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journal of materials processing technology 196 ( 2 0 0 8 ) 279–291 journal homepage: www.elsevier.com/locate/jmatprotec Experimental and numerical investigations: Alleviating tensile residual stresses in flash-butt welds by localised rapid post-weld heat treatment David Tawfik a,, Peter John Mutton b , Wing Kong Chiu a a Mechanical Engineering Department, Monash University, Melbourne, Australia b Institute of Railway Technology, Monash University, Melbourne, Australia article info Article history: Received 7 August 2006 Received in revised form 13 November 2006 Accepted 29 May 2007 Keywords: Rail welding Residual stresses Post-weld heat treatment Finite element model abstract Flash-butt welding is commonly used in the manufacture of continuously welded rails (CWR). The finished welds typically exhibit high levels of tensile residual stresses in the rail web. In addition, the surface condition of the web may contain shear drag or other defects resulting from the shearing process. When combined with torsional loading of the web under service loading (particularly at high axle loads), these conditions may contribute to fatigue failure of the weld in a horizontal split web mode. The risk of weld failure may be alleviated by reducing the magnitude of the tensile residual stresses. A sequentially cou- pled thermo-mechanical finite element (FE) model incorporating the phase transformation characteristics of the rail material has been used to predict the residual stress distribution developed during welding of AS60 and AS68 rails, by approximating the thermal distribution after upset from the heat-affected zone (HAZ) characteristics. The effect of localised, rapid post-weld heat treatment on residual stresses in the web region of the weld was also inves- tigated. An experimental program covering measurement of post-weld cooling rates using infra-red thermography, and residual stresses by the strain-gauge and trepanning, was used to validate the finite element model. The results have shown that the FE model can satisfac- torily predict the residual stress distribution. In addition, the (tensile) residual stress levels in both vertical and longitudinal directions of the web can be reduced by rapidly reheating the base of the foot directly after welding. Thereafter, both numerical and experimental approaches will be used to develop modifications to the flash-butt welding procedure that should result in improved weld performance under high axle load conditions. © 2007 Elsevier B.V. All rights reserved. 1. Introduction Continuously welded rail (CWR) has largely replaced mechan- ically jointed track as the accepted method of rail joining. Welded rails reduce impact loading, thereby facilitating a reduction in track inspection and maintenance require- ments. Although a welded joint may be superior to a bolted Corresponding author. E-mail address: david.tawfi[email protected] (D. Tawfik). joint in strength, welds still represent a discontinuity in the track structure due to variations in microstructure, mechanical properties and residual stress levels relative to the parent rail (Marich et al., 1983). These variations may contribute to an increased risk of weld failure under high axle load conditions (Marich et al., 1983; Mutton, 1994). 0924-0136/$ – see front matter © 2007 Elsevier B.V. All rights reserved. doi:10.1016/j.jmatprotec.2007.05.055

Experimental and numerical investigations: Alleviating tensile residual stresses in flash-butt welds by localised rapid post-weld heat treatment

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Page 1: Experimental and numerical investigations: Alleviating tensile residual stresses in flash-butt welds by localised rapid post-weld heat treatment

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j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291

journa l homepage: www.e lsev ier .com/ locate / jmatprotec

xperimental and numerical investigations: Alleviatingensile residual stresses in flash-butt welds by localisedapid post-weld heat treatment

avid Tawfika,∗, Peter John Muttonb, Wing Kong Chiua

Mechanical Engineering Department, Monash University, Melbourne, AustraliaInstitute of Railway Technology, Monash University, Melbourne, Australia

r t i c l e i n f o

rticle history:

eceived 7 August 2006

eceived in revised form

3 November 2006

ccepted 29 May 2007

eywords:

ail welding

esidual stresses

ost-weld heat treatment

inite element model

a b s t r a c t

Flash-butt welding is commonly used in the manufacture of continuously welded rails

