International Journal of Fracture 97: 249–261, 1999. c 1999 Kluwer Academic Publishers. Printed in the Netherlands. Burnthrough prediction in pipeline welding A.S. ODDY1 and J.M.J. McDILL2 1 Research Associate, Carleton University, 1125 Colonel By Drive, Ottawa, Ontario, K1S 5B6, Canada. Associate Professor, Dept. of Mechanical and Aerospace Engineering, Carleton University, 1125 Colonel By Drive, Ottawa, Ontario, K1S 5B6, Canada 2 Received 24 March 1997; accepted in revised form 4 December 1997 Abstract. Welding of branch connections on gas pipelines at full line pressure is frequently an operational necessity. Weld parameters must be selected so that heat inputs are low enough to avoid burnthrough yet not so low that hydrogen-assisted cold cracking occurs. Current techniques rely on the observation that burnthrough does not occur if peak temperatures on the inner surface are kept below 980◦ C. At these temperatures, rate-dependent flow is the dominant mechanism. The problem is one of creep rupture occurring at temperatures in excess of 980◦ C with times of the order of seconds. Material constitutive models for the analysis of welding must include both rate-dependent and rate-independent plastic flow as well as the effects of phase transformations. Material properties at elevated temperatures are usually not available for pipeline steels and must be extrapolated from values at lower temperatures. An exploratory study using 3D thermal-mechanical finite element analysis of welding on pressurized vessels is presented and includes comparisons with experiment. The agreement is encouraging. The material failure does occur in appropriate locations but the predictions are generally overconservative. Estimated material properties, especially damage and rupture properties at high temperatures could be improved. Key words: Welding, burnthrough, creep, finite element analysis. 1. Introduction Adding branch connections or repairing defects in pipelines is a common industrial problem. Welding at full line pressure and gas flow is the preferred technique but requires careful selection of the welding parameters in order to avoid two problems. If the heat input is too low then hydrogen-assisted cracking of the heat-affected zone (HAZ) may occur. The causes and undesireable effects of hydrogen-assisted cracking are well known; e.g., (Grong, 1994). If the heat input is too high, the pipe wall may locally soften, rupture, and leak. The heat source need not melt through the pipeline wall completely for burnthrough to occur. The latter problem, predicting material rupture and burnthrough during welding of pressurized pipelines is examined here. This exploratory study examines the feasibility of such predictions and illustrates deficiencies in material data for pipeline steels. Current practice uses empirical guidelines to prevent burnthrough. It has been observed that burnthrough does not occur as long as temperatures on the inside surface never exceed 980◦C. In practice, it is difficult to keep temperatures below 980◦ C. In addition, while the the empirical guidelines highlight the principal role played by the heat input, they neglect the influence of the internal pressure which may be present. In any case, rupture of the pipeline can occur even though the fusion zone only partially penetrates the wall. The safe upper limit of 980◦C is well below the liquidus temperature which, for steel, is somewhat less than 1600◦C. In an excellent series of experiments with short welds (Bruce et al.,1993) on water filled, pressurized vessels, one case showed slight thinning of the wall with a fusion zone (FZ) penetration to thickness ratio of only 13 . In another case, substantial thinning, partial rupture and incipient failure was seen with a penetration of only half the thickness. Material failure 250 A.S. Oddy and J.M.J. McDill at these temperatures is clearly due to creep rupture. Of course complete penetration of the molten FZ through the pipe wall also causes failure. The analysis of burnthrough in pipelines requires material properties to describe the thermal, elastic, plastic, creep and rupture characteristics particularly of the base and HAZ materials. All material properties show large variations with temperature. Creep properties, especially those describing damage growth, are well known to be strongly dependent on not only the alloy content and temperature but also on the material processing. For instance, considerable scatter was observed in the short rupture-life data for the same alloy as shown in (Klueh, 1983). This was mainly due to the differences in processing used by different suppliers. These variations in creep response persisted even after identical reprocessing; e.g., heat