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Journal of Materials Processing Technology 211 (2011) 1547– 1552

Contents lists available at ScienceDirect

Journal of Materials Processing Technology

jou rna l h om epa g e: www.elsev ier .com/ locate / jmatprotec

hermo-mechanical modelling of residual stresses induced by martensitic phaseransformation and cooling during quenching of railway wheels

iva N. Lingamanaik ∗, Bernard K. Chenepartment of Mechanical and Aerospace Engineering, Monash University 3800, Australia

r t i c l e i n f o

rticle history:eceived 7 December 2010eceived in revised form 12 April 2011ccepted 14 April 2011vailable online 22 April 2011

a b s t r a c t

A finite element (FE) method was used to study the formation of residual stresses in low carbonbainitic–martensitic rail wheels. The FE model combines a commercially available heat treatmentsoftware DANTE to the finite element analysis software ABAQUS. Material data which include thermo-mechanical properties and kinetics of phase transformations for low carbon bainitic–martensitic (LCBM)steels were obtained from dilatometry experiments and added to DANTE material library. The results

eywords:esidual stressuenchingailway wheelsow-carbon bainitic–martensitic steels

showed that quenching conditions can be designed to promote the development of compressive residualstresses in the rim of LCBM rail wheels making it possible to produce LCBM steel rail wheels, which havesuperior properties compared to conventional pearlitic steels.

© 2011 Elsevier B.V. All rights reserved.

inite element modelling

. Introduction

During manufacture, rail wheels are quenched to achieveavourable mechanical properties as well as to promote beneficialesidual compressive circumferential stress in the rim of the wheel.utton and Lynch (2004) have shown the importance of compres-

ive residual stresses in rail wheels and their role in increasing theife of rail wheels by retarding the formation and growth of cracks.

Conventional rail wheels are made of medium to high carbonteels and have a characteristic pearlitic–ferritic microstructure aseported by Zhang and Gu (2008) and Devanathan and Clayton1991). A variety of grades are used within the Australian rail indus-ry ranging from the AAR Class A (nominally 0.47–0.57% C), Class B0.57–0.67% C) and Class C (0.67–0.77% C) depending on the typef service and degree of tread braking involved. Kwon et al. (2006)ave demonstrated that conventional pearlitic wheels are suscep-ible to the initiation of thermal fatigue cracks at the wheel/railontact zone. Significant thermal loads which occur during heavyread braking transform the pearlitic steel into brittle martensitefter subsequent rapid cooling of the wheels. Although theoriesor the formation of such cracks differ as reported by Mutton and

oelen (1989), there is evidence that pearlitic steels with higherarbon content are more susceptible to this problem. Therefore,ompressive residual stresses within the rim of the wheel are seen

∗ Corresponding author. Tel.: +61 03 990 53647.E-mail addresses: siva.lingamanaik@monash.edu

S.N. Lingamanaik), Bernard.Chen@monash.edu (B.K. Chen).

924-0136/$ – see front matter © 2011 Elsevier B.V. All rights reserved.oi:10.1016/j.jmatprotec.2011.04.007

to both retard fatigue crack growth and guard against sudden catas-trophic failure from large overloads.

Performance of pearlitic steel rail wheels has been improvedmainly by the removal of impurities in steels and strengthening ofwheel steels through the additions of alloying elements as demon-strated by Yokoyama et al. (2002). Lonsdale and Stone (2002) haveshown the advantages of micro-alloying elements in rail steels inwhich hardness levels have been increased up to 321 HB and wearresistance has also been improved by 20–25% over conventional railsteels. However, the current trend in heavy haul traffic for heavieraxle loads and higher speeds for passenger trains have seen theconventional pearlitic steel pushed to their limit and urged thedevelopment of new, stronger and more fatigue resistant steels.

As a result, low carbon bainitic–martensitic (LCBM) steels haverecently been developed as a promising material for rail wheelsas reported by Constable et al. (2006). Constable et al. (2006)have shown that LCBM steels have superior hardness and strength,toughness and ductility compared to conventional pearlitic steels.High hardness levels up to 415 HB have been recorded in LCBMsteels which is an increase of 18% compared to conventionalpearlitic steels. LCBM steels have also showed superior resistanceboth to thermal fatigue and formation of cracks which are benefi-cial for reliability and safety in railway wheels. Lonsdale and Stone(2002) have shown that martensitic rail wheels could significantlyimprove wear resistance of rail wheels. However, the effects of

austenitic to martensitic phase transformation which causes a netvolumetric expansion during quenching of LCBM steel rail wheelson the residual stress distribution in the rim of rail wheel have notbeen adequately evaluated.