(CWR). The finished welds typically exhibit high levels of tensile residual stresses in the

rail web. In addition, the surface condition of the web may contain shear drag or other

defects resulting from the shearing process. When combined with torsional loading of the

web under service loading (particularly at high axle loads), these conditions may contribute

to fatigue failure of the weld in a horizontal split web mode. The risk of weld failure may

be alleviated by reducing the magnitude of the tensile residual stresses. A sequentially cou-

pled thermo-mechanical finite element (FE) model incorporating the phase transformation

characteristics of the rail material has been used to predict the residual stress distribution

developed during welding of AS60 and AS68 rails, by approximating the thermal distribution

after upset from the heat-affected zone (HAZ) characteristics. The effect of localised, rapid

post-weld heat treatment on residual stresses in the web region of the weld was also inves-

tigated. An experimental program covering measurement of post-weld cooling rates using

infra-red thermography, and residual stresses by the strain-gauge and trepanning, was used

to validate the finite element model. The results have shown that the FE model can satisfac-

torily predict the residual stress distribution. In addition, the (tensile) residual stress levels

in both vertical and longitudinal directions of the web can be reduced by rapidly reheating

the base of the foot directly after welding. Thereafter, both numerical and experimental

approaches will be used to develop modifications to the flash-butt welding procedure that

should result in improved weld performance under high axle load conditions.

to the parent rail (Marich et al., 1983). These variations

. Introduction

ontinuously welded rail (CWR) has largely replaced mechan-cally jointed track as the accepted method of rail joining.

elded rails reduce impact loading, thereby facilitating aeduction in track inspection and maintenance require-

ents. Although a welded joint may be superior to a bolted

∗ Corresponding author.E-mail address: [email protected] (D. Tawfik).

924-0136/$ – see front matter © 2007 Elsevier B.V. All rights reserved.oi:10.1016/j.jmatprotec.2007.05.055

© 2007 Elsevier B.V. All rights reserved.

joint in strength, welds still represent a discontinuity inthe track structure due to variations in microstructure,mechanical properties and residual stress levels relative

may contribute to an increased risk of weld failure underhigh axle load conditions (Marich et al., 1983; Mutton,1994).

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n g t e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291

Table 1 – Chemical composition of 60 kg rail (wt%)

up to 1500 ◦C with a field of view focus of 24◦ × 18◦ with a320 × 240 pixel resolution. Hence, the measured planar arearepresented by 1 pixel would cover a 1.3 mm × 1.3 mm objectarea per metre length away from the subject sample. Atmo-

280 j o u r n a l o f m a t e r i a l s p r o c e s s i

Flash-butt welds produced using both stationary (fixed)and mobile welding machines have been reported to incurhigh tensile residual stress levels in the web region of theweld (Skyttebol and Josefson, 2004; Mansouri et al., 2004). Themagnitude of these stresses and the hardness distributionacross the weld (which in turn may influence the magnitudeof impact loading) are influenced by a range of welding param-eters, including preheating conditions, upset conditions (inparticular upset distance), post-weld cooling conditions andthe characteristics of the rail material.

Limited experimental residual stress measurements, par-ticularly in the web region of flash-butt welds, have showninconsistent results. Several reports (Marich et al., 1983;Mutton, 1994; Carpenter and Sonon, 1983; Urashima et al.,1986; Tawfik et al., 2005) have shown that vertical resid-ual stresses are more dominant in this region, but neutrondiffraction measurements (Ezeilo et al., 1997) revealed thatlongitudinal strains were more significant. Finite elementtechniques have been used in a limited number of cases toexamine the thermo-mechanical behaviour of rails duringflash-butt welding (Skyttebol and Josefson, 2004).

Attempts to examine the behaviour of residual stressesin welds or in manufactured steel components by rapid heattreatments such as induction (Mansouri et al., 2004), rapid fur-nace (Vershinin, 1972) and flame heating (Canas et al., 1996)are limited, with little information on the effect of such treat-ments on the microstructural and hardness characteristics.Moreover, other research has revealed that certain post-weld-heating techniques may lower residual stresses in one planardirection (vertical direction) but actually raise tensile residualstresses in the other planar direction (longitudinal direction)and cause worsening of the weld material properties (Canaset al., 1996).

Currently, the accepted technique to alleviate residualstresses in welded joints is by thermal annealing. While FEsimulations and experimental results have demonstrated thatlong-duration thermal annealing is effective (Sedek et al.,2003), its practical application is ultimately time consumingand cost ineffective, particularly for production welding dur-ing fabrication of rail strings.

This paper described the results of an investigation, usingnumerical modelling techniques, of the influence of variouswelding parameters such as increasing heat affected zonewidths on the residual stress distribution in flash-butt welds inAS60 and AS68 rails. In addition, a preliminary analysis on theeffects of localised rapid post-weld heat treatment on resid-ual stress levels was undertaken. The numerical results werecompared with experimental results including residual stressdata obtained using strain-gauge techniques, and microstruc-tural and hardness characteristics of finished welds. Theoverall objective of this program is to develop a rapid post-weld heat treatment procedure than can be used to reducethe risk of fatigue failure in flash-butt welds under high axleload conditions.

