Contents lists available at ScienceDirect

Journal of Materials Processing Tech.

journal homepage: www.elsevier.com/locate/jmatprotec

Vacuum laser welding of SA508 steel

J.A. Francisa,⁎, N. Holtumb, S. Olschokb, M.J. Roya, A.N. Vasileioua,c, S. Jakobsb, U. Reisgenb,M.C. Smitha

a School of Mechanical, Aerospace and Civil Engineering, The University of Manchester, Manchester, M13 9PL, UKb RWTH Aachen University, Welding and Joining Institute (ISF), Pontstraße 49, 52062, Aachen, Germanyc Dalton Nuclear Institute, The University of Manchester, Manchester, M13 9PL, UK

A R T I C L E I N F O

Associate Editor: C.H. Caceres

Keywords:Contour methodDuctile-to-brittle transitionNuclear fabricationPressuriserReactor pressure vesselWeld toughness

A B S T R A C T

Vacuum laser welding was employed to manufacture 80mm thick welds in SA508 Grade 3 steel in two weldpasses, using a 16 kW laser, while travelling at 150mm/min. The motivation was to explore the potential for theapplication of the process to the joining of large, safety-critical nuclear components, such as the steam generatorsor the pressuriser in a pressurised water reactor (PWR). The advantages of vacuum laser welding are first re-viewed, and compared to those of electron beam welding, in terms of the process physics. Preliminary devel-opment work is then summarised, together with an evaluation of weld quality, mechanical properties and re-sidual stresses. Vacuum laser welding warrants further development, as it offers significant promise for futurenuclear build programmes.

1. Introduction and background

Safety-critical nuclear components such as the reactor pressurevessel, the steam generators and the pressuriser within a pressurisedwater reactor (PWR) are currently fabricated using narrow-gap variantsof traditional welding technologies, such as submerged-arc welding(SAW) or gas-tungsten arc welding (GTAW). Indeed, codes of practicesuch as those that are overseen by the American Society of MechanicalEngineers (ASME) currently exclude the application of other weldingprocesses. Changes to fabrication codes for the nuclear sector must beunderpinned by extensive development work and testing, but the ap-petite for the requisite development programmes diminished in thewake of the Three Mile Island accident in 1979 and the Chernobyldisaster in 1986. As a consequence, the following two decades sawdevelopments in relevant welding technologies occurring only at themargins.

Comparatively recently, Sanderson et al. (2000) undertook devel-opment work on the electron beam (EB) welding process to make itmore amenable to the manufacture of large nuclear pressure vessels.These workers pointed out that it is possible to avoid placing a largepressure vessel inside a vacuum chamber if welding is carried out at areduced pressure, rather than in a high vacuum, since localized sealingsystems can cope with such scenarios. Ayres et al. (2010) subsequentlydemonstrated that reduced pressure electron beam (RPEB) weldingcould be applied to the welding of a pressure vessel steel with a

thickness of 160mm. These developments stemmed from the need toreduce the costs and lead times associated with plant construction, sothat nuclear power can remain competitive with alternative forms ofgeneration. Indeed, the EB welding process offers the ability to makesingle-pass welds at thicknesses exceeding 100mm (in contrast to∼100 passes for GTAW or SAW) and, therefore, it can deliver a step-change in productivity.

Jeyaganesh et al. (2014) recently provided an overview of theNNUMAN research programme in nuclear manufacturing. In this pro-gramme, significant efforts were devoted to evaluating welding tech-nologies for safety-critical pressure vessels. For example, Feng et al.(2017) reported on the development of multipass narrow gap laserwelding for nuclear pressure vessels, while Balakrishnan et al. (2018)described a cross-process comparative study on residual stress devel-opment in 30mm thick welds in SA508 Gr. 3 Cl. 1 steel. Furthermore,Rathod et al. (2019) described the manufacture of 130mm thick weldsin detail, using technologies (GTAW and SAW) that are currently em-ployed in nuclear construction, as well as a candidate process for futurebuild programmes (EB welding). A similar cross-process comparison onthese 130mm thick welds was then described by Vasileiou et al. (2019).However, in contrast to EB welding, laser welding has received com-paratively little attention, particularly with respect to the welding ofcomponents with thicknesses exceeding 50mm. As such, the authorsformed the view that laser welding warranted further investigation.

In a noteworthy study, Zhang et al. (2011) applied multipass laser

https://doi.org/10.1016/j.jmatprotec.2019.116269Received 14 December 2018; Received in revised form 10 May 2019; Accepted 16 June 2019

⁎ Corresponding author at: School of Mechanical, Aerospace and Civil Engineering, The University of Manchester, Manchester, M13 9PL, UK.E-mail address: john.francis@manchester.ac.uk (J.A. Francis).

Journal of Materials Processing Tech. 274 (2019) 116269

Available online 18 June 20190924-0136/ © 2019 The Authors. Published by Elsevier B.V. This is an open access article under the CC BY license (http://creativecommons.org/licenses/BY/4.0/).

T

welding to 316 L austenitic stainless steel plates with a thickness of50 mm, by depositing weld metal into a narrow groove while em-ploying the heat source in conduction mode. Approaches of this typemust overcome several challenges, such as the need to deliver a focusedlaser beam and shielding gas to the bottom of a deep and narrowgroove, while concurrently ensuring that the heat source and filler wireare precisely aligned. These challenges are easier to meet if the grooveis widened. On the other hand, it is also necessary to ensure that thewalls of the weld groove melt to the extent that is necessary to avoidlack of side-wall fusion; a defect which is not tolerated at any level insafety critical pressure vessels. The avoidance of lack-of-fusion defects,and demands associated with welding productivity, are favoured bybringing the opposing walls of the groove closer together. There is,therefore, an inescapable tension between the various challenges formulti-pass narrow gap laser welding that, in the authors’ view, rendersthe process impractical for the welding of very thick sections.

In the past decade, Jakobs and Reisgen (2015) demonstrated thatlaser welding in a vacuum (LaVa) achieves comparable weld penetra-tion depths to the EB process when the same beam power and weldingspeed are used. Laser beams offer some advantages over electronbeams, in that they are not susceptible to deflection by the residualmagnetism that invariably exists in thick sections of steel, and theshielding requirements are also less daunting, since laser welding doesnot generate X-rays. Limitations of the process are associated with theavailable laser power (typically much lower than for EB welding) andthe fact that it is more challenging to oscillate a laser beam than anelectron beam. However, laser power sources are rapidly increasing inpower, and the limitations on laser beam scanning are receding owingto advances in the performance of guiding mirrors. As such, there arestrong motivations for exploring the potential of LaVa for nuclear fab-rication.

