simulacion y evaluacion fatiga termica

7/31/2019 Simulacion y Evaluacion Fatiga Termica

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International Journal of Fatigue 26 (2004) 10951107www.elsevier.com/locate/ijfatigue

Simulation and evaluation of thermal fatigue crackingof hot work tool steels

Anders Persson a,, Sture Hogmark b, Jens Bergstrom a

a Department of Materials Engineering, Karlstad University, SE-651 88 Karlstad, Swedenb The Gngstrom Laboratory, Uppsala University, SE-751 21 Uppsala, Sweden

Received 30 June 2003; received in revised form 2 February 2004; accepted 3 March 2004

Abstract

Die casting is a very cost efficient method to manufacture near net-shaped and complex cast products. One limitation for fur-ther cost reduction is fatigue cracking of the tool due to thermal cycling, which is observed as a crack network on the tool sur-face. Hot work tool steels are commonly used as die material.

In this study, an experimental test machine for simulation of thermal fatigue is described. The test is based on cyclic inductionheating and internal cooling of hollow cylindrical test rods. The surface strain is continuously recorded during the thermal cyclingthrough a non-contact laser speckle technique. The applicability of the test is demonstrated on two hot work tool steel grades,hardened and tempered to different conditions, and heat cycled between Tmin 170

v

C and Tmax 600850v

C.It is shown that the test method can simulate surface cracking of tools exposed to thermal fatigue. The surface strain recordings

proved to give sufficient information to successfully deduce the strains and stresses behind the mechanism of thermal fatigue sur-face cracking, without knowledge of the temperature distribution below the surface. It was also found that low-cycle fatigueoccurs for the tests with Tmax 600 and 700

v

C, although the estimated tensile stress after cooling does not exceed the initial yieldstrength of the steel. Most probably, the reason is the gradual softening of the tool steels during the thermal cycling. Addition-ally, the presence of stress concentrators play a critical role during these conditions.# 2004 Elsevier Ltd. All rights reserved.

Keywords: Thermal fatigue; Heat checking; Surface strain; Hot work tool steel; Die casting

1. Introduction

Die casting is a very cost-efficient method of formingnear net-shaped cast products of, for example, alu-minium, zinc, magnesium, and copper based alloys ofalmost any shape [13]. Prior to casting aluminium and

copper alloys, the die is normally preheated to a tem-perature within the range of 250300 and 300350

v

C,respectively, to reduce the thermal shock, and the aver-age tool temperature is usually kept at those levelsthrough internal cooling. During a casting cycle, mol-ten metal is forced into the mould by the application ofpressure, the peak of which can exceed 70 MPa. A dis-tinguishing characteristic of the process is that the

liquid metal flows with high velocity during injectionand provides rapid filling of the die cavity, typicallywithin milliseconds. For aluminium alloys, the entrancevelocity of the melt is usually 2060 m/s and the melttemperature is approximately 700

v

C, whereas those forcopper alloys is about 110 m/s and 970

v

C. The high

melt velocity is necessary to completely fill the mouldof thin-walled and complex shaped products. Continu-ous internal cooling of the tool during the processmakes the solidification of the casting efficient, andhigh rate manufacturing of typically 100 castings perhour is possible. When the casting has solidified, the dieis opened and the casting ejected. Thereafter, the diemay be externally cooled and lubricated by spraying.

Thermal fatigue cracking, gross fracture, erosion,corrosion and local adherence of the casting alloy (sol-dering) are important life-limiting tool failure mechan-isms in aluminium and brass die casting [13]. Thermal

Corresponding author. Tel.: +46-54-700-18-21; fax: +46-54-700-14-49.

E-mail address: [email protected] (A. Persson).

