried out at Battelle Columbus Laboratories by for “Pipeline maintenance personnel to select welding and other parameters to optimize the chances for obtaining welds of satisfactory structural integrity on pressurized natural gas pipelines”. (Ref 2)This was published in 1981 as “Development of guidelines for repair and hot tap welding on pressurized pipelines” and has been used extensively ever since for in-service welding using the MMAW welding process either with cellulose or low hydrogen electrodes. It is normally referred to as “Battelle’s tables” as the results were published in graph and table form for quick reference to select safe parameters for welding.The design curves used in the graphs were predicted by 2 Dimensional thermal analysis modeling and actual experimental results. Joint geometry, alloy content, flow rate and gas pressure were all used in the modeling. The time of cooling from 800 C to 500 C (T8/5) was used as a guide for estimating the likely hood of martensite formation as this had a direct relationship to the time cooling rates at lower temperatures where this would take place. The Battelle laboratories carried out further research that was published in January 1983 as “Experimental verification of hot tap welding thermal analysis to repair and hot tap welding group”. “The purpose of using the models is to define acceptable welding heat input parameters that will avoid either burn-through and release of the pipe contents or excessively high cooling rates that may cause adverse micro structural features and/or cracks at the welds”.(Ref 3)To verify the thermal analysis models 143 test welds were made in total. 90 fillet welds similar to that used for welding a sleeve to a pipe and 53 groove welds similar to that used to attach a branch pipe to a main pipe. A range of pipes were used from 0.125 inch (3.175mm) to 0.5 inch (12.7mm), 6 different pipe sizes. The pipes were pressurized with flowing nitrogen gas.The temperatures of these pipe welds were monitored both inside and outside to see if the rates of cooling to see if they matched the predicted temperatures. About 1/3 of the welds were sectioned to ascertain the hardness and

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microstructures of the Heat Affected Zone (HAZ). The results were found to correspond to that predicted in the modeling.

2.2.1 Battelle results

It was found that the internal diameter (I.D.) temperature increased as the pressure and flow decreased and inversely the higher the pressure and flow for a given heat input the lower the internal surface temperature. This is as would be expected with the heat sink effect of the flowing product.During the experiments it was found that an I.D. temperature of 982 C for hydrogen controlled (E7018) and 760 C for cellulose (E7010) electrodes was the maximum allowable surface temperature for preventing burn through. The differences in the 2 allowable temperatures are a direct result of the deeper penetration characteristics of the cellulose electrodes.In all the testing the only burn-through achieved were intentional and some of the attempted burn-through were unsuccessful. These unsuccessful attempts were based on the thermal modeling, with the cutoff of 760 C for cellulose and 982 C for the hydrogen controlled.

On one successful burn-through the recorded I.D. temperature was actually 1288 C. This shows that the figures of 760 C and 982 C were conservative

figures. This burn-through was on a 0.188 inch (4.77mm) test plate without a flowing medium to act as a cooling medium.

Table 1 The above table was taken from “Development of guidelines for repair and hot tap welding on pressurized pipelines” and shows predicted safe welding variables.

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2.3 Increasing risk of burnthrough for thin walled-high strength pipe lines

Most research carried out has been on thick walled low strength pipe. This has been fine for the conditions for X52 or X60 transmission pipeline in the US where typical wall thicknesses of 9.5mm are common. Over the last few years in Australia in particular higher strength materials have been used for pipeline construction. This has led to thinner pipes being used as the allowable pipewall thickness is inversely proportional to the material yield strength. (Ref 5)

This has led to significant cost savings due to the use of thinner pipe walls. By substituting X80 (551MPa) for X60 (413Mpa) this allows the pipe wall to be reduced by 25% for the same strength. At normal ambient temperature these higher grade materials have a significantly higher strength than the lower grades however when heated, as in welding, there is very little difference. This has led to in-service welding being much more difficult to predict due to the wall thicknesses being thinner than what was formally thought safe to weld on (Battelle Tables) and the increase in the carbon equivalent used in order to increase the strength. So the research carried out in the past are of limited use on these higher strength pipelines. (Ref1)

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Figure 1 Type of pipe wall failure (bulging) due to in-service welding (upper diagram), and type of pipe wall failure (burn through) due to in-service welding (lower diagram). (Ref 4)