This paper describes an attempt to carry out vacuum laser weldingon thick sections of nuclear steel. It begins with a comparison of thephysical aspects of the EB and vacuum laser welding processes, beforedescribing the process of identifying suitable welding parameters, andreporting on the assessment of weld quality through distortion mea-surements, residual stress determination, metallography, tensile testingand, finally, Charpy impact testing. It is hoped that this article willreveal the promise that this technology holds for future pressure vesselfabrication, and that it will stimulate further developmental work.

2. Physical considerations

2.1. Electron beam welding and laser beam welding

Despite differences between the beam-material interactions, EBwelding and laser beam (LB) welding have many similarities. For ex-ample, both processes employ a focused, high-intensity beam to im-mediately evaporate the base material and form a vapor capillary(keyhole) that allows deep penetration welding. By moving this capil-lary along a weld preparation, it is possible to join thick walled com-ponents. Using power beam welding, a high aspect ratio (bead depth/bead width) can be achieved.

The high-intensity beam in EB welding consists of electrons gener-ated at a heated cathode, which are accelerated by high voltages, andfocused by an array of coils in a beam generator. During welding, theseelectrons hit the surface of the work piece while travelling at up to 70%of the speed of light, subjecting the work piece to power densities thatapproach 107 W/cm². In contrast, the LB process uses a focused beam ofelectromagnetic waves in form of light with a defined wavelength totransfer energy to the work piece.

Generally, the wavelengths of modern laser beam sources that areused for materials processing are in the near infrared region. Due to thelow absorption rates of common construction materials, only a smallfraction of the beam power is absorbed by the work piece in a singlebeam-material interaction (e.g. in heating or conduction-mode

welding). Depending on the material, the reflection rates are between70% (iron/steel) and 95% (aluminum). As with an EB, an LB forms avapor capillary when the power density on the work piece surfacereaches a threshold. Energy losses from reflection reduce to valuesbelow 10%. Multi-mode solid-state lasers allow focal diameters as smallas 0.1mm, and power densities of more than 107 W/cm² can beachieved. Single-mode fiber lasers offer a beam profile near to the idealGaussian distribution with focal diameters down to 0.03mm. Althoughgenerally available with lower output powers, these modern laser typescan achieve power densities up to and beyond 108 W/cm².Conventional LB welding at atmospheric pressure uses a shielding gasto protect the molten pool from oxidation and to reduce the inherenttendency for spatter.

Despite the similarities between the processes, when welding atatmospheric pressure, the LB process is not able to match the perfor-mance of the EB process in terms of achievable weld penetration andweld seam quality at comparable beam power levels. To overcome thisdrawback, much effort has been devoted to the development of laserbeam welding in vacuum.

2.2. Laser beam welding in a vacuum

The impetus for the development of laser beam welding in a vacuumis based on isolated research work carried out by Arata et al. (1985).This study investigated the extent to which the harmful plasma plumeabove the vapor capillary can be suppressed by reducing the ambientpressure when welding with a CO2 laser. The benefits associated with areduction in pressure were proven and, as a side finding, an increase inachievable weld penetration depths was observed.

Motivated by the earlier finding, work in this field started again in2009 using modern solid-state laser beam generators (single-mode fiberlasers). Welding in a vacuum showed both a reduction of the vaporplume above the capillary and a clear increase in the weld penetrationdepth when using comparable parameters. While welding at atmo-spheric pressure, it was only possible to operate at the transition be-tween conduction-mode welding and keyhole welding, but a reductionin the ambient pressure to 10mbar led to the formation of a pro-nounced vapor capillary and the quadrupling of the weld penetrationdepth, as has been described by Holtum and Jakobs (2018). The sig-nificant increase in achievable weld penetration depths is caused by acombination of effects resulting from welding at reduced pressure. Ashas been pointed out by Luo et al. (2015), the first and most visuallystriking effect of these may be the reduction of the bright metal vaporplume above the keyhole. Other benefits have also been reported forspecific materials. Teichmann et al. (2018), for example, have reportedthat reducing the ambient pressure during laser welding of aluminiumcastings can lead to significant reductions in porosity and corre-sponding improvements in mechanical properties.

An interesting point arises in that laser radiation that is generatedby commonly-used solid-state lasers (which typically have wavelengthsin the 1 μm range) does not tend to be absorbed by metal vapor to thesame extent as radiation from CO2 lasers (which have wavelengths inthe vicinity of 10.6 μm). Clearly, the suppression of the plasma plume inisolation cannot explain the gain in weld penetration depth. A far morepotent influence of the reduction in pressure lies in the change thatoccurs to the boiling point of the base material. With a simple reductionof the working pressure to 10−1 mbar, the boiling temperature ofmolten iron is reduced by approximately 1300 °C, while the solidifica-tion temperature is largely unaffected, as is evident from the review byHonig and Kramer (1969).

The effects of such a reduction in pressure on the boiling point, andthe corresponding reduction in the temperature range for the liquidstate, lead to several effects on the laser welding process. Firstly, theenergy that is needed to transform the molten metal to vapour, whichcreates and stabilises the keyhole, is significantly reduced. Secondly,even the earliest trials performed by RWTH Aachen using a laser power

J.A. Francis, et al. Journal of Materials Processing Tech. 274 (2019) 116269

2

of 600W suggested that the threshold power density that is needed toform a keyhole is reduced by the lower boiling point. Thirdly, thecombination of the lower boiling point and an unchanged melting pointwill lead to a lower temperature at the interface between the keyholeand the surrounding molten material, when comparing welding in avacuum to welding at ambient pressure. When the fact that the thermalconductivity and the solidification temperature are largely independentof pressure is taken into account, this suggests that the layer of moltenmetal surrounding the vapour capillary will be thinner when welding ata reduced pressure. The combined influence of a thinner layer of moltenmetal and an increase in the size of the vapour capillary leads to sta-bilisation of the keyhole and produces deep and narrow weld seamgeometries.

A final point of comparison is that, unlike electrons, photons do notcarry charge. Therefore, photons are less likely to interact with atoms inthe atmosphere. As a result, when applying laser beam welding undervacuum it is only necessary to decrease the pressure to ∼ 10−1 mbar inorder to achieve results that compare well with EB at working pressuresof 10−3 mbar or better. That is to say, laser welding is less sensitive tothe quality of the vacuum, as was pointed out by Jiang et al. (2017) intheir review article.