0142-1123/$ - see front matter # 2004 Elsevier Ltd. All rights reserved.doi:10.1016/j.ijfatigue.2004.03.005


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fatigue cracking results from rapid alternations in tem-perature of the die surface during the casting process.The temperature cycling may induce stresses highenough to impose an increment of plastic strain in thetool surface during each casting cycle. Surface cracksdevelop generally within a few thousand cycles, or even

earlier, and are, consequently, formed in the low-cyclefatigue range (


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of sensors give the axial surface strain and the horizon-tal pair the tangential. The speckle displacements areacquired with a sampling rate of 500 Hz.

2.2. Materials

Two Uddeholm hot work tool steels, QRO 90Supreme and Hotvar, were used as test materials. QRO90 has a nominal chemical composition (wt.%) of 0.38C, 0.30 Si, 0.75 Mn, 2.6 Cr, 2.25 Mo, 0.9 V, and Febalance. Hotvar has a nominal chemical composition(wt.%) of 0.55 C, 1.0 Si, 0.75 Mn, 2.6 Cr, 2.25 Mo,0.85 V, and Fe balance. For the QRO 90 specimens,two variants of heat treatment were used: hardening(austenitizing 30 min at 1030

v

C, followed by airquenching) and tempering 2 2 h at 640

v

C to a hard-ness of 430 10 HV30 and hardening as above fol-

lowed by tempering 2 2 h at 625v

C to a hardness of510 10 HV30, respectively. The Hotvar specimens

were hardened (austenitizing 30 min at 1050v

C, fol-lowed by air quenching) and tempered 2 2 h at 575v

C to a hardness of 640 10 HV30. All heat treatmentsresulted in microstructures of tempered martensite. Thehardness of the steels was assessed by Vickers indenta-tions on polished cross-sections, using a load of 30 kg.For all materials, the austenitizing treatment gives anominal austenite grain size of about ASTM 9, and nosignificant microstructural differences between thematerials could be detected. The heat treatment of thespecimens was followed by grinding to a surface rough-ness (Ra) of 0:38 0:05 lm, as obtained using optical

surface profilometry.Relevant mechanical properties of the two tool steels

are given in Fig. 2. The true yield strength values(Fig. 2f) are obtained from visual inspections of tensilecurves such as that of Fig. 2e, and are defined as thestresses when the stressstrain curves deviate from alinear relation. Note that the true yield stress values arewell below Rp0.2 in Fig. 2b,e.

2.3. Thermal fatigue testing

Three temperature cycles were used to simulate vari-ous die casting conditions (see Table 1), and designated

according to their maximum temperature. The maxi-mum temperatures for the three temperature cycleswere set to 600, 700, and 850

v

C, respectively, toinclude the thermal conditions for aluminium and brassdie casting, respectively. The latter temperature corre-sponds to the maximum tool surface temperature dur-ing actual brass die casting [13]. The temperature cyclesincluded a steep ramp to the maximum temperature,no holding time at maximum temperature, and sub-sequent cooling to the minimum temperature. Continu-ous cooling was performed by circulating silicon oil(flow rate % 2.5 l/min) of 60

v

C through the specimen,

and also externally with either argon (forced convec-tion) or air (natural convection).

Argon was used as cooling medium because the toolsare exposed to an environment with reduced oxygencontent during actual die casting. The oxygen in the diecavity is partly consumed through oxidation of toolmaterial and casting alloy. For comparison, some testswere performed in air.

Prior to testing, the specimens were pre-oxidised toobtain a thin oxide layer, which facilitates the pyrom-eter temperature control during heating. This wasmade by electrochemical oxidation in a NaOH-solution(containing deionized water and 300 g/l NaOH) at70

v

C for about 5 min, followed by 1 h heat treatmentat 200

v

C in air. A K-type ChromelAlumel thermo-couple with thin wires (1 0.13 mm) was spot welded tothe specimen to measure the surface temperature dur-ing testing. The thin wires enable rapid response of anychange in temperature. Finally, to obtain a good

speckle pattern for the surface strain measurements, anarea of approximately 10 10 mm located in the mid-dle of the specimens was roughened by a 1000 meshabrasive paper.