2.3.1 3D Modeling and pipewall cavity

The Commonwealth Scientific and industrial Research Organization of Australia (CSIRO) recognized that whilst these thinner walled pipelines made economic sense they made welding on them much harder. During welding the thinner walls are more prone to burn through due to the strength loss. The thinner walls are also more easily cooled by the flowing product and hence the increased risk of HAZ cracking. The CSIRO commissioned a series of reports into In-service welding and in a survey of the Australian pipeline industry welding discovered that welding procedures and practices in use at that time were very conservative. With gas flow and pressure being reduced for the welding. It was estimated that in the four years from 1998 this cost the Australian industry between $2 and $4 million. (Ref 6) The project hoped to fill in some of the spaces left from the previous research where it affected the different industrial conditions found in Australia. From this research a new 3D model was created that did not rely on the maximum ligament temperature from Battelle.Instead it considers that due to the loss of strength exhibited by the heated metal of the weld pool is similar to the loss of overall strength that occurs during corrosion pitting of metal. The 3D model develops an Isotherm pattern around the weld pool to predict where the material is expected to have no strength. Outside of this isotherm the material is assumed to have full strength. This area of weakness is then considered to be a cavity in the pipewall. If this cavity is then taken out then this creates a reduction in the effective wall thickness. This remaining effective wall thickness is then used to predict when the pipes internal pressure will become too great and overcome the remaining strength causing burn through or blowout. It is thought that because weld orientation and working pressure are used in this model that it will be more accurate than Battelle modeling. To use this information a modified B31G equation for assessment of corrosion cavities is used.

SMP= MAOP.(1+69SMYS ).(

1+0.85 . dt

1+0.85 . dt. 1M

)

Where M+ Folias factor

M=(1+1.255

2 . L2

D. t -0.0135

4. L4

D2 . t2)1 /2

SMP= Safe maximum pressure D=Pipe diameter.MAOP=Maximum allowable operating pressure d=Cavity depth SMYS=Specified minimum yield strength t=Pipe wall thicknessL=maximum cavity width in the axial direction

This method of simulation is only good to predicting down to a cavity of 80% of wall thickness after small alterations to the equation have to be used.

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Figure 2 The above shows modeled metallographic (left) compared to the actual macro of a test weld using the modeling in this CSIRO project.

The above shows the acceptable working range for in-service welding on a 5mm thickness pipe with an internal pressure of 5MPa using the CSIRO modeling.

Graph 1 The above 2 graphs (CSIRO modeling) use the predicted remaining effective wall thickness to show effect of the reduced effective wall thickness on an X70 pipe. On a 4mm thick pipe burnthrough will occur using a heat input of1.2kj/mm at only 0.1 Maximum allowable operating pressure but if this heat input is reduced to 0.6kj/mm then this value will rise to 0.8 Maximum allowable operating pressure.

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Graph 2 The 4 graphs above show the predicted effects of preheat on hardness with the inverse effect this has on the ligament temperature and consequently the reduced effective wall thickness using the CSIRO modeling. (Ref 7)

Figure 3 Casual factors involved in burn-through during in service welding

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2.4 Heat affected zone cracking (HAZ cracking)

Catastrophic failure can occur during normal service conditions through Heat Affected Zone Cracking due to the effects of Hydrogen. (Ref 8)First of all when we are talking about Hydrogen cracking we must understand the conditions that need to arise in order for Hydrogen cracking to occur.For HAZ cracking to occur we need diffusible hydrogen, a susceptible microstructure and local concentrations of stresses, either residual or applied.

2.4.1 Hardenability and its affect on microstructure

Hardenability can be described as the “susceptibility to hardening by rapid cooling” or “The ease with which steel can be quenched to form Martensite”. Steels with high hardenability form Martensite even on slow cooling. It is not an indication of how hard a material is but more of a measure of its potential to become hard through quenching.In order for Martensite to form from Austenitic grains, which steel is made of above 800C, it must go through a phase change to form Martensite. If the rate of cooling is delayed (through the use of preheat etc) then a softer phase is created called pearlite or bainite as opposed to the harder and more brittle Martensite. The time taken to form the favourable pearlite is a direct result of the carbon content of the steel and/or other alloying elements. The main alloying elements that have an effect on this time are Manganese, Nickel, Copper, Chromium, Molybdenum and Vanadium. The equation below is used to predict the likelihood of HAZ martinsite forming.