treatment, of the samples from different suppliers (Klueh, 1983). Abundant evidence exists to show that wide variations in minimum creep rate and rupture life can occur for a specific alloy at a single temperature. The design of constitutive models must capture the essential features of the material’s response but the analyst must always be aware of the likely variation in properties. These are factors which directly influence the repeatability of any experiments, quite apart from the difference between numerical predictions and experimental measurements. The more complex a material model is, the more material parameters are required. Variations in some parameters will have more significant effects on predictions than others. Part of the purpose of this investigation was to identify which material properties have the largest effect on the predictions. At least a dozen different alloys and grades of steel are used in pipelines. It is very unlikely that all the properties and their variation with temperature will ever be known for all the steels currently in service. The effects of in-service degradation add additional difficulties. Designing constitutive models which require large numbers of material parameters will be fruitless if industrial users must guess at appropriate values or if those parameters vary widely. The process chosen here is to use reasonably simple models requiring the minimum number of material parameters. Properties representing conservative bounds; i.e., chosen to overestimate the likelyhood of creep rupture, are used wherever possible. Part of the task was to generate a qualitative understanding of the interactions between material parameters as an aid in estimating properties. Finite element analysis (FEA) simulations of five experimental welds on pressurized, water-filled vessels (Bruce et al., 1993) were performed. The results of the FEA simulations were compared to those of the physical tests. Finite element methods can make accurate predictions of transient and residual stresses in welds. The thermal and mechanical FEA simulations were performed using a wellestablished three-dimensional multipurpose code; e.g., (Oddy et al.,1989; Oddy et al.,1990; McDill et al.,1992) which incorporates a finite deformation algorithm with the Green–Naghdi stress rate and centred strain. The thermal analysis provides nodal temperature histories. The stress analysis takes the thermal history and uses it to compute thermal expansion, stresses and plastic strains. Predicting temperatures and stresses is complicated by the temperature dependence of all the material properties, by phase transformations, as well as the steep temperature gradients around the heat source. Thermal-mechanical analyses of welds in (Oddy et al., 1989; Oddy et al.,1990; McDill et al.,1992; Goldak and Oddy, 1992) show the importance of including phase transformation effects, transformation plasticity and the necessity of three-dimensional (3D) Burnthrough prediction in pipeline welding 251 analyses. The constitutive model divides the total strain rate into the sum of the elastic, plastic, thermal, creep, transformation volume change and transformation plasticity strain rates. ε̇tot = ε̇el + ε̇pl + ε̇th + ε̇cr + ε̇trv + ε̇trp (1) The temperature cycle in welding leads to thermal expansion/contraction and phase transformations. Because both these effects are not uniform over the volume they in turn create stresses, leading to plasticity at low temperatures and creep at higher temperatures. During phase transformations a further phenomenon, transformation plasticity can be important. The constitutive model incorporated in the FEA software of (Oddy et al.,1989; Oddy et al.,1990; McDill et al.,1992) follows the effective stress function algorithm of Kojic and Bathe (Kojic and Bathe, 1987), modified to include phase transformation effects. It allows a mixed elastoplastic-creep material response, avoiding abrupt transitions from elasto-plastic to visco-plastic to linear-viscous material models as temperatures are increased. Predicting burnthrough involves two possible failure modes. The first is complete penetration of the molten pool through the pipe wall. This is not a case of accumulated damage and eventual rupture. It is a case of a plastic instability. The more important problem is the prediction of material rupture as temperatures approach the melting point. If the heat source completely melts through the wall, the liquid metal will be unable to contain the pressurized fluid inside and the molten metal will be blown out. This is more a case of weakness and excessive flow than it is one of material rupture. In fact, the material loses all strength before it melts. This effect