3. Experiments

3.1. Materials

The objective of the study was to make two vacuum laser welds inSA508 Grade 3 Class 1 steel at a thickness of 80mm. This thickness waschosen on the basis that it had already been used by Rathod et al.(2017) for investigations on GTAW and SAW joints in the NNUMANresearch programme. A thickness of 80mm was also thought to besufficient to demonstrate the viability of vacuum laser welding forjoining thick sections of pressure vessel steel, while not being so thick asto reduce the likelihood of making successful welds within the time andresource limitations that were applicable. To this end, blocks measuring290mm×160mm×80mm were machined from a larger forging,which also acted as the source of material for 130mm thick welds madeusing the GTAW, SAW and EB welding processes. The details of these130mm thick welds and the processing history of the forging materialhave been described by Rathod et al. (2019). The chemical compositionof the parent steel, as quoted on the mill certificate, is given in Table 1,while the quoted room-temperature mechanical properties include a0.2% proof stress of 460MPa, an ultimate tensile strength (UTS) of610MPa, and an elongation of 26%. The final dimensions of the va-cuum laser welded test pieces in SA508 Gr. 3 Cl. 1 steel after comple-tion of the joints were 290mm×320mm×80mm.

Prior to making the welds, a number of trials took place on othersteels in order to identify suitable welding parameters. Early bead-on-plate trials were carried out on a 50mm thick plate of S355 steel inorder to establish a suitable welding speed. Further bead-on-plate trialsthen took place on a 65mm thick plate of S690QL steel with a view tofinalising the parameters for beam oscillation. Nominal chemicalcompositions for S355 steel and S690QL steel are also included inTable 1. Once suitable welding parameters were identified, a trial weldwas attempted between two plates of A533 steel, each measuring285mm×100mm×75mm. This steel was used for a full weldingtrial because it has the same nominal chemical composition as SA508

Gr. 3 steel. The final dimensions of the welded test piece in A533 steelwere 285mm×200mm×75mm.

3.2. Welding set-up

A Trumpf disc laser with a maximum power of 16 kW was used in alltrials and experiments. The laser had a beam parameter product of8mm*mrad, and was fed through a fiber optic cable to the laser opticsassembly, which was hard mounted to the wall of the vacuum chamber,as shown in Fig. 1. The optic that was used had an image ratio of 1:2.75with a focal length of 932mm. The main components of the optical unitused here were manufactured by Kugler, while the oscillating mirrorwas manufactured by the company ILV, and has been described by Arlt(2012). A fiber with a 200 μm diameter was used. In combination withthe optic the spot size at the focal point was 550 μm in diameter.

In the first instance, the laser beam coming from the fiber must beconverted back to a parallel raw beam. The laser light comes out of thelaser light cable with a total aperture angle of 0.2° and is converted intoparallel light beams by the collimator. The collimator unit consists of ametal mirror with a diameter of 35mm and a focal length of 125mm.The laser beam then hits a plane mirror, which deflects the beam by90°. Instead of a conventional deflecting mirror, an ILV DC oscillationoptic can also be used. This so-called scanner optic consists of a planemirror that, in the switched-off state, acts as a deflecting mirror anddeflects the beam by 90°. Its frequency range is between 3 and 1000 Hz.The oscillation optic control unit, which is manually operated, controlsand monitors the scanner mirror. The scanner mirror is driven by agalvo scanner whose control signal can be a sine wave, triangular sawtooth wave or a pulse signal. In addition to the amplitude, a positive ornegative offset can also be applied to the control signal of the mirror.

The vacuum chamber was cylindrical, and was mounted on rollers,so that the welding position (e.g. 1G (beam vertical) or 2G (beamhorizontal)) could be changed with ease. Prior to welding, a dome-shaped lid was clamped in position and a vacuum pump was used toreduce the pressure inside the chamber to< 10mbar. In this work, thepressure in the chamber was set to be below 1×10−1 mbar in all cases.A motorised x–y table was mounted inside the vacuum chamber toenable the test piece to translate during the welding process. Themaximum possible traverse distance was 300mm, and this was thelimiting factor regarding the total length of the weld seams in this work.When a welded joint was made, double-sided welding was employed,and the welding direction was kept consistent on either side of the testpiece (see Fig. 1).

3.3. Welding procedure development

It was clear from the outset that, with a maximum available powerof 16 kW, it would be necessary to complete an 80mm thick weld intwo passes. Since the weld pool would not penetrate through the fullthickness in a single pass, a decision was taken to carry out welding inthe 1G position, i.e. with the beam vertical, as there was no risk of theweld pool falling through during welding. Thus, the objective for thefirst set of bead-on-plate trials was to establish a suitable welding speed.A higher welding speed was assumed to be desirable from a pro-ductivity standpoint. However, the choice of welding speed was con-strained by the requirement to achieve a penetration of at least 45mmin each pass, to ensure that complete fusion would be achieved in a two-

Table 1Chemical compositions (wt.-%) for SA508 Gr. 3 Cl. 1, S355 and S690QL steels. An asterisk indicates that the values given are nominal values.

Element C Si Mn Cr Co Ni Mo S P Cu Fe

SA508 Gr.3 Cl.1 0.16 0.27 1.43 0.23 0.004 0.77 0.52 0.002 0.005 0.04 Bal.S355 Steel* 0.2 0.55 1.6 – – – – 0.030 0.030 0.55 Bal.S690QL Steel* 0.2 0.8 1.7 1.5 – 2.0 0.7 0.015 0.025 0.5 Bal.

J.A. Francis, et al. Journal of Materials Processing Tech. 274 (2019) 116269

3

pass weld.In the first set of three trials, a 50mm thick plate of S355 steel was

used and the welding speed was set to values of 100, 150 and 200mm/min. In each case, the focal position of the beam was 5mm below thetop surface of the plates (z =−5mm), and the beam was oscillated at afrequency of 50 Hz using a sinusoidal waveform. The oscillation width(i.e. the distance between the extremities) was 4.5 mm, as measuredusing a pilot laser to mark a thin steel sheet that was coated with en-amel. The trial at a speed of 100mm/min resulted in the weld poolfalling through the 50mm thick plate, while the trial at 150mm/minresulted in a penetration depth of ∼ 45mm, and the trial at 200mm/min resulted in a penetration depth of ∼ 30–35mm. On this basis, atravel speed of 150mm/min was identified as the most suitable.However, in these trials, a spiking defect was observed in the beadprofile near the root (Fig. 2a). The mechanisms of spiking were in-vestigated by Wei et al. (2012). Defects of this type are undesirablebecause they increase the likelihood of problems associated with por-osity and cracking. In this case, the authors were of the view that thespiking was a consequence of the sinusoidal oscillation pattern leadingto an excessive fraction of the time being spent with the beam directedat the extremities of oscillation.