Carbon Equivalent (CE) =%C + %Mn

6 + %Ni+%Cu

15 + %Cr+%Mo+%V

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This equation (there are others for higher alloy and TMCP steels) estimates the influence of each element has on the formation of martinsite. Carbon has the largest single effect (it is added to steel to increase the mechanical properties of steel e.g. tensile strength which is a function of martinsite formation) and is weighted as a full percentage the others have less effect and therefore only a percentage of each is used. Once this carbon equivalent is established it is used to estimate the preheat necessary to slow down the cooling rate so we can achieve the formation of pearlite rather than martinsite.

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The above continuous cooling transformation diagrams for a steel containing 0.76% carbon shows that if the material cools at rates of 140C or greater (IE to the left of the 140C line) the material will form martensite when it reaches approximately 220C in this case. The critical cooling rate depends on the material composition, if the carbon content (or carbon equivalent) is lower than the critical cooling rate line shifts to the left.Therefore lower carbon steels require higher cooling rates for martinsite to form IE the time available to form martinsite will be much less.For steels with higher carbon content the opposite is true in that the line will move to the right giving the Martinsite more time to form.With the cooling effects of the product flowing through the pipe is would be very easy for the Austenite to form straight into the martinsite range. With the use of preheat and the relatively high minimum inter-pass temperatures the goal is to form into a microstructure that is less prone to cracking.

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Fig 4 continuous cooling transformation diagrams

The above graph shows the different hardness values, at room temperature. As can be seen the martinsite has much higher hardness values. The hardness value is also a good indication of the other mechanical properties of the material EG strength, microstructure, cracking susceptibility, likelihood of plastic collapse, ductility etc (Ref 9) AS 2885.2 states that the results of its hardness test for Procedure qualification shall be no greater than 350HV (250HV for sour service). This is considered to be a conservative limit for low hydrogen welds. (Ref 10)

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Fig 5 Microstructure hardness values

Figure 6 to the left (taken from BS 4515.1) shows the positions for the hardness reading to be taken for a fillet weld.“In general, increased rates of cooling lead to higher hardness levels for a given steel composition. Combined with the introduction of diffusible hydrogen through the welding process, and the increased levels of diffusible hydrogen trapped within the weld zone at higher cooling rates, this increase in hardness causes an increase in susceptibility to hydrogen cold cracking.” (Ref 11)

Figure 7 Factors influencing the possibility of hydrogen assisted cracking during in service welding

2.4.2 But what’s Hydrogen got to do with it?Hydrogen can enter the weld through various ways from surface contaminants (grease or moisture), oxy-fuel from heating, atmosphere and the flux on the welding electrodes.Whilst the weld metal is in the austenitic stage hydrogen is readily dissolved in the weld pool but as the weld solidifies and cools to ferrite and pearlite the hydrogen is rejected due to the lower solubility of hydrogen in ferrite. This causes the hydrogen (“in its atomic form, H, as opposed to the molecular form H2”) to diffuse to the HAZ zone near the fusion boundary which at this stage is

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still austenite. The austenite has a lower diffusion coefficient for hydrogen than ferrite so the hydrogen stays in this zone rather than diffusing further across the fusion line and into the ferrite structure of the base material. The austenite then transforms into martensite (If the cooling conditions allow).

The above diagram shows the passage of the hydrogen from the weld pool and into the area where hydrogen cracking is likely to occur.

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Figure 8 Diffusion of hydrogen from weld metal to HAZ during welding

2.4.3 How does Hydrogen cause cracking?

Several theories have been put forward as to why the hydrogen causes cracking by Troiana, Petch, Beachem, Savage etc (Ref 11) but it seems that hydrogen leads to:

Stress induced hydride formation and cleavage Hydrogen enhanced localized plasticity (HELP) Hydrogen induced decohesion (ref 12)

A good analogy for hydrogen cracking in the susceptible microstructure as opposed to pearlite is: Imagine a chisel placed into a crevice or crack of a brick (martensite) and another placed into a freshly cut log (pearlite). If the chisel is hammered into the brick it will crack because it is hard but brittle. If the chisel is hammered into the green log the wood will not split but will yield to the chisel because it is softer and more ductile. The hydrogen builds up as more and more is expelled from the weld metal. This hydrogen creates an internal pressure with the resulting cracks forming.

Figure 9. The above photographs show the effect of hydrogen on crack propagation.

Figure 10. The diagram to the left shows how the hydrogen diffuses forward of the crack and makes the cracks in the previous photos take place.