is easily included by specifying a zero-strength limit at a temperature just below the melting point. This allows very low stiffness and stress values to be assigned to elements with temperatures above the melting point. In this investigation the elastic modulus is reduced by 5 orders of magnitude for melted material. This ensures that molten material will be unable to resist large displacements. In FEA it is incorrect to assume that this instability will create a bulge-like bubble in the pipe wall. At an instability, small increases in the load cause infinite displacements. The best that can hoped for is increasing difficulties in achieving convergence of the nonlinear solution followed by failure to converge on a solution. At this point the displacements may still not be very large. In addition, real failures occur very rapidly and unless time/load steps are kept extremely small the numerical analysis will be unable to capture the onset of this instability. The more important problem is including creep and creep-rupture since the phenomenon known as burnthrough occurs even though temperatures on the inner surface remain well below the melting point. This has been done in the FEA software of (Oddy et al., 1989; Oddy et al.,1990; McDill et al.,1992) following the method outlined by (Leckie, 1986; Hall and Hayhurst, 1991). The equivalent creep strain rate is ε̇cr = G σ 1−ω n , (2) where G, n are material properties, σ is the equivalent stress and ω is the internal damage (ω = 0 for undamaged, ω = 1 at rupture). The growth rate of the internal damage is computed with ω̇ = C ∆ν , (1 − ω)η (3) 252 A.S. Oddy and J.M.J. McDill where C, ν, η are material properties and ∆ is an isochronous effective stress to account for the effects of multiaxial stresses. For reasons outlined below, if the first principal stress is negative or if σ is zero, then the isochronous effective stress and thus the damage growth rate are zero. In welds, the heating half of the thermal cycle creates large compressive stresses and considerable deformation. At high temperatures where creep is the dominant mechanism, the equivalent stress is large, but all principal stresses are strongly compressive. Furthermore, temperatures are well above those used in hot-isostatic pressing to eliminate voids. Therefore it must be expected that creep strain rates will be large but damage growth will not occur. Here, the isochronous effective stress is used to allow creep damage growth only when two conditions are met simultaneously. First, damage grows only when creep strains occur (σ > 0). Second, damage grows only when the first principal stress is tensile (σ1 > 0). A multiplicative decomposition of the isochronous effective stress is used to enforce both conditions simultaneously. ∆ = σ1α σ (1−α) , (4) where α, the triaxiality factor, is 0 < α < 1. 2. Properties Standard constitutive models in weld analyses require, at a minimum: 2 thermal properties (conductivity, specific heat or enthalpy); 2 elastic properties (elastic modulus, Poisson’s ratio or bulk modulus); the thermal dilation; 4 plastic properties (yield strength and hardening modulus for ferritic/pearlitic and austenitic phases); 6 phase transformation properties (austenite formation start and end temperatures and volume change, austenite decomposition start and end temperatures and volume change). Apart from the phase transformation properties all the properties show strong dependence on temperature. The addition of creep strain rates and creep rupture adds to the number of material properties. Creep rates; i.e., rate dependent plastic strains, require two properties, the creep index n, and coefficient G. The rupture properties are the stress index ν , coefficient C , and damage exponent η , plus the isochronous triaxiality factor α. All of these show some temperature dependency and several are known to be strongly temperature dependent. Burnthrough, or creep rupture during welding occurs at temperatures well above the austenite formation temperatures. While these properties are often available for steels intended for service at high temperatures, material data for pipeline steels in this range are very difficult to obtain. Given the number of different properties and the number of different alloys used in pipelines it is very unlikely that any analyst will have a complete set of properties at all temperatures between room temperature and melting. An important part of any analysis will be selection or estimation of appropriate properties with some informed judgement of the likely effects. Although it is not always easy to judge, wherever possible conservative material properties; i.e., those likely to promote material rupture were selected. In