A second series of bead-on-plate trials was undertaken on a 50mmthick plate of S355 steel, employing a welding speed of 150mm/minand a triangular oscillation waveform. The objective with these trialswas to establish whether the spiking defect could be avoided throughthe choice of waveform. The oscillation width was unchanged from thefirst set of trails, while values for the oscillation frequency of 50, 100and 200 Hz were used. Spiking was observed to persist with a frequencyof 50 Hz, but it was not seen when oscillating at higher frequencies.However, as the frequency increased, the weld bead profile becamedeeper and narrower, as shown in Fig. 2b. In this image, the cavity atthe top of the weld resulted from the beam penetrating through the50mm thick plate and the weld pool partly falling through under theinfluence of gravity. (This type of defect can be mitigated by welding inthe 2G position.) Although a high oscillation frequency offered an in-crease in penetration, the increase would not have been sufficient tojoin 80mm thick plates in a single pass, so the benefits associated withthe higher penetration were not significant. Furthermore, the authorsopted to avoid bead profiles that were very narrow at the root, on the

basis that beads that were somewhat wider would minimize the po-tential for incomplete fusion if there happened to be any misalignmentbetween the passes made on opposing sides of the plate. Consequently,an oscillation frequency of 100 Hz was selected.

A final set of parameter trials was carried out on a 65mm thick plateof S690QL steel to establish whether there were any benefits associatedwith changing the focal position so that it was aligned with the topsurface of the plates (i.e. to z= 0) and also to establish whether, byreducing the oscillation width, it would have been possible to increasethe penetration depth to the point where the welding speed could havebeen increased. The matrix for the final set of parameter trials is listedin Table 2. In these trials it appeared that the probability of defects, orspikes at the weld root, was reduced when a focal position coincidingwith the surface (z= 0) was employed. However, reductions in theoscillation width tended to produce welds that were narrow, with morepointed profiles at the weld root and, in some cases, there appeared tobe a region coinciding with the centerline of the weld bead in which themicrostructure was different (see Fig. 2c). The authors opted to main-tain the oscillation width at 4.5mm in order to maintain a roundedprofile at the root of the weld bead, and to minimize the extent to whicha distinct microstructural band formed on the weld centerline. Fol-lowing this rationale, condition number 6 in Table 2 was chosen as thefinal set of welding parameters, and the macrograph from this bead-on-plate trial appears in Fig. 2d.

When the trial weld and final welds were made, the preheat tem-perature was between 100 and 105 °C. Preheating was achieved byplacing the component in a furnace overnight at a temperature20–30 °C above the desired preheat temperature. Prior to welding, thetest piece was transferred to the vacuum chamber, and the chamber wasevacuated to a pressure not exceeding 1× 10−1 mbar. Welding tookplace when thermocouples that were attached to the test piece in-dicated that the desired preheat temperature had been reached. Theinterpass temperature was identical to the preheat temperature, andthis was also controlled by monitoring the output from thermocouples.

3.4. Details of instrumentation

Eight k-type thermocouples were attached to the trial weld and toboth of the welds that were made in SA508 Grade 3 Class 1 Steel. In all

Fig. 1. The set-up for the vacuum laser welding experiments. During welding, the laser assembly remained stationary and the test piece translated on a supportingtable.

J.A. Francis, et al. Journal of Materials Processing Tech. 274 (2019) 116269

4

cases, temperatures were recorded at a frequency of 10 Hz. Four ther-mocouples were attached to the top surface of the welds and four to thebottom face. The thermocouples were arranged symmetrically aboutthe weld centerline to ensure that there was redundancy at each loca-tion. The configuration of the thermocouples is shown schematically inFig. 3.

3.5. Distortion measurements

A Creaform HandySCAN 3D™ laser scanning system (Handyscan700 model) was used to measure the distortion associated with eachpass in each welded joint. The surface profiles and distortion werecaptured in three dimensions using this hand-held system, which had amaximum resolution of 0.05mm and a maximum accuracy of 0.03mm.The measured profiles were used to calculate a butterfly distortionangle, θ. If the weld test piece were to remain perfectly flat upon thecompletion of welding, the value of θ would be zero, indicating nobutterfly distortion.

Distortion measurements took place at three instances: (a) beforewelding, when the sample was in its initial state, (b) after the first weldpass and (c) after the second weld pass. Reflective targets were placedon the specimen at strategic locations, to allow rapid data acquisitionduring scanning. The shutter time was set at 4.7 ms and the volumetricaccuracy was set at 0.020mm/m. The point cloud for the weldment was

acquired using VX elements® version 6.1 SR1 software, which was thenconverted to a stereolithographic file and processed using GeomagicControl 2014.0®.

3.6. Nondestructive evaluation

Each weld was subject to radiographic examination and, in all cases,the acceptance criteria published by the American Society ofMechanical Engineers (2017) were met. However, it should be men-tioned that, while this result was encouraging, radiography is onlylikely to detect volumetric flaws. Ultrasonic testing would also be ne-cessary as a more effective check for planar defects such as lack-of-fusion. Macrographs were also prepared for both the trial weld and anSA508 steel weld.

3.7. Residual stress measurements

The contour method was applied to an 80mm thick SA508 testpiece in the as-welded condition, to establish whether the vacuum laserwelding process behaved in a similar manner to the EB welding process.Kundu et al. (2013) were among the first researchers to demonstratethat the contour method can capture the “M-shaped” residual stressdistribution that is expected for EB welds in P91 steel, for which solid-state phase transformations are expected to influence the development

Fig. 2. Macrographs prepared from bead-on-plate trials for: (a) S355 steel travelling at200mm/min with sinusoidal oscillation, re-vealing spiking defects; (b) S355 steel travel-ling at 150mm/min with triangular oscillationat 200 Hz; (c) 65mm thick S690QL steel,showing a pointed weld root profile and abanded centerline microstructure; and (d) thetrial parameters that were selected for welding.

Table 2Nominal Welding Parameters for the Bead-On-Plate Trials on 65mm thick S690QL Steel.

Condition Number Power (kW) Speed (mm/min) Focal Position (mm) Waveform Shape Frequency (Hz) Oscillation Width (mm)

1 16 150 −5 Triangle 100 4.52 16 200 −5 Triangle 100 33 16 200 −5 Triangle 100 1.54 16 150 −5 Triangle 100 35 16 150 −5 Triangle 100 1.56 16 150 0 Triangle 100 4.57 16 200 0 Triangle 100 38 16 200 0 Triangle 100 1.59 16 150 0 Triangle 100 310 16 150 0 Triangle 100 1.5

J.A. Francis, et al. Journal of Materials Processing Tech. 274 (2019) 116269

5

of stress. Smith et al. (2014) subsequently demonstrated, using acombination of standard and incremental deep hole drilling, that the M-shaped nature of the stress distribution was also a feature of EB welds inSA508 Gr. 3 Cl. 1 steel at the much greater thickness of 160mm. Thesestudies provide useful background information, since one might an-ticipate that vacuum laser welding in keyhole mode would producesimilar stress distributions to those that are generated by EB welding.