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Figure 11. The photo to the left shows a classic Hydrogen crack forming at the toes of the weld where high stress concentrations susceptible microstructure and the availability of hydrogen have led to this crack.

Figure 12. The photo to the left shows the effect of hydrogen cracking on pipe socket weldment.

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2.5 Welding Processes

The current process used for in-service welding is almost exclusively SMAW welding.

John Norrish (Ref 13) of Wollongong University looked at different process and possible future developments in welding. He concluded that SMAW would become increasingly challenging on thinner and higher strength pipelines in the future. This is because it is a manual process that needs a highly skilled welder to correctly keep within the safe welding parameters. A slowing down of travel speed can have a big difference in the heat input. He looked at other process and made this table of current welding processes and their suitability for in-service welding.

Table 2 Welding processes suitability

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He concluded the variation in travel speed was the biggest problem with SMAW. And that the Pulsed GMAW using an automated process would give much more predictable results and with a lower potential for HAZ cracking due to it being a low hydrogen process.

2.5.1 Temperbead welding

A type of SMAW electrode called a “Post weld heating electrode” was developed that has a flux coating and leaves no weld deposit. When used it was known as a non-depositing electrode.British Gas introduced this electrode in the 1980’s and it was shown to be effective in lowering the hardness of the welded joint by as much as 100HV by tempering the course grained HAZ of the weld close to the cap. It was hoped that this would dramatically reduce the likelihood of HAZ cracking.This electrode was used on the weld toes on both the pipe and welded sleeve.This technique was used successfully for a while but on one occasion when this tempering technique was used on a previously sound weld, problems were found. Cracking occurred after this tempering weld that were not previously present. On a subsequent investigation it was found that the arc length of 20-25mm needed to use these rods had proved too difficult to maintain and in fact several arc strikes were found in the weld area.It was intended that the intended tempering temperature of 650C was exceeded in several areas above the solid transformation temperature required causing hard spots. Because of these problems the use of this technique was discontinued. (REF16)The use of a temperbead weld using conventional rods can be used instead of the non depositing ones (in fact it’s quite common) tried by British Gas. Because in-service welding normally uses a stringer bead technique the subsequent welds help temper the previously deposited welds underneath. However where the weld is most susceptible to cracking, the weld toe the weld would have cooled for too long before the next weld is deposited. To overcome this when the final weld is deposited on the pipe the next weld overlays this whilst it is still relatively hot. The diagram below show how this welding sequence is used. (Ref 17)

Figure 13 Weld bead sequence for temperbead welding.

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3 Methodology

3.1 Qualifying the weld procedure for use on an in-service weld

Now we’ve looked at some of the problems to avoid whilst carrying out in-service welding now let’s see how they can be overcome. A client recently commissioned the welding consultancy where I work (The Weldnet) to help them develop a welding procedure for an in-service weld.The client supplied the following information from a previous in-service that had recently been carried out:

1. Pipe contents: Natural gas2. Linear flow rate: 0.88m/s to 1.38m/s3. Service pressure: 8.0MPa4. Service temperature: 20C5. Run pipe: DN 450 x 6.75mm 6. Sleeve DN 450 x 25.4mm WT CEq 0.397. PQR records that were used with the above variables.

The client supplied the following information for the service conditions of the pipe to be welded.

1. Pipe contents: Natural gas2. Linear flow rate: 0.88m/s to 1.38m/s3. Service pressure: 8.0MPa4. Service temperature: 27C-29C5. Run pipe: DN 450 x 7.98mm 6. Sleeve DN 450 x 19.05mm WT CEq 0.39

Using the information supplied by the client we used the PRCI (Pipeline Research Council International) “Thermal Analysis Model for Hot-Tap Welding-V4.2” to determine what welding parameters would provide acceptable HAZ hardness values. The modeling also predicts the internal pipewall (ligament) temperature.