1 rials Processing Technology 211 (2011) 1547– 1552

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548 S.N. Lingamanaik, B.K. Chen / Journal of Mate

Therefore, the aim of this paper is firstly to determine if theonventional rail wheel quenching process applied to LCBM steelail wheel will result in sufficiently high compressive residualtresses in the rim of the wheel required to retard the initiation andropagation of cracks: achieving the compressive residual stressesquivalent to those attained in pearlitic steel rail wheels. If thiss unsuccessful i.e. if conventional quenching process results inelatively low compressive residual stresses or even tensile resid-al stresses in the rim of the LCBM rail wheel, then, the next aimould be to determine if the quenching conditions can be altered

o achieve the desired level of compressive strength in the rim ofhe wheel.

. Methods

.1. Development of thermo-mechanical model for quenchingrocess

In this work, the modelling of quenching rail wheels was carriedut using a commercially available heat treatment software DANTE.3® coupled to the finite element analysis software ABAQUS.ANTE 3.3® (Deformation Control Technology) has inbuilt materialata for a range of steel which include the kinetics of metal-

urgical phase transformation and thermo-mechanical propertieshich have been obtained from an expansive range of temperatureependent tension-compression and dilatometry tests as shown byrantil et al. (2003).

In DANTE’s thermal module, heat transfer coefficients aressigned at surfaces of the model and local temperature and heatux histories are calculated. A phase transformation module inANTE (based on an internal variable framework) then calculatesnd tracks the evolution of the different phases from the local tem-erature data and also updates the thermal module for latent heateneration occurring from phase transformations as described byammann et al. (1996).

Phase transformation kinetics parameters can be obtained byeveral sources which includes CCT diagrams, TTT diagrams, Jominyardness test and dilatometry data. While TTT diagrams are mainlysed for diffusive transformations such as pearlite, CCT diagramsffer data for both diffusive and martensitic transformation aseported by Li et al. (2004). Jominy tests alone are not adequate foretermining kinetic phase transformations since strain–time dataannot be obtained as shown by Li et al. (2004). However, Li et al.2004) demonstrated that Jominy and TTT can be used to verifyinetic parameters.

In DANTE, dilatometry data are preferred over the above sourcess time, temperature and strain can be obtained from dilatometryxperiments and cooling transformation kinetic parameters can beasily verified against TTT and CCT data as described by Li et al.2004).

DANTE has an inbuilt fitting utility which can be used to obtaininetic parameters from dilatometry experiments. For marten-itic steels, phase transformation kinetics for martensite have beenetermined using the DANTE fitting function and good agreementas been found between predicted and experimental dilatometryata as reported by Ferguson et al. (2005). Since TTT and CCT dia-rams are available for common steels, they provide further checksn the fitting model performance as shown by Li et al. (2004). Fur-hermore, dilatometry data can be used and has shown to providehermal expansion and various phase transformation strains at dif-erent temperatures as described by Li et al. (2004).

DANTE mechanical module/solver is also based on an internalariable framework to track the evolution of different metal-urgical phases as austenite transforms into product phases ofearlite, bainite and martensite as described by Warke et al. (2009).

Fig. 1. Sample fitted in the jaws of a Gleeble machine.

Material data for the mechanical module are obtained from tem-perature dependent tension and compression tests as functions ofphases, temperature, carbon content, strain level and strain ratewhich therefore takes into account work hardening in the differ-ent phases. For low alloyed steels, mechanical data were found tobe mainly dependent on time, temperature and carbon contentas reported by Ferguson et al. (2000). Strains induced by phasetransformation and those by thermal contraction are then calcu-lated (stress–strain response) from the local temperature and phasefractions as a function of time.

In the absence of material data for LCBM rail wheel steels,a set of dilatometry experiments were undertaken to determineand quantify volumetric changes associated with martensite phasetransformation for a number of different cooling rates. Thermalexpansion data and kinetic rate equations from the dilatometryexperiments were incorporated into DANTE in a similar fashion asdescribed in Warke et al. (2009).