In this work, pilot holes 3mm in diameter were machined throughthe thickness of the test piece. These holes were 25mm from the outeredges of the welded plate, and half way along the length in the directionof welding. The pilot holes allowed a transverse contour cut to takeplace across the majority of the cross-section, so that the distribution oflongitudinal stresses could be evaluated. The ligaments of material thatremained at either edge would have prevented translation of the testpiece, on either side of the contour cutting line, during cutting.

The contour method measurements involved the following steps:

• Making the contour cut;

• Measuring out-of-plane deformations on the newly created surface;

• Processing of the surface profile data (averaging, fitting andsmoothing); and

• Estimating the original residual stress distribution by conducting anelastic finite element analysis.

The contour cut was made by electric discharge machining (EDM),with the test piece left un-restrained during cutting. An GF AgieCharmilles FI 440 ccS wire-EDM machine was employed to cut fromone pilot hole to the other, using parameters that were optimised overseveral earlier trials.

Surface profilometry was carried out with a Mitutoyo Crysta-Apex776 coordinate measurement machine fitted with a 2mm diameter balloperating in hen-peck mode. The maximum distance between pointswas 2.5mm, resulting in 6500 points describing each side of the cut.The data processing involves several steps, as has been described byJohnson (2008). Sun et al. (2017) demonstrated that due considerationneeds to be given to errors arising from cutting-induced plasticity whenselecting a cutting strategy and when subsequently processing the data.In this work, data from each side of the cut was first processed togenerate an outline, and to eliminate data outside of the focal range ofthe scanner. The profiles for the two cut surfaces were then averaged toobtain a single surface contour. This surface contour underwent asmoothing operation that involved fitting a bi-variate cubic spline tothe profile, using a least-squares approach. The spline comprised seg-ments of 3rd order polynomials, and these were joined at locations that

are referred to as ‘knots’. The spacing of these knots can be chosenbased on a judgement as to what gives the best compromise betweenachieving a good fit to the data, while also avoiding the possibility ofoverfitting, as was pointed out by Prime et al. (2004). A knot spacing of0.5 mm was chosen by the authors after some trials were carried out.

An elastic finite element analysis (FEA) was carried out withABAQUS, employing 120,000 s-order tetrahedral (C3D10) elements.The smoothed surface profile was imported and used as a basis forgenerating displacement boundary conditions. Further conditions wereapplied to preclude the possibility of rigid body motion. The analysiswas purely elastic, since the contour method was originally developedby Prime (2001) based on the principle of elastic superposition. Themodulus and Poisson’s ratio were assumed to be 210 GPa and 0.301,respectively. The contour method has been applied to welded joints andto other applications with considerable success over the past decade orso but, in common with other techniques that work on the principle ofelastic stress relaxation, it is prone to errors associated with cutting-induced plasticity. The analysis demonstrated that there were wireentry artefacts, particularly at the edges of the pilot holes, but this didnot affect the estimates for the stresses arising as a consequence ofwelding.

3.8. Post-weld heat treatment

After the completion of the contour method measurements, the re-maining intact weld and one half of the weld that had been subject tocontour cutting were subjected to PWHT. Both of these pieces were heattreated concurrently in the same furnace. The remaining half of the testpiece that was subject to contour cutting was left in the as-weldedcondition for archiving purposes.

At temperatures below 300 °C there were no restrictions on theheating and cooling rates. Above this threshold the maximum permis-sible heating or cooling rate was 20 °C per hour. The hold took place ata temperature of 607 +/− 13 °C, which was maintained for 4 h. Theheating and cooling rates were lower than is necessary to comply withguidelines published by the American Society of Mechanical Engineers(2017), and they were chosen in order to maintain consistency with thewelds described by Rathod et al. (2019), which were made using otherwelding processes on the same steel, in order to assist with bench-marking of the vacuum laser welding process.

3.9. Metallography and hardness testing

The preparation of macrosections is a destructive examination

Fig. 3. Details of thermocouple arrays for the welded samples. The same arrays were used on both sides of the test piece. Thermocouples were located at distances of12mm and 20mm from the weld seam.

J.A. Francis, et al. Journal of Materials Processing Tech. 274 (2019) 116269

6

method. Transverse slices were sawn from the weld test pieces, milled,ground and then etched. During mechanical separation, care was takento ensure that heating of the samples was minimized, though the use ofcoolant, in order to prevent changes in microstructure. Etching wasthen undertaken with 10% nitric acid (i.e. 10% Nital) to reveal thefusion boundary profiles.

Hardness testing was carried out after the slices had been polishedto a 1 μm surface finish, but with the material in an un-etched condi-tion. A Vickers microhardness testing machine (Durascan™) was em-ployed to make the measurements, using a load of 1 kg and a dwellperiod of 20 s. The rectangular measurement array covered a regionthat included the fusion zone, the heat-affected zone (HAZ) and suffi-cient parent material for the hardness values to stabilise. The intervalsbetween measurements varied according to the microstructural gra-dients that were expected, with measurement intervals between 0.25and 1mm being typical.

3.10. Mechanical testing

In order to assess the quality of the vacuum laser welds, cross-weldtensile testing, bend testing and Charpy impact testing were carried out.The extraction of specimens involved machining twelve slices of ma-terial, each with a thickness of 12mm, from the weld test pieces, andone 18mm thick slice. After the contour method measurement wascompleted, only one of the test pieces in SA508 steel remained intact.This weld provided ten of the 12mm thick slices and the 18mm thickslice, as shown in Fig. 4. The remaining two 12mm thick slices wereextracted from the remnant of the second SA508 steel weld after it hadbeen subjected to PWHT.

Three cross-weld tensile blanks each measuring18mm×18mm×125mm were extracted from the 18mm thick slice.This slice of material coincided with the weld mid-length position(Fig. 4). The blanks were extracted in such a way that the fusion zone ofthe weld was at the mid-length position of the blank. The top andbottom 10mm of the slice/weld were avoided. Each blank was thenmachined to a round cross-section with a nominal gauge diameter of12.7 mm and a parallel length of 72mm. The specimens were testedaccording to the ASTM A370 standard (2017), using a universal screwdriven machine with self-aligning fixtures. All tests took place at am-bient temperature and a single machine displacement rate was usedthroughout.

Two of the 12mm thick slices were used as full wall thickness bendspecimens. Bend testing was carried out following the guidance given inSection IX of the Boiler and Pressure Vessel Code issued by The

American Society of Mechanical Engineers (2017), using a formerdiameter of 40mm.