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3.2 How to work out the carbon equivalent on an unknown pipe material

Quite often the carbon equivalent is unknown with the relevant certification of the pipe material mislaid or never kept in the records in the first place. The ideal method of working out the chemical composition, and through that the carbon equivalent is to use a portable Atomic Emission Spectroscopy appliance (AES).The test procedure for the AES is as follows: It is carried out on the coupon which is to be removed during the hot-tap. The paint coatings etc must be removed so the test can take place on bare metal. The area to be tested is NDT’d to ensure no laminations or unduly thin pipewall thickness’s are present. This is necessary as what you’re actually doing during the test is similar to striking a welding arc on to the pipe wall and analyzing the chemicals as they vaporize. The area must be thoroughly degreased with solvents to ensure cleanliness. The analysis can now be carried out. The results should be available in a few minutes as a computer printout. The area of the test is now basically an arc strike on the pipewall and it should be considered as such. It would be dangerous to leave it as it could act as a crack initiation point. It may be months before the actual in-service weld takes place (it may even get cancelled). So the area must be blended into the pipewall. Care must be taken not to reduce the wall thickness to below the design thickness. Afterwards the area must be tested using crack detection methods; magnetic particle inspection is normally carried out. (Ref 14)

3.3 Calculating cooling rates

Although the Battelle tables can be used to predict cooling rates of the pipe if you know the pressure, diameter, flow rate etc another, more simple, method has been developed.The Edison Welding Institute (EWI) devised a simple field test to measure the actual pipes heat sink capabilities. The areas that will be welded upon are prepared as for the previous chemical analysis with the pipework cleaned back to bare metal and NDT’d for thickness/laminations. The areas to be welded are marked out with chalk and 50mm circles drawn at suitable locations around this chalked line. The 50mm circles are then heated with an oxy-fuel torch (being careful to move the heat around inside the circle and not to dwell in one spot) till it reaches a temperature of 300-325C. The temperature is measured with a suitable temperature gauge and the time it takes for the temperature to fall between 250C and 100C is measured with a stopwatch. After checking the temperature is normal on the next test position the procedure is repeated on the upstream circle on the opposite side of the pipe and an average of these times is calculated. (Ref 15)

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3.4 AS 2885.2 Guidelines for running a PQR

Australian standard AS 2885.2 Pipelines-Gas and liquid petroleum Part 2 Welding requires that a weld procedure must be qualified for welding on live pipelines and that the normal qualification without testing criteria (clause 5.4.4) does not apply in these cases. The purpose of the welding procedure being to ensure the weld has the required hardness, ductility, strength and mechanical properties and is risk free of Hydrogen Assisted Cold Cracking (HACC). It is also a requirement that the procedure must “design out” the chance of HACC happening.The normal ranges of essential variables that are used for welding procedures do not apply to welds that can be directly affected by pressure and cooling effects of the product (I.E. the fillet onto a sleeve). But they can be used on the longitudinal welds that are not directly welded onto the live pipe if the joints have a backing strip to prevent penetration into the wall of the pipe.A simulated procedure must be run if the weld will be affected by the products pressure or cooling effects, I.E. directly onto the live line.

Figure 14 The above diagram is a suggestion (AS 2885.2) for testing the procedure with water flowing through the test piece to simulate the products cooling effect.The test must be carried out under conditions that simulate the worst case that will be found during the actual weld. For example if a delay is expected in completing the joints, for NDT to take place for example, then this must also happen during the simulation. The welding preheat, burn off rate and heat inputs will all be as close to the lower end of the qualification to simulate the worst case scenario.

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3.5 Weld test simulation

By using the suggested method from AS 2885.2 to simulate the cooling effect it was thought that cooling would be too vigorous to simulate the cooling time of 60.9 seconds (this being the measured heat sink capacity of the service pipeline). Clause 13.16 of this code suggests if the recommended method produces a more severe cooling than that required then another media be used. Water mist or motor oil are alternative suggestions.So an alternative heat sink capacity method was used. A water mist and compressed air spray was used instead.

Figure 15 The air/mist cooling equipment

As can be seen above a calibrated pressure gauge was used to help monitor and the flow and mixture of air and water was adjusted till the desired cooling effect was established.Measurements were taken at 4 different locations around the pipe during these trials. The positions were chosen to simulate the worst case for the weld IE the positions around the circumference where the fillet weld will take place.

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3.5.1 PQR test piece set upThe test piece was a 5M long pipe, same CEq, dimensions etc as the in service pipe (It was thought to be an off-cut of the original pipe). It was tilted 5° to stop pooling of the coolant at the 6 o’clock position. As 2885.2 Welding Positions Designation (4.1) allows for a ± 5° from the nominal test position.

Figure 16 PQR test piece set up

The gap between the sleeve and pipe was adjusted using 2 chain blocks (used to hold the sleeve in position) attached and a gap of between 1mm and 2mm was achieved. The longitudinal seams in the 2G position had a 3mm root gap. These seams had a backing bar to prevent the weld directly penetrating the line pipe.