2. Experimental procedure

The chemical composition of the AS60 rail steel [referenceAS1085.1] used throughout the experimental program is pre-

C Mn Si S P Cr Al

0.8 0.87 0.19 0.014 0.017 0.02 0.001

sented in Table 1. The rail was used in the as-rolled condition,with a nominal hardness of 280 HB.

Ten flash-butt welds were produced during the weldingtrials conducted under normal operating conditions at therail welding depot owned and operated by the John HollandGroup as part of the Regional Rail Link project in Australia.The flash-butt welds were performed using a Plasser mobilewelder which operated at 6–7 V with an output power ofapproximately 300 kW. The final upset consumed approxi-mately 32 mm of rail. The finished welds were removed inlengths of 650 mm, using an abrasive rail saw, to preventany significant changes to the residual stress distribution inthe vicinity of the weld. Standard welding procedures whichinclude grinding of the sheared surface of the weld was notaccounted in the FE model and the experimental program forsimplicity of analysis. Previous research has shown that grind-ing of the weld surface at the correct (elevated) temperatureshould not have any significant effect on the final residualstress distribution (Mahdi and Zhang, 1999). Prior to welding,the rail ends and web surfaces were ground by small abra-sive grinders to ensure good electrical contact between theelectrodes and rail surface.

Six welds were performed under normal operating andcooling conditions and other four welds underwent localisedpost-weld heat treatment in the rail foot. A 250 mm × 145 mmarea on the underside of the welded foot was heated (Fig. 1)by means of high pressure gas flame burners applied approx-imately 80 s after final upset for durations of up to 300 s.

Cooling rates at the surface of the welded region weremonitored by a FLIR Therma-Cam PM695® Infrared (IR) ther-mographic long wave camera (FLIR, 2003). The infra-redthermographic camera was capable of recording temperatures

Fig. 1 – Simplified setup for a rapid localised post-weld heattreatment for a flash-butt weld.

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j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291 281

Fig. 2 – Etched longitudinal sections through flash-butt welds: (a) normal cooled and (b) short-term post-weld heat-treated.

sae

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2hnpandniztito

pheric relative humidity was set to 50% while both object andtmospheric ambient temperature were set to 20 ◦C. Objectmissivity was tested and found to be 0.95.

The thermographic infra-red camera commenced record-ng prior the final flashing stage. However, attempts to

easure peak temperatures directly after upset were delayedy 20–24 s due to visual obstruction by the shearing head ofhe welding machine. Thereafter, cooling rates were recordedn small increments along the foot and along the neutral axis,t 0 mm, 25 mm and 50 mm from the fusion line.

.1. Metallographic examination and hardnesseasurements

ections extracted from the head, web and foot regions ofnormal cooled weld, and a weld subjected to short-term

ost-weld heat-treatment in the foot, were ground and macro-tched with 2% nital. The full width of the visible heat-affectedone (HAZ) varied over the height of the rail section, as shownn Fig. 2(a) and (b). In addition, segregation within the parentail was visible in the web region of both welds. The full widthf the visible heat affected zone in the head was 42 mm and6 mm in the normal cooled and post-weld heat-treated welds,espectively. In the web the corresponding width approached6 mm for the post-weld heat-treated web. The foot showedhe narrowest HAZ widths, ranging from 30 mm in the in theormal cooled weld to 34 mm in the post-weld heat-treatedelds.

Vickers hardness measurements (5 kg load) were taken atmm spacing on longitudinal traverses across the welds at theead, web and foot regions. The results revealed similar hard-ess distribution in the head, between the normal cooled andost-heated weld (Fig. 3(a)). The corresponding measurementscross the web (Fig. 3(b)) revealed more variability in hard-ess within each weld, and between welds; however, theseifferences were primarily due to the location of the hard-ess indentations relative to the segregated region present

n the parent rail. In the rail foot, the hardness in the fusionone of the post-heated weld was approximately 30 HV lower

han that in the normal cooled weld (Fig. 3(c)). The variationn hardness between these welds may have been caused byhe application of rapid post-weld heat treatment to the footf the weld.