Ten of the 12mm thick slices were used to prepare standard sizedCharpy impact specimens. Three coupons were extracted from eachslice, at locations A–C (Fig. 4), and the top and bottom 15mm of theweld were avoided. The location from which each coupon was ex-tracted was noted so that it would be possible to check for through-thickness variability in the results. The notches of the specimens coin-cided with the heat-affected zone (HAZ), and they were nominally1mm from the fusion boundary, as determined by metallography andetching of the slice. The notch orientation was such that the direction ofcrack propagation was parallel to the welding direction. Tests werecarried out at ten different temperatures, with the lowest being−196 °C and the highest being 200 °C. Three tests were carried out ateach of the chosen test temperatures, meaning that thirty tests werecarried out in total.

4. Results

4.1. Temperature measurements

The temperature transients that were measured at each thermo-couple location are shown in Fig. 5, for both the first pass (a) andsecond pass (b) in the double-sided weld. It can be seen that the ther-mocouples that were located on the same side of the plates as the heatsource, and at a distance of 12mm from the weld centerline, recordedpeak temperatures of approximately 550 °C, while those at a distance of20mm recorded peak temperatures in the range between 350 and400 °C. In contrast, the peak temperatures that were recorded on theopposite side of the blocks to the heat source were close to 200 °C, ir-respective of the distance from the weld centerline. One of the en-couraging features of these results is that a high degree of symmetry isobserved between thermocouples that are equidistant from (and onopposite sides of) the weld centerline. This suggests that the measuredthermal transients are credible.

4.2. Distortion measurements

Fig. 6 provides examples of the output from the three-dimensionallaser scanning system. Three-dimensional images that were extractedfrom one of the SA508 steel welds are shown in Figs. 6a and b, while avirtual cross-section from the same dataset is shown in Fig. 6c. Theacquisition of data in this manner enabled the authors to track theevolution of butterfly distortion, which can be quantified by the angle,

Fig. 4. Schematic representation of positions at which tensile specimens, bend specimens and Charpy impact specimens were extracted from a vacuum laser weld inSA508 Gr. 3 Cl. 1 steel after PWHT.

J.A. Francis, et al. Journal of Materials Processing Tech. 274 (2019) 116269

7

θ, in Fig. 6c. If the plates do not undergo any rotational movement as aconsequence of the welding process, then the value of θ will be zero. Itcan be seen in Fig. 6d that, prior to welding, the measured value of θ isclose to zero, as one might expect. After the completion of the first weldpass, the value of θ increased to approximately 0.5°, a value that isobservable with the naked eye but also one that is certainly not ex-cessive. However, after the completion of the weld pass from the op-posing side (i.e. after pass number 2) the value of θ is once again veryclose to zero, which indicates that the two-pass, two-sided weldingstrategy that was employed in these experiments would lead to negli-gible butterfly distortion.

4.3. Residual stresses

The longitudinal residual stresses for a weld in SA508 Gr. 3. Cl. 1steel, as measured by the contour method, are shown in Fig. 7 for a testpiece in the as-welded condition. In this case, the stress distribution isshown for the case where the second weld pass appears on the top of thesample, which is consistent with the presentation in Figs. 1 and 4. Thecutting direction during the measurement process was from right to left.The locations at which pilot holes were machined are highlighted.

Smith et al. (2014) and Balakrishnan et al. (2018) found that thepeak values of tensile residual stress, for fusion welds in SA508 Grade 3steel, corresponded approximately with the ultimate tensile strength of

the material, which was measured by Rathod et al. (2019) to be610MPa. This is broadly supported by the data that are presented inFig. 7, although there are some small regions in the figure where themeasured stresses approach 700MPa. The distribution of tensile regionsis also in good agreement with the results published by Smith et al.(2014) and Balakrishnan et al. (2018) for electron beam welding, whenan account is given for the fact that the vacuum laser welds were madein two weld passes, rather than one. The characteristic M-shaped re-sidual stress distribution that is seen in EB welding is evident for thesecond weld pass in particular, but it is also evident for the first pass. Intheir review, Francis et al. (2007) described how this distribution arisesas a consequence of solid-state phase transformations taking placewithin the HAZ and fusion zone during a welding thermal cycle,whereby the decomposition of austenite on cooling results in transfor-mation strains that counteract the accumulation of thermal contraction.As a consequence, the regions that are subject to the highest tensilestresses are located immediately outside the outer boundary of the HAZ,since material in this region has experienced maximum thermal con-traction upon cooling without benefiting from compensatory transfor-mation strains. It can be seen that anomalous compressive line-shapedregions of residual stress appear in the vicinity of the pilot holes. Theseare believed to be artefacts introduced by cutting-induced plasticity inthe vicinity of the pilot holes.

Fig. 5. Temperature transients as measured by thermocouples in the first pass (a) and second pass (b) of the welding process. The highest temperatures were recordedby thermocouples 2 and 3 during the first pass, while for the second pass the highest temperatures were recorded by thermocouples 6 and 7.

Fig. 6. (a) and (b) – Three-dimensional surface profiles of a vacuum laser weld; (c) a virtual cross-section extracted from the same dataset showing the angle, θ, whichcan be used to quantify butterfly distortion; (d) the evolution of butterfly distortion with each weld pass.

J.A. Francis, et al. Journal of Materials Processing Tech. 274 (2019) 116269

8

4.4. Metallography and hardness

Macrographs are presented in Fig. 8 for the 75mm thick trial weldon A533 steel (Fig. 8a) and an 80mm thick weld in SA508 steel(Fig. 8b). It is evident that desirable fusion zone geometries wereachieved in both cases, and that no significant defects are visible.

Although some efforts were made to mitigate microstructural hetero-geneity through the choice of beam oscillation parameters, the micro-structures along the weld centreline appear to have etched differentlyto other regions within the fusion zone. Further evaluation would re-quire more detailed microstructural characterisation, which is beyondthe scope of this article.

The hardness distributions are shown in Fig. 9 for an SA508 steelweld, both before and after PWHT. The PWHT operation was effective,since it reduced the peak hardness values to ∼ 300 HV, and thereforethe joint would be expected to exhibit reasonable toughness. The regionof elevated hardness after PWHT appears to be narrower than it is in theas-welded condition, which suggests that it is likely to be confined tothe fusion zone. Based on these results, one would expect the fusionzone to be stronger than the parent material.

4.5. Mechanical properties

All mechanical testing was carried out after the PWHT operation.The cross-weld tensile tests all failed in the parent material, and themeasured values for the UTS were 621MPa (top), 619MPa (middle)and 617MPa (bottom). Thus, there was consistency between the threetests and there did not appear to be any significant variation in prop-erties through the thickness of the weld. Values for the yield stress andelongation are not available owing to the heterogeneous nature ofstrain and deformation in cross-weld tensile tests. The results of thebend tests were acceptable in both cases.

The results of the Charpy tests on the vacuum laser welds in SA508steel are plotted in Fig. 10 (blue markers), which also includes a plot forthe parent material (black markers), as well as corresponding resultsfrom the 130mm thick welds described by Rathod et al. (2019). Theresults for the arc welding processes appear in grey, while the results forthe EB process appear in red.