2.4. Evaluation techniques

During the heat cycling, the surface strain is continu-ously obtained by the laser speckle technique from thechange in the specimen dimensions, and represented assurface strain vs. temperature. Any thermal fatiguedamage of the specimen surface was revealed using

scanning electron microscopy (SEM). It was furthercharacterised with respect to crack growth (cracklength vs. number of cycles) and crack density (numberof cracks per unit of length along the specimen surface)by measurements on polished axial cross-sections per-formed in light optical microscopy (LOM). For eachspecimen, all evaluation of cracks is based on crackslonger than about 5 lm, detected along two lines, eachof 8 mm length.

Profiles of surface hardness and hardness vs. depthafter exposure to the heat cycling were assessed byVickers indentations on polished specimen surfaces andcross-sections, respectively, using a load of 25 g.

Finally, the residual stress state in the surface layerwas measured by X-ray diffraction (XRD) using CuKaradiation and the sin2 w method.

2.5. Tests for verification of the strain measurements

To investigate if the surface strains obtained by thelaser speckle arrangement correlate with the actual sur-face deformation, a cylindrical pressure vessel of alu-minium (diameter 100 mm, wall thickness 2 mm, andlength 80 mm) was selected. The vessel was designedand positioned in such a way that its external surface

A. Persson et al. / International Journal of Fatigue 26 (2004) 10951107 1097


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was kept in the same position as that of an ordinary

thermal fatigue specimen. To achieve a good specklepattern for the speckle measurements, the surface were

roughened as above. Consequently, the speckle move-

ments during deformation of the pressure vessel were

detected by the CCD-sensors under equal optical con-

ditions as those during heat cycling of the thermal fati-gue specimens. A strain gauge was glued to the surface,

next to the spot illuminated by the laser beam, to

obtain reference values of the surface strain.

Table 1Thermal cycles used in the thermal fatigue tests

Max. temperature [v

C] Min. temperature [v

C] Heating time [s] Total cycle time [s] External cooling

600 170 0.2 11.2 Argon or air700 170 0.3 14.3 Argon or air850 170 2.2 26.2 Argon or air

Fig. 2. Nominal mechanical properties of Hotvar and QRO 90 hardened as above and tempered to various conditions. (a) Temper resistance ashardness at room temperature vs. holding time for temperatures within the range of 550650

v

C. (b) Hot yield strength as Rp0.2 vs. temperature.(c) Modulus of elasticity vs. temperature. (d) Hot ductility as reduction of area at tensile fracture vs. temperature. (e) Schematic of sample tensilecurve. (f) Estimated true yield strength ry vs. temperature. The solid lines are the second order polynomial fit to the data.

1098 A. Persson et al. / International Journal of Fatigue 26 (2004) 10951107


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The design of the pressure vessel enabled defor-mation under well-controlled conditions and simul-taneous strain measurements by the laser speckletechnique and the strain gauge, respectively. For thetwo methods, the tangential surface strain (hoop strain)after loading or unloading of the pressure vessel wasused for the evaluation, since the induced deformationis significantly larger in the tangential than in the axialdirection. The results from the two techniques werecompared.

3. Results

3.1. Verification of the laser speckle technique

By pressuring the aluminium vessel, it proved poss-ible to verify that the surface strains obtained by thelaser speckle technique correlates very well to the refer-

ence values measured by the strain gauge (see Fig. 3).It is concluded that the laser speckle technique makes itpossible to detect surface strains with a resolution ofabout 25 10-6. Because of security reasons, it wasnecessary to limit the maximum pressure exposure ofthe pressure vessel, which as a result, restricted themaximum and minimum surface strains of this verifi-

cation to about 160 106.