2.2. Microstructural observations

Sections extracted from normal cooled and post-weld foot-heated welds, close to the locations of the hardnessmeasurements, were metallographically polished to 1 �m andetched with 2% nital. Microscopic examination was carriedout at the fusion line, within the fusion zone, at the tran-sition between heat-affected zone and parent rail material,and within the parent rail material, under an optical micro-scope. Fig. 4(a) and (b) shows the microstructures within theseregions, for the head and foot regions, respectively, in the nor-mal cooled weld. The corresponding regions in the foot of thepost-weld foot-heated weld are shown in Fig. 4(c).

Both welds exhibited pro-eutectoid ferrite at the prior-austenite grain boundaries at the fusion line. This is typicalof flash-butt welds, and delineates the fusion line as shownin the macro-etched samples in Fig. 2. This was only region inwhich the prior-austenite grain size could be determined, asdelineated by the pro-eutectoid ferrite. In the remainder of thefusion zone the microstructure was fully pearlitic in both headand foot regions. In the foot, some variation in the pearlitespacing (consistent with the differences in hardness evidentin Fig. 3(c)) may have been present; however, these were notdetectable at the magnifications used. In the HAZ/parent railtransition, the pearlite was partially spheroidised, with nodetectable difference between normal cooled and post-weldfoot-heated conditions. The microstructure in the parent railwas fully pearlitic.

3. Numerical procedure

3.1. Numerical dilatometry validation

Numerical simulations were conducted on a thin walled plaincarbon steel cylinder similar in chemical composition to thatof the rail steel, to evaluate the volumetric expansion dueto thermal and phase transformation effects compared withexperimental results. These parameters are known to be

major contributors to the formation of residual stresses inwelding (Rammerstorfer et al., 1981; Runnesson et al., 2003;Rohde and Jeppsson, 2000; Ronda and Oliver, 2000; Storesundand Tu, 1995; Toshioka, 1986). The numerical simulations were
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282 j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291

rmat: 90

Fig. 3 – Longitudinal hardness distributions measured in no5 kg): (a) head (height above rail foot: 164 mm), (b) web (heigh

undertaken using the commercial code ABAQUS Standard(Hibbit, 2003).

Experimental dilatometry was undertaken on a high-speedquenching and deformation dilatometer manufactured byMaterials Measuring Corporation, using RF heating from a sin-gle turn induction coil (Prior and Mitchell, 1993). A plain carbon

grade WK1072 (0.72% carbon) FE model was constructed of573 tetrahedral solid elements and 1245 nodes. The nomi-nal dimensions of this tubular specimen were 12.7 mm longwith an external diameter of 7.3 mm and a wall thickness

l cooled and short-term post-weld heat-treated welds (HV,mm), (c) foot (height: 6 mm).

of 1 mm. The tubular specimen was subjected to coolingfrom 1000 ◦C to 100 ◦C at 1 ◦C/s. The normalized length ofthe tube was recorded at 50 ◦C intervals, and compared withthe numerical results as shown in Fig. 5. The results showedrelatively good agreement between the experimental andnumerical measurements over the temperature range. More-

over, it showed the ability of the ABAQUS to model thevolumetric expansion of the material due to phase trans-formation in the temperature interval between 650 ◦C and700 ◦C.
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j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291 283

F ledp

3v

At2

FntW

become increasingly common for the simulation of both weld-ing and post-weld heat treatments to be broken up in twoconsecutive steps (Fig. 6(b)) by a sequentially coupled thermo-

ig. 4 – Optical micrographs of normal cooled: (a) normal cooost-weld foot-heated (foot only) weld.

.2. Sequentially coupled thermo-mechanical analysisalidation

lthough the welding process is ideally described by a coupledhermo-mechanical-metallurgical analysis (Ronda and Oliver,000) which occurs in the heat-affected zone (Fig. 6(a)), it has

ig. 5 – Comparison of experimental measurement andumerical simulation of dilatometric behaviour due to

hermal expansion and phase transformation behaviour ofK1072 grade steel.

weld in the head, (b) normal cooled weld in the foot, (c)

mechanical analysis (Marich et al., 1983; Carpenter and Sonon,

Fig. 6 – Thermo-mechanical analysis: (a) fully coupled(1-2-3-4-5-6-7), (b) sequentially coupled (1-2-4-6).