this study, properties of a fine-grained steel and an Si-killed normalized steel were used. The thermal, elastic and plastic properties of a fine-grained steel were taken from (Andersson, 1978), with some modifications. Temperature dependent enthalpy rather than specific heat was used. Above 500◦ C the elastic modulus was taken from (Harste, 1992). Poisson’s ratio was modified from (Andersson, 1978) to keep a fixed ratio between the elastic and bulk moduli. Burnthrough prediction in pipeline welding 253 Yield strengths at temperatures above 900◦C were not permitted to drop below 30 MPa, for reasons to be described later. The melting, or zero strength temperature, was set at 1600◦C. The minimum creep rate for austenite as compiled in (Harste, 1992), is described as ε̇cr = 8.8 × 105 2.3 × 1011 − 9.9 × 107 T ! −32714 6.014 × 10−5 σ exp sinh T T ! σ3 , (5) where T is the temperature in K , and σ is the uniaxial stress in Pa. The strong temperature dependence of the minimum creep rate is clearly evident. This effect arises from several sources. As temperatures rise there is a weakening of the interatomic forces as described by a reduction in the elastic properties, and an increase in the mobility of dislocations. The apparent activation energy for the temperature dependence of the dislocation mobility is also affected by stress as reflected by the hyberbolic sine term. Thus in equation 1, G is dependent on temperature and stress. Equation 5 describes the compilation of minimum creep rate data for carbon steels from 9 different sources, with temperatures ranging from 850◦ C to 1400◦C and strain rates covering up to 5 orders of magnitude. Even so, the scatter in the measured strain rates is such that predicted minimum strain rates for a given stress are accurate to only ± 1 order of magnitude. This however is not exceptional. Measurements of minimum creep rates, under the same temperature and stress conditions for multiple samples of the same material, from different heats can range over 2 orders of magnitude (Boyer, 1988). Creep rupture properties are also strongly temperature dependent. It would be preferable to have rupture data at many temperatures but it is rare for such to be available. More often one is forced to extrapolate from values of ν, C, η and α at a single temperature. Realistic methods of estimating temperature effects are required. The most important requirement is that the rupture life be estimated plausibly. In the absence of data to the contrary, α and η are assumed to be constant since, in a uniaxial case at least, they don’t strongly affect the rupture life. Although η affects the damage growth rate its principal role lies in describing the shape of the creep strain curve with time; i.e., the amount of tertiary creep up to rupture. It is ν , and the combined factor C(1 + η) which determine the rupture life at a given stress. In any case, rupture lives at constant stress and temperature frequently range over 2 orders of magnitude (Boyer, 1988). Various parameters exist for the extrapolation of creep rupture data at one temperature to another temperature. The Larson–Millar parameter was used here. L = T (m + log(t)), (6) where T is temperature in K, t is time and m is a constant which also shows some alloy dependence. For this investigation, and until better data becomes available, m is assumed to be 20 if time is in hours or 16.4 for time in seconds. If the rupture life changes with temperature such that L remains constant one obtains ν2 = ν1 T1 , T2 C2 (1 + η2 ) = (7) (C1 (1 + η1 ))T1 /T2 10 m T1 −1 T2 . (8) 254 A.S. Oddy and J.M.J. McDill Figure 1. Melt-through instability in external spot heating of a pipe. Contours from inner surface to spot centre at 500◦ C, 750◦ C, 1000◦ C, 1250◦ C, 1500◦ C. Displacements magnified 3x. As temperatures increase, ν2 decreases slowly whereas C2 (1 + η2 ) increases by orders of magnitude. For time in seconds and stress in Pa, if the rupture properties for a Si-killed, normalized steel at T1 = 400◦ C are taken as: C1 = 5.42 × 10−72 , ν1 = 7.63 (Boyer, 1988) with η = 5 and α = 0.5 it becomes possible to compute ν2 and C2 at any T2 . The nucleation and growth of new austenite grains makes extrapolation of this data to temperatures above about 750◦ C unjustifiable. Constant values are used above 750◦C. Before undertaking the FEA simulations of the experimental welds, a series of numerical tests on geometries and conditions different from those of (Bruce et al., 1993) were performed to determine the limits on extrapolating properties. One such case was an air filled pipe (0.3 m OD, 0.012 m wall thickness, internal pressure 5 MPa) spot heated (4.6 KJ/s heat input) until the melting instability occurred. Figure 