Cylindrical hollow LCBM specimens (length of 0.01 m, outer dia.∅5 mm and inner dia. ∅3.5 mm) were prepared from LCBM steelslabs and inserted into the jaws of a Gleeble machine (Fig. 1). Eachspecimen was conditioned in the same manner at the start of thetest to remove residual stresses and to ensure that each specimenhas the same starting microstructure. The specimen was heated at anominal rate of 1 C/s to 920 C, held for 10 min, and then quenchedat the controlled cooling rate to room temperature. Experimentswere repeated with LCBM steels of varying carbon content (0.21%C, 0.15% C, and 0.10% C) and for different cooling rates (0.9 C/s,1.5 C/s, 3 C/s and 9 C/s).

Using DANTE fitting utility, the dilatometry data were used todetermine the kinetic parameters for LCBM steels. Thermal expan-sion data from the dilatometry experiments were incorporated intoDANTE in a similar fashion as described in Warke et al. (2009).Experiments for pearlitic steels were previously undertaken byDANTE and material data are available in DANTE’s proprietarymaterial library.

2.2. Development of ABAQUS/DANTE FE rail wheel model

The finite element model of the rail wheel was developed usingABAQUS 6.7.1 and DANTE 3.3®. A two dimensional model of halfthe cross-section of an as-forged rail wheel was created using 9999

elements and 10182 nodes (Fig. 2(a)). A selection of four noded andthree noded linear asymmetric heat transfer elements, DC4X4 andDC3X3 respectively, have been used in the meshing of the thermalmodel with an acceptable mesh aspect ratio not greater than 2. An

S.N. Lingamanaik, B.K. Chen / Journal of Materials P

Fig. 2. (a) Two-dimensional axisymmetric geometry and mesh of half cross-sectionof as-forged rail wheel showing wheel sections and surfaces. (b) Dilatometry dataf9

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Favourable compressive stresses were predicted on the rim’s sur-

or low carbon bainitic–martensitic steel (0.21% C) at 0.9 C/s, 1.5 C/s, 3 C/s andC/s.

ncremental FE analysis allowed the time–temperature fields fromhe heat transfer analysis to be used as time-dependant loadingsor the mechanical analysis in which the residual stresses follow-ng the quenching process were predicted. The four noded andhree noded bilinear asymmetric stress ABAQUS elements CAX4nd CAX3 respectively were used in the mechanical model analysis.

The boundary conditions and thermo-physical quantities thatave been assumed in the model are listed below:

1) At the start of the analysis, a stress-free state is assumed for therail wheel model as the rail wheel is set to austenization tem-perature of 871 C. At this elevated temperature, differentialwork hardening which may have been formed during forgingis assumed to be significantly relieved and hence not consid-ered at the start of the analysis. Subsequent induced strainsduring the quenching process and associated work harden-ing effects have been built into the constitutive equations andhence accounted for in the prediction of residual stress.

2) Carbon content of rail wheel is assumed uniform across wheeland specified at the start of analysis.

3) Values of heat transfer coefficients for water spray quenchingwere based on laboratory experiments as reported by Khulmanand Gallagher (1988). Convection from wheel to water at thetread region is 3066 W/m2 K and heat loss from wheel to air is28 W/m2 K.

4) Surface emissivity of 0.95 and Stefan–Boltzmann constant of5.67 × 10−8 W/m2 K4.

5) Density of steel was 7.83 × 10−6 kg/mm3.6) Elastic modulus of steel was 1.800 × 105 MPa.

7) Thermal conductivities of individual phases were temper-

ature dependent; austenite = 0.016 + 1.3 × 10−5 (T) W/mm C;martensite = 0.025 + 3 × 10−6 (T) W/mm C.

rocessing Technology 211 (2011) 1547– 1552 1549

(8) Specific heat capacities individual phases were temper-ature dependent; austenite = 370 + 0.298 (T) J/kg C; marten-site = 450 + 0.387 (T) J/kg C.

2.3. Modelling of quenching process for pearlitic steel and lowcarbon bainitic–martensitic steel rail wheels

The combined ABAQUS/DANTE FE model was used to investigatethree different quenching conditions labelled Case A, Case B andCase C.

In Case A, the conventional quenching process is modelled forpearlitic rail wheel AAR Class A (0.55% C). The conventional quench-ing process is modelled in four stages as described by Gordon andPerlman (1998). The tread of the wheel is first quenched for 120 swhile the other sections of the wheel lose heat through convectionand radiation. The rail wheel is left to dwell at room temperaturefor 240 s and then tempered in a 500 C oven for 5 h followed by aircooling to room temperature.

The conditions selected in Case A also allowed the predictivecapability of ABAQUS/DANTE FE model to be assessed and com-pared against experimental data as reported by Mutton and Lynch(2004).