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n g t

284 j o u r n a l o f m a t e r i a l s p r o c e s s i

1983; Gordon et al., 1998; Fischer et al., 1988; Sarkani et al.,2000). Although this allows the simulation to be performedmore efficiently using less computational time, it must beused with caution as the volumetric change due to phasetransformation and the associated localised deformation donot generate sufficient latent heat to significantly alter thethermal distribution (Runnesson et al., 2003; Sarkani et al.,2000). A sequentially coupled analysis was implemented tomodel flash-butt welding. However, a preliminary quenchingsimulation was performed in ABAQUS in order to comparethe thermal stress behaviour using the sequentially coupledthermo-mechanical analysis with results obtained by anothernumerical code (ANSYS) performed by Sen et al. (Sen et al.,2000).

Fig. 7 shows a 30 mm long, 10 mm diameter cylinder madeof steel grade SAE1020. It was assumed that the specimen ini-tially held at 600 ◦C was completely homogenized and stressfree before quenching in water. In addition, symmetry con-ditions along the axial centre line (OM) and radial centreline (OK) only required a quarter axis symmetrical geometrymodel. The model incorporated a bilinear elastic-plastic mate-rial model including linear kinematic hardening. The effectsof phase transformation on residual stresses were not takeninto account in the published work. The thermal and mechan-ical properties and boundary conditions were implemented asdetailed by Sen et al. (Sen et al., 2000).

Fig. 8 shows a comparison of the thermal history of pointsO, K and M of the first sequence of decoupled analysis betweenthe numerical model and published results. Temperature pre-dictions between ABAQUS and ANSYS showed very goodagreement, see Fig. 8(a). The thermal results were then usedas the basic load inputs to produce the transient thermalstress plots for the correlating elements. Again, Fig. 8(b) showsclose agreement between the transient thermal stresses inthe radial, axial and circumferential (hoop) direction on thesurface (K), core (O, M) of the cylinder and that of publishedresults. In addition, the final residual stress distribution alongthe radial “OK” elements was verified as shown in Fig. 8(c).The results confirmed the creditability of the sequentially cou-pled thermo-mechanical technique when used to estimate theresidual stresses distribution in rail welds.

3.3. Finite element model of AS60 and A68 flash-buttwelded rail

A finite element model of flash-butt welds in AS60 and AS68rails has been developed using the commercial code ABAQUSStandard (Hibbit, 2003). Fig. 9(a) and (b) show orthographicprojections of the three dimensional FE AS60 and AS68 railmodels, respectively. Only one quarter of the 3000 mm longrail geometry was simulated due to the symmetrical nature ofthe rail geometry and welding process in the transverse andlongitudinal directions. The AS60 model comprised 8692 brickelements and a total of 10,138 nodes; the AS68 model com-prised 3888 brick elements and 4536 nodes. A finely refinedmesh was implemented in the region where the highest ther-

mal gradients occur in the heat affected zone. Thereafter, agradual coarsening of the mesh density was applied as the rateof temperature change decreases when moving away fromthe HAZ. This biased meshing technique has previously been

e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291

implemented in numerical welding and heat treating simu-lations and reported to provide reasonable accuracy (Sarkaniet al., 2000). The FE model was kept as long as possible tosimulate a continuous welded rail, thus minimizing the uncer-tainties associated with prediction of the thermal behaviourand resultant residual stress distribution.

The cooling conditions used for the initial FE modellingwere set for normal cooling. The convection parameterfor the normal cooled weld was set to 6 w/m2, the radi-ation emissivity at 0.9 and Stefan–Boltzmann constant of5.6697 × 10−8 W/m2 K4. The assumption during the experi-mental heat treatment was that the flame burners evenlyheated the bottom region of the weld. This heating rate wasapproximated numerically by applying a uniform heat fluxof 120,000 W/m2 to the underside of the welded foot span of250 mm. The simulated heat treatment commenced 80 s afterupset for a duration of 300 s.

The nonlinear elastic-plastic post-welding behaviour wasmodelled by a sequentially coupled thermo-mechanicalanalysis incorporating temperature-dependent phase-transformation material behaviour (Fischer et al., 1988;Ringsberg and Lindbak, 2002). The data used were represen-tative values for steels of similar chemical composition usedin these experiments (Fischer et al., 1988). A constitutivelinear kinematic hardening model was used in the modelas opposed to isotropic hardening, since the material willexperience cyclic plasticity despite the monotonic heatingprocess. This has been used in other models used for theanalysis of welding and heat treatment, and shown to yieldmore accurate results (Rammerstorfer et al., 1981).