1 shows the displacements, magnified 3 times, in the last step for which a solution was obtained. The 1200◦C isotherm encloses a roughly 3 mm diameter spot on the inner surface. In subsequent steps the nonlinear stress analysis would not converge. If the elastic modulus is reduced too much at melting then elements distort severely and negative volume errors occur. FEA of welds on flat plates without surface loads erroneously predicted material rupture when creep rupture properties were extrapolated to temperatures above 750◦ C. Extrapolating rupture properties above 750◦C makes the material is too weak to withstand any welding cycle. It ruptures during cooling, regardless of the loading. Finally, a combined visco-elastic-plastic model was essential. As would be expected from deformation maps, if stress rates are high enough then dislocation glide can occur even at elevated temperatures. Thus rate-independent plasticity and the yield strength play a role even at high temperatures. If the high temperature yield strength is too great, as it effectively is in a purely visco-elastic model, stress levels have no upper bound and are determined only by the rate of loading. High stresses promote damage growth and therefore overestimate both Burnthrough prediction in pipeline welding 255 Figure 2. Pipe mesh and close-up of refined region. Direction of source motion is from A to B. the rate of damage and volume of ruptured material. If the yield strength is too low then rate-independent plasticity keeps stresses low, reducing creep and damage growth. Here, the yield strength did decrease with increasing temperature but was not permitted to drop below 30 MPa, a value typical of austenite at these temperatures. 3. Boundary Conditions Table 1 has the weld parameters and internal pressures used. Cases A, D, E and F correspond to physical tests from (Bruce et al., 1993) which resembled short repair welds. In all cases the welds were made on a 0.4064 m (16 inch) diameter pressure vessel. Short notches were machined in the outside surface to mimic damaged areas requiring repair. The nominal 6.4 mm (0.25 inch) thick wall had been machined to 3.2 mm thick at the site of the welds. The vessel was filled with water and pressurized. Weld O represents a FEA simulation which had the same thermal conditions as A but did not include the internal pressure in the mechanical portion of the FEA. Identical meshes (Figure 2) were used for the FEA simulations of cases A, D, E, and F as well as case O, although the deposit heights differed. Symmetry was assumed along the weld path. The direction of the weld path is from A to B (Figure 2). There are 1900 elements and 2773 nodes. The variable node, isoparametric, grading bricks of (McDill et al., 1987) are used with selective reduced integration. The thermal FEA had temperature dependent surface convection to air on the outside as in (McDill et al., 1992). Boiling surface heat transfer on the inside was included. Surface flux vs. surface temperature (Figure 3) was computed for the pressure conditions by nondimensionalizing the surface data measured in (Dubost et al., 1988). A generic distribution of y vs T − Tsat , where Tsat is the saturation temperature, was extracted from the experimental data in (Dubost et al., 1988) where, y= q − qmin qmax − qmin (9) Here qmax is the maximum surface heat transfer per unit area and qmin is the minimum surface heat transfer per unit area corresponding to the formation of a stable gas layer. Nondimensional relations for maximum surface flux (Lienhard, 1988) and minimum surface flux (Baumeister, 1978) were used to compute values for qmax and qmin . Heat flow across 256 A.S. Oddy and J.M.J. McDill Figure 3. Surface flux on inside of water-filled, pressurized tests. Table 1. Weld Conditions and Maximum Temperatures on Inner Surface ∗ Weld Current Amps Voltage Volts Source Speed mm/sec Internal Pressure MPa Deposit Height mm Max. Inside Temperature Measured (Bruce et al., 1993) ◦ C Max. Inside Temperature (FEA) ◦ C A O D E F 94.6 94.6 108.6 116.7 111.0 21.3 21.3 21.7 22.1 21.8 4.2 4.2 4.4 4.3 2.8 5.52 0.0 5.52 4.14 1.36 1.3 1.3 1.3 1.6 1.8 520∗ — — 1349 >1600 1230 1230 1380 1380 1520 (Bruce et al., 1993) suggests poor thermocouple placement the surfaces which connect the meshed portion to the unmeshed pipeline was small but was included by using an equivalent surface heat transfer coefficient on those surfaces. The height of the deposit was found to be a critical variable in the thermal analysis, exactly as it was in (Goldak and Oddy, 1992). Changing the deposit height by as little as 0.3 mm changed the peak temperature on the inner surface by 100◦C. In the mechanical analyses, the inside surface was subjected to pressure loads. The surfaces which joined the meshed portion to the unmeshed pipeline received a mix of prescribed displacements and surface tractions. The axial surfaces were not permitted to displace tangentially. Radial displacements appropriate for the internal pressure were applied to the parallel surfaces on the ends. On one end the axial displacements were not permitted. On the other end, surface tractions were applied to create appropriate axial stresses. 