The results for the vacuum laser weld appear to at least match thoseof the parent material at most temperatures, although it is possible thatthe upper shelf energies are slightly lower than those for the parentmaterial. Curiously, the results for the vacuum laser weld exceed thosefor the EB weld, both in terms of the ductile-to-brittle transition tem-perature and the upper-shelf energies. The reasons for this are notimmediately clear, since both processes are autogenous. In contrast, thearc welding processes employed a filler material with a lower carbon

Fig. 7. An overview of the mesh density andboundary conditions that were employed (top)in the contour method analysis, and estimatedlongitudinal residual stresses (bottom) for thetransverse half-plane at the mid-length posi-tion along the weld. The stresses correspond tothe as-welded condition, and the cutting di-rection for the contour method was from rightto left.

Fig. 8. Macrographs for (a) a 75mm thick trial weld in A533 steel and (b) aweld in SA508 Gr. 3 Cl. 1 steel. Sound welds were produced in both cases.

J.A. Francis, et al. Journal of Materials Processing Tech. 274 (2019) 116269

9

concentration to that of the parent material, and both of these processesinvolved a large number of weld passes.

5. Discussion

The results of this work auger well for vacuum laser welding as acandidate process for the future fabrication of nuclear pressure vessels.All indications in this work suggest that the process can reliably pro-duce welds of a suitable quality for safety-critical nuclear components.Clearly, further evaluation will be necessary before the process can beimplemented, particularly with respect to characterising the micro-structures in such welds, assessing the response of vacuum laser weldsto irradiation, and developing an understanding of how the laser beamoscillation parameters can be chosen in such a way as to ensure that theroot bead profile is acceptable. The latter aspect, however, may ulti-mately become moot if, in the future, available laser powers enable fullpenetration welds to be made in a single pass.

Vacuum laser welding clearly offers the potential for significantproductivity gains over arc welding processes. Indeed, the naturalcompetitor to vacuum laser welding for thick section components is EB

welding. Given that there is a greater body of experience with the EBprocess, and that the EB process is already capable of making very thickwelds in a single pass, one might ask why there might be interest invacuum laser welding for such applications. This question can be an-swered in part by pointing out that laser beams (or photons) are notprone to deflection when there is residual magnetism within the steel.This is pertinent to the welding of thick sections, because large steelcomponents are often lifted using magnets. Laser beam radiation doesnot generate X-rays when it impinges on a metallic surface, so theshielding requirements for vacuum laser welding are likely to be lessonerous than for EB welding. Vacuum laser welding is also less sensitiveto the quality of vacuum than the EB process, as pointed out by Reisgenat al. (2015). The vacuum technology is therefore cheaper, and thecycle times are faster than for EB welding. These factors suggest that thepotential exists for vacuum laser welding to offer improved robustnesswith simpler shielding arrangements. On the other hand, EB weldingoffers a superior penetration capability at present, and it is easier tocontrol the oscillation of an electron beam (through electromagneticfields) than it is for a laser beam (which requires oscillating mirrors).On balance though, the authors believe that the potential advantages of

Fig. 9. Macro-section (a) and hardness maps for a vacuum laser weld in SA508 Gr. 3 Cl. 1 steel in the as-welded condition (b) and after PWHT (c).

Fig. 10. A comparison of the Charpy impacttest results for samples extracted from the HAZof a vacuum laser weld with those obtainedfrom the parent steel, and with results fromsimilar samples that were extracted from theHAZ of the 130mm thick welds described byRathod et al. (2019). All welds were made inthe same heat of material but with differentwelding processes. Results are also shown forGTAW, SAW and EB welding.

J.A. Francis, et al. Journal of Materials Processing Tech. 274 (2019) 116269

10

vacuum laser welding are significant enough for the process to warrantfurther development for thick-section nuclear components.

In terms of the mechanical performance of the vacuum laser welds,it would appear that the residual stress distributions are similar innature to those for the EB process, which is unsurprising, given thatboth processes employ a high-energy beam to weld in keyhole mode.However, the results in Fig. 10 are interesting, since the Charpy energylevels for vacuum laser welding are clearly superior to those for the EBprocess in the same steel, both in terms of the DBTT and the upper shelfenergy. The reasons for the differences are not obvious, and the resulttherefore needs to be treated with caution at this stage. For example, itis possible that the requirement for two weld passes played some role inaffecting the Charpy values. Clearly, a more systematic comparisonbetween the mechanical properties of EB welds and vacuum laser weldsis needed, where the same plate thickness, welding speed and beamoscillation parameters are used in both cases. However, if there were tobe a physical reason for vacuum laser welding offering improvements intoughness, this would add to the list of incentives for further develop-ment. Any future development work would certainly need to involveboth experimental trials and modelling.

6. Conclusions

The following conclusions can be drawn from this work:

• Vacuum laser welding can be applied to thick-section nuclearcomponents while achieving the necessary weld quality;

• The process offers significant productivity gains when compared toarc welding processes, given that 80mm thick joints were com-pleted in two weld passes;

• Vacuum laser welding is likely to be a robust process, since initialbead-on-plate trials translated to immediate success in producingacceptable 80mm thick steel welds;

• The mechanical performance of the vacuum laser welds at leastmatched the performance of a 130mm thick electron beam weld inthe same heat of SA508 steel;

• Further work is required to understand how the beam oscillationparameters can be optimised to achieve an acceptable root beadprofile, and to establish whether variability in the grain structurewithin the fusion zone can be minimized through an appropriatechoice of welding parameters.

Acknowledgements

The authors are grateful to the Engineering and Physical SciencesResearch Council (EPSRC) for financial support provided through theNNUMAN programme grant (Grant number: EP/J021172/1). The au-thors would also like to thank Ioannis Pantelis, Ian Winstanley and PaulEnglish at The University of Manchester (UOM) for technical assistance.Finally, the authors would like to thank Jacqui Grant (UOM) for projectmanagement support.

References

American Society of Mechanical Engineers, 2017. ASME Boiler and Pressure Vessel Code.Section IX. .

Arata, Y., Abe, N., Oda, T., Tsujii, N., 1985. Fundamental phenomena during vacuumlaser welding. Proc. of ICALEO’ 85 Materials Processing Symposium 1–7.

Arlt, A.G., 2012. Laser Welding: Line Scanners for Beam Shaping and Guiding. Fügen imKarosseriebau, Bad Nauheim, Germany.

ASTM International, 2017. ASTM A370-17a: Standard Test Methods and Definitions for

Mechanical Testing of Steel Products.Ayres, K.R., Hurrell, P.R., Gill, C.M., Bridger, K., Burling, L.D., Punshon, C.S., Wei, L.,

Bagshaw, N., 2010. Development of reduced pressure electron beam welding processfor thick section pressure vessel welds. Proceedings of the ASME 2010 PressureVessels and Piping Division Conference, Paper PVP2010-25957 Bellevue,Washington, USA.