3.5.2 Welding the PQRLPG was used to apply the pre-heat but this took too much time to achieve and keep the work piece at the minimum pre-heat temperature with the cooling medium used. When this was switched to Oxy/LNG the Pre heat of 150C was easily maintained.The test piece was welded out using 3.2mm and 4mm low hydrogen vertical down electrodes.The 2G seams were welded simultaneously. The contraction of this welding pulled the gap between the sleeve and the line-pipe down to a maximum of 1mm. The gap was necessary so as to allow for the contraction during the fillet welding and so helping to minimize the residual stresses left in the weld.One fillet weld was completely welded out before progressing onto and welding the next fillet. The purpose for this sequence is to further try and control the residual stresses in the fillet welds so as to help “design out” the factors that contribute to HACC cracking.

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Figure 17 PQR weld joint sequence

A total of 5 welds were carried out. W3 and W4, 2G butt welds with backing were welded first, followed by W1 then W2, both 5F vertically down and finally W5 in the 4F position to simulate a worst case cap repair.

3.5.3 TestingA visual examination of the completed weld was carried out to AS 2885.2 tier 1 (clause 22.2) criteria. This inspection was followed by MT (magnetic particle) inspection to confirm the absence of immediate flaws.UT (ultrasonic testing) and MT examination were carried out 48hrs later. AS 2885.2 only requires a 24hour waiting period but 48 hours is again a worst case period. These tests were carried out to AS 2885.2 tier 1 acceptance criteria.

Immediate NDT Final NDT

Figure 18 Completed welds

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The following mechanical tests were carried out to AS 2885.2 section 6 requirements.Weld 1: 1 macro and hardness survey at the 12, 3 and 9 o’clock positions. The 6 o’clock position was covered in weld 5 (below). Weld 2: 1 macro and hardness survey at the 12, 3, 9 and 6 o’clock positions

The 3 and 9 o’clock positions were taken approximately 50mm below W3 and W4 so as not to include any tempering effects off the butt welds.

Weld 3 and 4: 2 x macros and hardness survey, 4 x transverse tensiles, 2 x 3 sets CVN (Charpy V-notch test) at the WMCL at -20C and 2 x 3 sets of CVN in the HAZ at -20C.Weld 5: 1 macro and hardness survey at the 6 o’clock position (cap repair weld).The acceptance criteria for the mechanical tests:

CVN: Weld and HAZ 41J average and 34J minimum. Hardness: 350 HV10 maximum. Tensile: Minimum 483 MPa.

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4 Results

4.1 Previous PQR data

Using the results provided by the client for the previous in-service weld we modeled it on the service conditions for the new hot-tap pipe (see methodology). The Average cooling time from 250C to 100C on the previous weld was 38 Seconds. This was assuming a 150C preheat. PRCI software predicted the following results (for the fillet welds only).

Electrode diameter and pass

Heat Input(kJ/mm)

Cooling rate800C-500C(seconds)

Ligament temperature(C)

Predicted HAZ hardness (HV10)

2.5-pass 1 0.40 1.17 475 4502.5-pass 1 0.76 (actual) 2.90 619 4002.5-pass 1 1.00 4.09 684 3703.2-pass 2 0.40 1.17 475 4503.2-pass 2 0.71(actual) 2.69 606 4053.2-pass 2 1.0 4.09 684 3704.0-pass 3 0.50 1.70 531 4334.0-pass 3 0.80(actual) 3.10 631 3944.0-pass 3 1.2 5.08 726 348Table 3 Previous PQR data

Figures in red were the actual results from the previous test.As can be seen in the table only 2 of the passes achieved an acceptable hardness value (in blue) of less than 350HV. The actual heat inputs gave hardness values well outside the acceptable limits. The PQR passed hardness testing on the previous testing (not available to us) leading to the assumption that during the testing the cooling regime was insufficient to achieve the cooling rate of 38 seconds. From these predicted results the previous WPS was rejected for use on this weld.

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4.2 New PQR modeled data

Using the service conditions provided by the client (see methodology) for the service conditions on the new pipe we modeled this new data with the PRCI software to establish the ideal heat input and preheat for the PQR weld test. The results of the spectroscopy test confirmed the CEq of the line pipe was 0.40. The Average cooling time from 250C to 100C was found to be 60.9 Seconds.