In Case B, the conventional quenching process (as used in CaseA) is applied to LCBM steel rail wheels.

In Case C, an alternative set of quenching conditions (Table 1)was applied to LCBM steels rail wheels to determine if favourablecompressive residual stresses can be achieved in the rim region ofthe wheel. Instead of quenching only the tread of the wheel (asin Case A), the inner hub and the tread were also quenched in thefirst quenching stage. Following the first quenching stage, the treadof the wheel is quenched for 1100 s and then left to cool at roomtemperature. Dwelling and tempering stages have been omitted inCase C.

3. Results and discussion

3.1. Dilatometry results for low carbon bainitic–martensitic steels

Fig. 2(b) shows the dilatometry curves obtained for low carbonbainitic–martensitic steels (0.21% C) for different cooling rates of0.9 C/s, 1.5 C/s, 3 C/s and 9 C/s. As shown in Fig. 2(b), there is asmall shift in martensite start temperature which is expected to bedependent on cooling rate. This behaviour is in agreement with thatobserved for martensite phase transformation for martensitic steelsas reported by Weise and Fritsche (1997). The relative change oflength in the direction of deformation (dL) depends on the degree ofnet volumetric expansion as well as on the fraction of recrystallisedaustenite transforming into martensite as described by Weise andFritsche (1997). Also, the start temperature in martensite transfor-mation in LCBM steels is found to be dependent on carbon content.The martensite phase transformation kinetic equations reflect thisbehaviour being a function of both cooling rate and carbon content.

3.2. Residual stress distribution for pearlitic and low carbonbainitic–martensitic rail wheels

Fig. 3(a) shows the distribution of residual stresses predictedfor AAR Class A (0.55% C) rail wheel under conventional quenchingconditions (Case A). The legend shows the contours of circumfer-ential stresses (negative and positive signs are used to indicatecompressive stresses and tensile stresses respectively in MPa).

face (≈−500 MPa) and below the tread surface (≈−300 MPa). Also,neutral stresses and compressive stresses are predicted in platesection.

1550 S.N. Lingamanaik, B.K. Chen / Journal of Materials Processing Technology 211 (2011) 1547– 1552

Table 1Quenching conditions for pearlitic and LCBM steel rail wheels.

QuenchSpray property,HTC = 3066 W/m2 K

Dwelling Tempering QuenchSpray property,HTC = 500 W/m2 K

Air cooling

Duration (s) Spray location Duration (s) Spray location

Case A (conventional quenching for pearlitic wheels) 120 Tread Yes Yes N/A N/A YesCase B (conventional quenching for LCBM wheels) 120 Tread Yes Yes Yes

ner H

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below the tread of the wheel (Fig. 3(b)). The results predicted for

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Case C (Alternate quenching for LCBM wheels) 120 Tread, In

In Case A, hot austenite in the tread section of the wheel coolsnd transforms to pearlite. The rim of the wheel shrinks and a dif-erential change in volume is created between the inner rim and theread section which results in the inner rim being in tension andhe tread of the rim to be in compression. The final distribution of

esidual stresses predicted by the present model for conventionalearlitic rail wheels is in agreement with finite element analysis ofearlitic rail wheels by Lonsdale and Stone (2002).

ig. 3. (a) Circumferential stress for AAR Class A (0.55% C) rail wheel under conventionarious rail wheel grades as reported by Mutton and Lynch (2004). (c) Circumferential stuenching procedure (Case B). (d) Circumferential stress for low carbon bainitic–martens

ub N/A N/A 1100 Tread Yes

Mutton and Lynch (2004) quantified residual stress distribu-tions in common rail wheel grades using ultrasonic methodsobtained by summing the stresses along the path of the ultrasonicwave and plotted as averaged stress values at different locations

AAR Class A (0.55% C) rail wheels by the present model at variouslocations in the rim have also been averaged (in a similar fash-ion) such that they can be compared against experimental data as

al quenching conditions (Case A). (b) Ultrasonic residual stress measurements onress for low carbon bainitic–martensitic steel (0.21% C) wheels under conventionalitic steel (0.21% C) wheels under alternate quenching procedure (Case C).

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eported by Mutton and Lynch (2004). Fig. 3(b) shows the resid-al stresses predicted by the ABAQUS/DANTE FE model are in goodgreement with experimental residual stress measurements in railheels.