Balakrishnan, J., Vasileiou, A.N., Francis, J.A., Smith, M.C., Roy, M.J., Callaghan, M.D.,Irvine, N.M., 2018. Residual stress distributions in arc, laser and electron-beam weldsin 30 mm thick SA508 steel: a cross-process comparison. Int. J. Press. Vessel. Pip.162, 59–70.

Feng, J.C., Rathod, D.W., Roy, M.J., Francis, J.A., Guo, W., Irvine, N.M., Vasileiou, A.N.,Sun, Y.L., Smith, M.C., Li, L., 2017. An evaluation of multipass narrow gap laserwelding as a candidate process for the manufacture of nuclear pressure vessels. Int. J.Press. Vessel. Pip. 157, 43–50.

Francis, J.A., Bhadeshia, H.K.D.H., Withers, P.J., 2007. Welding residual stresses in fer-ritic power-plant steels. Mater. Sci. Technol. 23 (9), 1009–1020.

Holtum, N., Jakobs, S., 2018. Potentiale des Laserstrahlschweißens unter stationärem undmobile Vakuum. Im Blickpunkt. Deutschlands elite-Institute: Institut fürSchweißtechnik und Fügetechnik der RWTH Aachen University, AlphaInformationsges, Lampertheim, pp. 33–38.

Honig, R.E., Kramer, D.A., 1969. Vapor Pressure Data for the Solid and Liquid Elements.RCA, pp. 285–305 Review 30, June.

Jakobs, S., Reisgen, U., 2015. Laser beam welding in vacuum of fine-grained steel –learning the good things from EB. Proceedings of the 8th International Conference onBeam Technologies & Laser Application.

Jeyaganesh, B., Callaghan, M.D., Francis, J.A., English, P.D., Vasileiou, A., Roy, M.J.,Guo, W., Irvine, N.M., Smith, M.C., Li, L., Sherry, A.H., 2014. Overview of weldingresearch under the New Nuclear Manufacturing (NNUMAN) Programme. Proceedingsof the ASME 2014 Pressure Vessels and Piping Conference. https://doi.org/10.1115/PVP2014-29015. Paper PVP2014-29015.

Jiang, M., Tao, W., Chen, Y., 2017. Laser welding under vacuum: a review. Appl. Sci. 7(9), 909. https://doi.org/10.3390/app7090909.

Johnson, G., 2008. Residual Stress Measurements Using the Contour Method. PhDThesis. The University of Manchester, pp. 243.

Kundu, A., Kumar, S., Venkata, K.A., Bouchard, P.J., Francis, J.A., Paradowska, A., Dey,G.K., Truman, C.E., 2013. Residual stresses in P91 steel electron beam welds. Sci.Technol. Weld. Join. 18 (1), 70–75.

Luo, Y., Tang, X., Lu, F., Chen, Q., Cui, H., 2015. Effect of subatmospheric pressure onplasma plume in fiber laser welding. J. Mater. Process. Technol. 215, 219–224.

Prime, M.B., 2001. Cross-sectional mapping of residual stresses by measuring the surfacecontour after a cut. J. Eng. Mater. Technol. 123 (2), 162–168.

Prime, M.B., Sebring, R.J., Edwards, J.M., Hughes, D.J., Webster, P.J., 2004. Laser sur-face-contouring and spline data-smoothing for residual stress measurement. Exp.Mech. 44 (2), 176–184.

Rathod, D.W., Francis, J.A., Roy, M.J., Obasi, G., Irvine, N.M., 2017. Thermal cycle-de-pendent metallurgical variations and their effects on the through-thickness me-chanical properties in thick section narrow-gap welds. Mater. Sci. Eng. A 707,399–411.

Rathod, D.W., Francis, J.A., Vasileiou, A.N., Roy, M.J., English, P.D., Balakrishnan, J.,Smith, M.C., Irvine, N.M., 2019. Residual stresses in arc and electron-beam welds in130 mm thick SA508 steel: part 1 – manufacture. Int. J. Press. Vessel. Pip. 172,313–328.

Reisgen, U., Olschok, S., Jakobs, S., Turner, C., 2015. Laser beam welding in vacuum –overview of thick-plate steel application and beyond. Proceedings of Lasers inManufacturing 2015 - LIM 2015 and World of Photonics Congress.

Sanderson, A.S., Punshon, C.S., Russell, J.D., 2000. Advanced welding processes for fu-sion reactor fabrication. Fusion Eng. Des. 49–50, 77–87.

Smith, D.J., Zheng, G., Hurrell, P.R., Gill, C.M., Pellereau, B.M.E., Ayres, K., Goudar, D.,Kingston, E., 2014. Measured and predicted residual stresses in thick section electronbeam welded steels. Int. J. Press. Vessel. Pip. 120–121, 66–79.

Sun, Y.L., Roy, M.J., Vasileiou, A.N., Smith, M.C., Francis, J.A., Hosseinzadeh, F., 2017a.Evaluation of Errors Associated with Cutting-Induced Plasticity in Residual StressMeasurements Using the Contour Method. Experimental Mechanics 57, 719–734.

Sun, Y.L., Roy, M.J., Vasileiou, A.N., Smith, M.C., Francis Hosseinzadeh, J.A.F., 2017b.Evaluation of Errors Associated with Cutting-Induced Plasticity in Residual StressMeasurements Using the Contour Method. Experimental Mechanics 57, 719–734.

Teichmann, F., Muller, S., Dilger, K., 2018. On the occurrence of weld bead porosityduring laser vacuum welding of high pressure aluminium die castings. Procedia Cirp74, 438–441.

Vasileiou, A.N., Smith, M.C., Francis, J.A., Rathod, D.W., Balakrishnan, J., Irvine, N.M.,2019. Residual stresses in arc and electron-beam welds in 130 mm thick SA508 steel:part 2 – measurements. Int. J. Press. Vessel. Pip. 172, 379–390.

Wei, P.S., Chuang, K.C., Ku, J.S., DebRoy, T., 2012. Mechanisms of spiking and humpingin keyhole welding. IEEE Trans. Compon. Packaging Manuf. Technol. 2 (3), 383–394.

Zhang, X., Ashida, E., Tarasawa, S., Anma, Y., Okada, M., Katayama, S., Mizutani, M.,2011. Welding of thick stainless steel plates up to 50mm with high brightness lasers.J. Laser Appl. 23, 022002. https://doi.org/10.2351/1.3567961.

J.A. Francis, et al. Journal of Materials Processing Tech. 274 (2019) 116269

11