In addition, the temperature-anneal effect was accountedfor in both the welding and rapid post-weld heat treatment.The temperature-anneal effect initiates in the welding pro-cess when the localised temperature of a material exceedsa user-specified anneal temperature (Ta), causing the mate-rial point to lose its hardening memory. The effect of priorwork-hardening is removed by setting the equivalent plasticstrain to zero. For kinematic hardening the backstress ten-sor is also reset to zero. Once the temperature of the materialpoint drops below the annealing temperature at a point in timethe material may work harden again. Depending on the tem-perature history a material point may lose and accumulatememory several times. The anneal effect has been reported asan important parameter (Skyttebol and Josefson, 2004; Hibbit,2003) for both mechanical and material behaviour during anal-ysis of welding and post-weld heat treatments, and preventsthe material model over-predicting the accumulated plasticstrains. The anneal temperature for this finite element modelrail steel was set at 650 ◦C (Ringsberg and Lindbak, 2002). Thisis a representative value taken from a similar rail steel.

The material model did not take into account the effectsof time-dependent creep at high temperature due to theshort duration of the thermal cycle during welding. Previ-ous research has shown that creep effects (Sedek et al., 2003;Sarkani et al., 2000) have a minor effect on the overall residualstress distribution in similar welding and short duration heat

treatment processes.

The effects of rail fabrication have not been accounted forin the FE simulation although previous research (Schleinzerand Fischer, 2001) has shown that some aspects of the rail

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Fs

Fig. 7 – Diagram of (a) solid cylinder and (b) cons

ig. 8 – (a) Transient temperatures at point K, M and O. (b) Transitress distribution along the normalized line O-K.

Fig. 9 – Finite element model of (a) AS60

tructed axisymmetric mesh solid cylinder.

ent thermal stresses at element K, M and O. (c) Residual

and (b) AS68 flash-butt welded rails.

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286 j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291

istribution in flash-butt welds immediately after upset for HAZ

Fig. 11 – Heating and normal air cooling validation:Comparison between type K thermocouples and FLIR

Fig. 10 – (a) Visible HAZ in flash-butt weld. (b) Temperature dhalf widths between 15 mm and 30 mm.

manufacturing process, in particular roller straightening, havea significant influence on residual stresses in the rail. However,the residual stress distribution in the parent rail was assumedto have minimal effect on the final residual stress distributionin the vicinity of the weld. During the welding process, thefusion zone reaches melting temperature which is expectedto reset the rail material to a homogenised (stress free) state,relieving all previously accumulated residual stresses.

3.3.1. Initial thermal conditions by HAZ approximationAn alternative technique was implemented to model theflash-butt weld process commencing from the preheatthrough to the final upset stage. An approximation of thethrough-thickness thermal gradient in the heat-affected zonewas used to establish the peak thermal distribution obtainedjust before the cooling phase, i.e. immediately following upset.The width of the HAZ depends on the preheat and upset condi-tions, and tends to vary between 15 mm and 30 mm dependingon the rail size and grade, and the welder type, as illustratedby the visible HAZ widths in Fig. 10(a). The thermal gradientof the weld was approximated by an exponential temperaturedecay as a function of distance from the fusion line (Canas etal., 1996; Storesund and Tu, 1995). The two arbitrary thermalconstant in the equation can be determined by the outermostposition of the visible heat affected zone, which is known toreach a peak temperature just below the pearlite to austenitetransformation temperature (i.e. approximately 720 ◦C), andthe melting temperature of the rail material in the fusion zone.Hence, an exponential curve fitted to these two points wassufficient to describe the temperature distribution across theweld immediately after upset, as shown in Fig. 10(b) for typi-cal half width HAZ ranging from 15 mm to 30 mm. The averagewidth of the HAZ was measured by sectioning several welds.For the analysis of residual stresses in both normal cooled andrapid post-weld heat-treated welds the HAZ half width was setat 25 mm.

3.3.2. Thermal analysis validation using thermographic

techniqueA thermal validation test, involving a single heating and cool-ing cycle, had been performed on a 1200 mm long AS60 railto confirm the accuracy of the infra-red thermographic cam-

Therma-Cam Infra-red thermographic camera.

era. Type K thermocouples were embedded just beneath thesurface at the toe of the foot and 86 mm above the foot, at0 mm and 125 mm, respectively, from the rail centre. The railbase was steadily flame heated from ambient temperature of20 ◦C for approximately 330 s, and then allowed to air cool. Thetemperature–time history plot comparing the thermocoupleand thermographic temperature measurements is shown inFig. 11. The thermocouple and the infra-red thermographicdata showed satisfactory agreement, although the thermo-couple data indicated slight variances, possibly due to thestem of the thermocouple probes being exposed to the atmo-sphere and flame interference during the preliminary heatingstage.