Fig. 3(c) shows the residual stresses predicted for LCBM railheels under conventional quenching procedure (Case B). High

ensile residual stresses (>550 MPa) were predicted in the rim ofCBM steel rail wheel which are probably high enough to causeracking as the value is of the order of yield stress of LCBM steel≈860 MPa).

During quenching of LCBM rail wheels (Case B), the rim shrinksue to thermal contraction but experiences a net outward expan-ion due to the martensite phase transforming in the rim, plate andub of the wheel. This behaviour is expected in LCBM steels due to

net volumetric expansion for martensite phase transformation.artensite transformation occupies a higher net volumetric expan-

ion compared to pearlite transformation in conventional pearliticail wheels as reported by Lonsdale and Stone (2002).

These results predicted in Case B have been found to be consis-ent with results of experimental work undertaken by Lonsdale andshler (2001) on 420 monobloc rail wheels which were quenchedt the tread for 240 s and tempered for 4 h. Cuts made in the rim ofhe wheel showed a net 0.4 mm (positive) opening displacement inhe tread’s surface which suggests that significant tensile residualtresses were formed due to quenching in the rim of the wheel.

Quenching conditions (quenching durations, spray locationsnd spray configuration) for LCBM steel rail wheels selected in Case

were shown to result in favourable compressive residual stresses≈−300 MPa) in the rim of the wheel. Tensile residual stresses rang-ng from 70 MPa to 200 MPa were predicted in the plate of the

heel. However, they are not detrimental since cracks tend to formainly in the outer rim of the wheel.As described in Case B, the hub, plate and the inner rim con-

ribute to the net outward expansion in the tread of the wheelfter the tread has been quenched. By quenching the inner hub inase C, martensite is expected first to transform in the hub and thelate and changes the sequence of martensite phase transformationhich alters the resultant residual stress distribution.

Therefore changes to quenching positions on the wheel andhanges to quenching intensities were found to significantlyhange the residual stress distribution in the rim of LCBM railheels. The level of compressive residual stresses in the rim

f LCBM steel rail wheels was found to increase substantiallyy selecting appropriate quenching conditions as those used inase C.

Khulman and Gallagher (1988) have reported averaged val-es for heat transfer coefficients based on experimental work.owever, heat transfer coefficients generally follow a non-linear

elationship with part’s surface temperature due to the variousoiling stages during quenching and the formation of differentegimes which depends on a number of factors such as part’s tem-erature, part’s geometry, quenching configuration and quenchingedium’s temperature as shown by Schwalm and Tensi (1981).

herefore, pertinent parameters involved in the quenching processuch as heat transfer coefficients need to be validated against exper-mental data to confirm the assumed values used in the FE model.n experimental quenching rig is in its final stage of constructionnd will be fitted with thermocouples and real-time data acquisi-ion devices to provide temperature measurements to verify heatransfer coefficients used in the computational modelling.

Another issue that needs to be considered is the effect of machin-ng of the as-forged quenched rail wheel on the final residual stressistribution. A surface layer of up to 40 mm in thickness (in the

ub) is removed by machining to achieve an accurate dimensionalail wheel profile. Consequently, the residual stress distribution fol-owing quenching of rail wheels is believed to change in which the

rocessing Technology 211 (2011) 1547– 1552 1551

residual compressive stresses in the rim of the wheel are expectedto relax. Further work is being undertaken to evaluate more pre-cisely the changes to the residual stress distribution brought aboutby the machining process.

4. Conclusion

Dilatometry experiments for different grades of LCBM steelswere undertaken to quantify the volumetric changes occurringduring martensite phase transformation and the results were incor-porated in an ABAQUS/DANTE FE model. The residual stressespredicted by ABAQUS/DANTE FE model were found to be inagreement with published experimental results. Conventionalquenching process for as-forged pearlitic steel rail wheels resultedin favourable compressive residual stresses being formed in therim of the wheel whereas the conventional quenching process wasfound to be unsuitable for LCBM steel rail wheels due to the for-mation of high levels of tensile stresses in the rim of the wheel.However, this work has demonstrated that altering the quench-ing parameters (heat transfer coefficients, quenching duration andquenching locations) can promote favourable compressive residualstresses in the rim of LCBM rail wheels.

Acknowledgements

The authors would like to thank CRC for Rail Innovation(established and supported under the Australian Government’sCooperative Research Centres program) for the funding of thisresearch. Project No. R3.101 New Wheel Steel and also Tim Con-stable for supporting this work.

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