4. Results Experimental results for A, D, E, F are available (Bruce et al.,1993). In these physical tests, Weld A exhibited no signs of material rupture while Weld D showed some slight, localized stretching and thinning of the pipe wall under the fusion zone. Weld E, on the other hand, was a case of incipient failure where thinning and bulging of the wall was observed under the fusion Burnthrough prediction in pipeline welding 257 Figure 4. Weld A, ruptured volume shifted. zone with substantial material failure and growth of a centreline crack into the fusion zone (Bruce et al.,1993). As Weld E progressed, weld penetration increased, eventually melting some of the inner surface, but no leakage was observed. Weld F underwent complete failure as soon as the source arrived on the thinned section. The heat input in F was sufficient to completely melt through the pipe wall and a 1.6 mm diameter hole formed (Bruce et al.,1993). Numerous other welds described in (Bruce et al.,1993) were found to have centreline cracks associated with incipient burnthrough. Table 1 also shows the maximum temperatures on the inner surface, both measured and predicted by the FEA simulations. Measurement of peak temperatures on the inner surface of the vessel is a difficult task. Poor placement of the thermocouples can lead to erroneously low measured values as was noted in (Bruce et al.,1993). The moderate differences in heat input between A and E make the measured value for A difficult to credit. Furthermore comparisons with other measurements in (Bruce et al.,1993), for welds with the same heat input as A, show 520◦C to be far too low. The agreement for E and F is quite encouraging. The peak temperatures predicted for D and E are the same even though the heat inputs are not. Weld D had a smaller deposit thickness which raised the maximum temperature on the inside. In the mechanical portion of the FEA simulations, material rupture was predicted for all the pressurized vessels; i.e., A, D, E, F. Absolutely no material rupture was predicted for O, the unpressurized case. Ruptured elements are shown (Figures 4-7) shifted by an arbitrary amount for the purposes of visualization. In most of the cases examined, larger failed zones can be found at the ends of weld where the thermal histories are quite different from the history in the midlength. Weld A shows a trail of ruptured elements on the plane of symmetry just below the fusion zone; i.e., an intermittent centreline crack which, for the most part, does not reach either surface. Weld D shows massive material failure; along the weld centreline and within the body, only occasionally breaking the surface. Weld E which should have shown 258 A.S. Oddy and J.M.J. McDill Figure 5. Weld D, ruptured volume shifted. Figure 6. Weld E, ruptured volume shifted. fairly extensive material failure does show some, but only near the beginning of the thin patch. Once again, the ruptured material lies along the centreline, within the interior of the material. The increase in peak temperatures along the weld path noted in experiment E was directly connected to the material failure. It was not present in the FEA simulation. Weld F showed Burnthrough prediction in pipeline welding 259 Figure 7. Weld F, ruptured volume shifted. a concentrated zone of failed elements at the start of the thinnest region, penetrating from the inside surface to just past the mid-thickness. This does correspond with the location of burnthrough in experiment. The volumes of ruptured material can be ranked in decreasing order as D, A, E, F, O. This follows the order of internal pressures from largest to smallest. In examples such as weld D and A which have the same pressure, the case which has the higher maximum inside temperature shows more material rupture. 