4. Full scale flash-butt welding andshort-term post-weld heat treatmentexperiment

The measured cooling rates were compared with calculatedresults for normal cooled and post-weld heat-treated con-ditions as illustrated in Fig. 12(a) and (b). The temperature

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j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291 287

Fig. 12 – Cooling behaviour of a flash-butt weld across the web and foot region, respectively. (a) Normal cooled weld, (b)short-term post-weld heat treatment of the foot, (c) thermographic image of normal cooled 20 s after upset, (d)t g the

htctetnomtrptocr

4

Tsu2adsgasm

mal cooling welding conditions in Fig. 14(a) and (b). Themeasured principal stresses 1 and 2 were plotted againstthe height of the rail in Fig. 14(a). The experimental resultsconfirmed that peak tensile residual stresses occurred in

hermographic image during reheating of the rail foot durin

istory in a normal cooled weld showed better agreementhan the post-weld heat-treated weld. Flame temperatureslose to the foot and dirt from the weld accumulated athe foot may have accounted for the differences betweenxperimental and numerical data. Fig. 12(c) and (d) showhermographic images of the temperature distribution for aormal cooled weld 20 s after upset and during the heatingf the underside of the weld. The early stage of the nor-al cooling prior to initiation of the rapid post-weld heat

reatment indicated slower cooling rates in the mid-headegion with respect to the toe region. The initiation of rapidost-weld heat treatment caused a noticeable widening ofhe temperature distribution across the toe and foot regionsf the weld. After the rapid post-weld heat treatment wasompleted, cooling rates in the foot region were significantlyeduced.

.1. Experimental residual stress measurements

he rail welds were prepared for residual stress mea-urements using the strain-gauge and trepanning method,sing rectangular SHOWA strain-gauge rosettes N32-MA--120-11 (Procter and Beaney, 1987). The diameter of thennular trepanned region was 20 mm, with an averageepth of 9.5 mm. The rosettes comprised of three gaugestacked counter clockwise at 45◦ increments from the lon-

itudinal direction of the rail. The rosettes were fixedlong the rail in the configuration shown in Fig. 13. Thetrains were recorded using a Vishay quarter bridge straineter.

post-weld cooling period.

The rail material properties used to calculate the accu-mulated residual stress at room temperature were: Young’smodulus—206 GPa, Yield stress—480 MPa, Poisson ratio—0.33.The strain relief after trepanning was calculated on theassumption that plastic deformation does not occur whenresidual stresses relax during trepanning.

4.1.1. Results for normal cooled weldThe residual stress distribution obtained from the FE modelis compared with the measured strain-gauge data for nor-

Fig. 13 – Strain-gauge rosette configuration for residualstress measurements in an AS60 flash-butt weld.

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288 j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291

Fig. 14 – Principal residual stress distribution in an AS60 flash-butt weld for a 25 mm HAZ boundary condition: (a) above thefoot (mm), (b) along the neutral axis away from the fusion line, (c) Principal (1) stress (red), Principal (2) stress (blue) andprincipal (3) stress (black) along the neutral axis of an AS60 flash-butt welded specimen, (d) vertical residual stress contour

s co

along the centerline of the rail, (e) longitudinal residual stres

the vicinity of the mid-web region, within the fusion zone.The numerical results showed good agreement, and revealedpeak tensile residual stresses occurred up to 30–50 mm awayfrom the fusion line. In addition, Fig. 14(c) shows the ori-entation and relation of principal stresses 1 and 2 (alongthe neutral axis), across the weld, HAZ and parent rail withthe vertical (z) and longitudinal (y) directions, respectively.Principal stress 1 and principal 2 stresses are aligned predomi-nantly in the vertical and longitudinal directions, respectively,for up to 20 mm from the fusion line, beyond which theybegin to rotate away from the vertical and longitudinal direc-tions at a distance approaching 50 mm from the fusionline.

A detailed contour analysis of the of the vertical (z) and

longitudinal (y) residual stresses along the centreline of therail weld in Fig. 14(d) and (e) provided a clearer insight tothe distribution of residual stresses as predicted by the FEmodel.

ntour along the centerline of the rail.