Case Heat input

(kJ/mm)

Maximum cooling rate

@ 538C (C/sec)

Cooling rate 800C-500C(seconds)

LigamentTemperature

(C)

Predicted HAZ

hardness(HV10)

1 0.75 33 2.53 538 4102 1.00 16 3.71 598 3793 1.25 5 5.23 649 3454 1.5 0 6.66 686 320

Table 4 Thermal modeling from the worst case service conditions

The results seem to show that the model predicts acceptable hardness values only for the heat input of 1.25 and over. These predictions closely follow the Battelle table’s results The chief Welding Engineer, Chris Cobain, when reviewing this data used is prior experience to estimate that the software estimates were too conservative. It was hoped that a heat input of 1Kj would give the required hardness of <325HV10.

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4.3 Actual PQR resultsThe following results have been taken from weld number W1 5F vertical down fillet (section 3.5.1).

Table 5 PQR running sheetsAs can be seen above for weld number 1the heat inputs range from 0.92 Kj/mm to 1.44 Kj/mm. The running sheets from the other welds in the test show a similar range and are close to the desired 1-1.5 Kj/mm

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Figure 19 Macros for weld one

The Macros above are all from weld one and comply with the code requirements.

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Table 6 The hardness survey of weld one

The tables above are of the hardness survey from weld one as can be seen the hardness values range from 164-298 HV10.

All 5 welds were tested and each complied and had similar hardness values. The Mechanical test results have been left out of the results but they all complied satisfactorily.

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5 Discussion

It is a widely held belief that the Battelle research and findings over estimate the cooling times for thin walled pipes and the opposite for thick walled pipes. That and the fact that the testing was performed on pipes of up to X52 can make them inappropriate for thinner higher strength pipelines.

If the cooling time for thin walled pipes is overestimated (IE the cooling rate predicted is greater than what actually happens) then this can lead to an overestimation of the heat input allowed to produce a safe weld.

In-service welding is a balancing act between keeping the hardness values below 325HV10 and an internal ligament temperature of less than 980C.

The PRCI software predicted (table 3) that the hardness values would be in excess of the acceptable values of 325 HV10. If we had followed this software tool we would have increased our heat input to the more acceptable range to 1.5-2.0 kj/mm heat input. Whilst this would certainly lower the hardness values it would have increased the risk of burn through, although the same modelling predicted that the chance of this was remote.

Previous experience of using the software led the Engineer to be wary of some of the results. It’s an expensive and time consuming business carrying out weld procedure qualification testing. By using the software to establish yardsticks of predicted results saves a lot of experimental hit and miss. But the software is only a tool to assist and is not the answer to all the questions of in-service welding. If it were there would be no need to carry out weld procedure qualification testing and we could use the results to go straight on and weld the live pipeline.

The software overestimated the predicted hardness value by a large margin. The predicted hardness for 1.00-1.5 kj/mm (table 3) was 379 HV10 the actual maximum hardness achieved in testing was 298 HV10. We did not have thermo-couples on the inside wall of the pipe to measure the ligament temperature so we have no idea of the wall temperatures achieved during the procedure testing but it is certain that if we’d followed the recommended heat input of 1.5kj/mm and over it would have been a lot higher.

The method of cooling used, Figure 15 air/mist does not simulate the internal pressure of the actual weld conditions but only the effect on the resulting hardness values using the real heat sink capacity of the pipe to be welded upon. The software predicted that the heat inputs of Table 3 would not result in burnthrough. The Battelle modelling puts more emphasis on the ligament temperature, 980C max, than it does the effect of internal pressure on

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burnthrough. The testing of the Battelle model was carried out on relatively thick pipes where the remaining strength of the pipewall below the weld pool gave the pipe a larger margin of safety over a thin walled pipe.

The CSIRO software (Graph 1) used more than the internal ligament temperature. It used welding orientation, heat flow, internal pressure and the Yorioka-1 carbon equivalent formula (this was found to have a better correlation than the conventional formula on X42 to X80 pipes) in its modelling. The CSIRO method of modelling the loss of through wall thickness strength due to the effect of the weld pool made a lot of sense. To then estimate this loss of strength as a form of cavity saves a lot of research as pipeline codes, EG AS2885 already allow for this and have appropriate calculations to estimate the remaining strength.