5. Discussion The agreement between the FEA simulations and experimental measurements is encouraging but could be better. Material failure from creep rupture was overestimated in A and D and underestimated in E. The instability expected in F from full penetration of the heat source through the pipe wall did not occur. However, predictions of volume of material rupture were generally conservative. The tendency to form centreline cracks did agree with reported results. The predicted volumes of ruptured material show a strong dependence on the internal pressure and, to a lesser extent, a dependence on the peak internal temperature. In the experiments (Bruce et al., 1993), weld F melted through the wall, forming a 1.6 mm diameter hole. The analysis on the other hand, showed material rupture where the hole was found but the expected instability did not occur. The modelled deposit height may have been too large or the penetration too small. The mesh may also be too coarse to permit the formation of a 1.6 mm diameter hole. The most likely problem is the zero-strength temperature. It probably should be lowered below the 1600◦C value which was used. The two-phase liquid/solid temperature is much lower than this, probably below 1500◦C. The zero-strength temperature in all probability should be even lower. The 1600◦ C isotherm in F only reached half way 260 A.S. Oddy and J.M.J. McDill through the pipe wall, evidently not enough to cause any instability. The same behaviour was seen in the spot-heating test where temperatures reached 1200◦C on the inner surface before any instability occurred. Welds A, D and E are instructive. Material rupture should not occur in A, be very small in D and show substantial stretching, bulging and extensive material rupture in E. The results for A and D show too much material rupture. E shows not enough material failure and not enough stretching and bulging. The material rupture properties seem to show too strong a dependence on stress and too small a dependence on temperature. The bulging and thinning which were missing in E suggest that creep properties underestimated the deformation at high temperatures. The material properties taken from (Harste et al.,1992) echo this; creep rate predictions are accurate to only ± 1 order of magnitude. The fact that O showed no material rupture whatsoever is very important. The properties chosen do not always predict material rupture, no matter what the heat input. The internal pressure is an important factor. In the thermal analysis there was little difference between welds D and E. The peak temperatures on the inner surface were indistinguishable, mainly because of the estimated deposit height. It is therefore, not surprising that a much larger volume of material ruptured in D than in E, given the higher pressure in the former. If the small difference in heat input between D and E truly represents the difference between safety and failure then the margin for error is extremely narrow. That is to say, underestimating the heat input by 11 percent is more dangerous than changing the internal pressure by 25 percent. The observed increase in weld penetration at incipient failure (Bruce et al.,1993) was not included in the analysis. The molten pool dimensions were assumed to remain constant. This increasing penetration was reported to be an important factor in weld E. Coupling the thermalmechanical analysis to a fluid dynamic prediction of circulation in the weld pool would be required to allow the molten pool dimensions to change during the analysis. In exploratory investigations there is a temptation to alter the material properties until the predictions agree with the measurements. This was not done, even though the material properties and basic boundary conditions seem to be the fundamental problem. First of all even very small changes in the deposit height can have very profound effects on temperatures on the inner surface. More importantly, the melting or zero-strength temperature is probably too high, causing the instability in weld F to be missed. The temperature dependence of the creep rupture properties was estimated in a very crude fashion and must be a major source of error. The estimation of the creep strain rate properties might be improved but given the known variability in such properties it is unlikely to be possible. In the final analysis rupture predictions were conservative in that some degree of material rupture was predicted in all cases where it was found experimentally and in several cases where it was not. The amount may have been underestimated but this is only to be expected given the difficulties with material properties. The known variability in the material response makes better agreement unlikely. 6. Conclusions and Contributions FEA simulations of welding on pressurized vessels have been performed and compared to experiment. The agreement between the thermal-mechanical FEA and the physical tests was encouraging. The material failure was found to occur in the appropriate locations although the FEA predictions tended to predicted material rupture even in cases where it was not found. 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