4.1.2. Welding parameter effectsVarying the HAZ width under normal welding conditions (i.e.in the absence of any post-weld thermal treatments) wasfound to have a significant effect on the vertical residual stresslevel in the web region, as illustrated in Fig. 15(a). In generalnarrower HAZ zones, and hence higher residual stress levels,tend to occur in welds produced using fixed or stationary flash-butt welding machine, where preheating conditions may bemore rapid, and more material is expelled from the rail ends,compared to welds produced using a mobile flash-butt welder.Vertical tensile residual stresses increase to just over 550 MPain the web when the HAZ half width was decreased to 15 mm(Fig. 15(a)); however longitudinal stresses were shown to belargely unaffected by changing the HAZ width (Fig. 15(b)). In

addition, a comparison of AS60 and AS68 flash-butt welds,at the same HAZ widths, revealed that the larger AS68 railhad higher vertical residual stresses and lower longitudinalresidual stresses (Fig. 15(c)). However, experimental strain
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j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291 289

Fig. 15 – Effects of varying HAZ widths on residual stresses in the (a) vertical direction (z), (b) longitudinal direction (y) and (c)in an AS60 and AS68 flash-butt weld rails.

Fig. 16 – Experimental and numerical principal residual stresses after reheating the foot for 300 s at approximately 80 s afterupset. (a) At the fusion line and (b) along the neutral axis.

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r

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measurements on welds in AS68 rail are needed to confirmthese observations.

4.1.3. Results for rapid post-weld heat treatment of therail footThe results of the FE modelling and strain-gauge measure-ments have confirmed the validity of the former method, withmajor principal stress 1 close to the yield strength of the par-ent rail in both AS60 and AS68 flash-butt welds. Moreover, inthe event of torsional loading on the head of the rail, stresseson one side of the weld may exceed the yield strength, andincrease the risk of failure of the weld in a horizontal splitweb failure mode.

The FE model was subsequently used to examine the effectof short-term modifications to the post-weld cooling condi-tions on the residual stress distribution. This involved anattempt to lower cooling rates in the foot by re-heating thebase, with the aim of reducing the magnitude of the tensileresidual stresses in the web.

Previous numerical simulations of rail flash-butt welding(Tawfik et al., 2005) had shown indications of a reduction inresidual stress levels by localised short-term re-heating ofthe foot. New simulations using improved modelling parame-ters for material behaviour were compared with experimentalresults (Fig. 16(a) and (b)); this revealed significant reduc-tions in principal stresses in both directions in the webregion, as shown by a comparison with Fig. 14(a) and (b),respectively. Principal stress 1 showed a reduction of 25%while principal stress 2 showed a reduction of approximately46%.

5. Conclusions

The objective of the research program outlined in this paperis to develop a short-term post-weld heat treatment proce-dure than can be used to reduce the risk of fatigue failure inflash-butt welds under high axle load conditions. The workundertaken to date has involved the development of a numer-ical model capable of predicting the residual stress levels asa function of the temperature distribution immediately afterupset and the post-weld cooling conditions. Thorough val-idation of the numerical model, with improved modellingof the material behaviour, has been undertaken on flash-butt welds produced in AS60 rail, using a mobile weldingmachine.

An approximation of the temperature distribution in flash-butt welds immediately after upset, based on the dimensionsof the HAZ, has enabled the residual stresses to be estimatedwith reasonable accuracy.

Major and minor principal stresses in the region extendingto ±60 mm from the weld fusion line are oriented in the ver-tical and longitudinal directions, respectively, with maximumvalues of approximately 400 MPa and 300 MPa, respectively, atthe fusion line in AS60 rails.

Altering the post-weld cooling conditions, by reheating

the underside of the rail foot, has been shown to reduce themagnitude of the residual stresses without having significanteffect on the microstructural characteristics of the rail mate-rial. However, further work is required to identify the optimum

e c h n o l o g y 1 9 6 ( 2 0 0 8 ) 279–291

post-weld cooling parameters that will enable residual stresslevels to be reduced to the extent necessary to reduce the riskof weld fatigue failure.

6. Further work

A refined program of residual stress measurements by neutrondiffraction will take place on normal cooled welds. In addition,experimental residual stress measurements carried out on aseries of test welds that have been subjected to a range of post-weld cooling conditions, including accelerated cooling of therail web. These results will be used to refine the FE model,which will then be used to identify the required post-weldcooling conditions.

Acknowledgements

The authors would like to thank the following organisa-tions for their support: John Holland Group (supply of theAS60 rail and flash-butt welding), Mechanical EngineeringWorkshop, Monash University, Australasian Infrared Systems(post-processing software), Spectrographic Services (rail anal-ysis).

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