Unfortunately this software did not get developed enough for a commercial release. One of the limiting factors with this and the Battelle modelling software was that the large amount computing effort taken to calculate the results made them prohibitive. This is less of a problem today and hence the availability of software such as the PRCI package.

Tempering beads are normally used to lower the hardness of the weld run most prone to high hardness values and therefore cracking, normally the toe weld (figure 13). The approach tried by British gas (Ref 16) in the early eighties for the use of a Post weld heating non depositing electrode was very encouraging with the a reduction in hardness values of up to 100HV10. This would allow welds to be deposited at very low heat inputs and so minimising the risk of burn through. These welds would then be tempered with a higher heat input, but now onto a pipe wall thickness that had already been reinforced with the first weld. Obviously care must be taken to ensure that the heat inputs, or the effect of the tempering electrode, are high enough to ensure satisfactory weld metal properties. It’s unfortunate that this method, if used incorrectly, sometimes created a HAZ cracking problem when none existed before.

Most in-service welding is carried out using the SMAW process. This relies on the skill of the welder and is therefore prone to operator error. Changes in travel speed, arc length and angle of the electrode can all have a huge impact on the completed weld. John Norrish (ref 13) looked at the other alternatives to SMAW welding and concluded that a pulsed GMAW was the most predictable in its results. Pulsed GMAW has better low hydrogen capabilities than low hydrogen SMAW. Although the welding equipment can have electronically regulated control it is still a manually operated process albeit semi automatic. To eliminate the human error he thought that to fully automate the process with an orbital machine would give the most predictable results. The problems with pulsed GMAW is it can be problematic for field use and shelters must be erected to shield the weld during use. That’s why SMAW is continually used due to its versatility, especially for field use.

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6 Conclusion

The use of thinner higher strength materials for pipelines will create more difficult and onerous conditions for carrying out in-service welding. The increased use of software simulation used in conjunction with effective weld procedure qualification testing will give results that can accurately predict the necessary variables needed to complete a successful weld.

How these variables are used in the field are of the utmost importance. The welding variables must all be tightly controlled with constant monitoring to ensure compliance.

The introduction of fully automatic welding packages will make this monitoring easier with all the works being carried out from a safe distance.

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7 REFERENCES

1. The prediction of burn-through during in-service welding of gas pipelines. P.N. Sabapathy, M.A. Wahab, M.J.Painter. September 2000.

2. Battelle Columbus Laboratories Report “Development of guidelines for repair and hot tap welding on pressurized pipelines”. J.F. Kiefner, R.D. Fischer and H.W. Mishler. September 1981.

3. Battelle Columbus Laboratories Report “Experimental verification of hot tap welding thermal analysis to repair and hot tap welding group”. R.C. Gertler, F.A. DeSaw and C.R. Barnes, J.F. Kiefner, R.D. Fischer and H.W. Mishler. January 1983.4. The onset of pipewall failure during “in-service” welding of gas pipelines. P.N. Sabapathy, M.A. Wahab, M.J.Painter. December 2004

5. Overview of in-service welding research at EWI. Bruce W.A.

6. In-Service welding of Gas Pipelines, CRCWS Project 96:34 Final Report. M. Painter. March 2000.

7. In-service welding of thin walled, high strength gas pipelines CRC-WS Research. M. Painter

8. Modelling of HAZ hardness in C-Mn pipeline steels subjected to in service welding procedures. IIW Document No. IX-2165-05. D. Nolan, Z. Sterjovski, D. Dunne.

9. WTIA (Welding and Technology Institute of Australia) Hardness testing of welds TGN-PE-01.10. Optimization of fillet weld sizes. Welding journal, research supplement. B A Graville and J A Read. April 1974.

11. Welding Metallurgy, second edition page 412. Sindo Kou 2003.

12. Hydrogen Embrittlement of Pipeline Steels: Causes and Remediation. P. Sofronis, I. M. Robertson, D. D. Johnson. May 2005.

13 Process consideration for in-service welding of transmission pipelines. John Norrish

14 WTIA (Welding and Technology Institute of Australia) Technical note 20 Repair of steel pipelines.

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15 Development of simplified weld cooling rate model for in-service Gas pipelines. Edison Welding Institute.

16 Hot tap welding in the UK gas transmission system recent experience. Boothby P.J. BG Transco plc.

17 Current practice and the gaps in In-Service welding technology a research perspective. M. Painter CSIRO.

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