3.2. Recorded surface temperatures

Induction heating very rapidly increases the surfacetemperature, as seen in Fig. 4a. After the short heatingtime, the surface will cool down following a more grad-ual slope. The three heat cycles tested all start with aspecimen temperature of 60

v

C (= oil temperature). Ittakes about four cycles to obtain equilibrium tempera-ture conditions (see Fig. 4b and Table 1). No signifi-

cant differences could be seen among the differentsteels.

3.3. Surface strain during thermal cycling

Typically, the surface strain increases with the sur-

face temperature, followed by an almost constant or a

slightly increasing strain level during the first steep part

of the cooling phase, whereafter it decreases with tem-

perature (see Fig. 5). As expected, the surface strain

level was strongly dependent on the maximum tem-

perature during each thermal cycle (see Fig. 5). During

heating to 850 vC, there is a sharp increase in the sur-face strain curve at about 760

v

C to a slope that is sig-

nificantly higher as compared to that at any lower

temperature (see Fig. 5c). After the cooling event and

at equilibrium temperature conditions, each surface

strain recording forms either a closed loop with practi-

cally no residual surface strain at the minimum

temperature (see Fig. 5a (Tmax 600v

C) and b

(Tmax 700v

C), or an open loop with tensile residual

surface strain at the minimum temperature (see Fig. 5c;

Tmax 850v

C). Note also that the tangential strain e/

Fig. 3. Comparison of tangential surface strain (hoop strain)obtained by the laser speckle technique and the resistance straingauge, respectively, after pressuring or depressurising the aluminiumvessel. Each circle represents the result from one loading (positivevalues) or unloading (negative values) event. The solid line is the lin-ear fit to the experimental data.

Fig. 4. Typical surface temperature recording for the 700v

C test atequilibrium (a), and of maximum temperature Tmax and minimumtemperature Tmin during the first 10 cycles extracted from the tem-perature recordings (b).



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and axial strain ez

are almost identical, and in the fol-

lowing only e/ is considered.

3.4. Surface cracking after thermal fatigue

3.4.1. Argon atmosphereThermal fatigue damage was typically observed as a

network of cracks on the specimen surface (see Fig. 6).

Polished cross-sections revealed that the appearance of

the cracks was strongly dependent on the maximum

temperature during each cycle. For thermal cycling up

to 700v

C, the crack path was relatively straight

(Fig. 7a), whereas branched cracks (Fig. 7b) in

addition to the single cracks were observed after heatcycling to 850

v

C.The crack length was strongly dependent on the

number of cycles and the maximum temperature dur-

ing each cycle (see Fig. 8). Since the crack propagation

rate was very high during the 700 and 850v

C cycles, as

compared to the 600v

C cycles, it was necessary to limit

Fig. 5. Typical surface strain recordings of QRO 90 at 510 HV30during thermal cycling in air at equilibrium temperature conditions to600

v

C (a), 700v

C (b), and 850v

C (c).

Fig. 6. Overview of typical crack network after 10 000 thermalcycles to 700

v

C in argon (SEM). The axial direction of the specimenis horizontal.

Fig. 7. Polished cross-section revealing typical crack after thermalcycling in argon (LOM). (a) 10 000 cycles to 700

v

C. (b) 1 000 cyclesto 850

v

C (QRO 90 at 510 HV30).



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the number of cycles to 10 000 and 1000 cycles,

respectively, for these temperature levels. In a relatively

early stage, the density of cracks was also strongly

dependent on the number of cycles and the maximum

temperature level. Thereafter, the crack density aver-

aged at almost constant levels. With few exceptions,

the crack length and crack density seem to have a tend-

ency to decrease with increasing hardness. Note also

that no cracks exceeding the evaluation criteria of

about 5 lm were observed on Hotvar after 100 cyclesto 850

v

C.

3.4.2. Air atmosphereThe crack length and crack density after 100 cycles

to 850v

C and 1000 cycles to 700v

C, respectively, inargon and air atmosphere were comparable.

Thermal cycling in air (1000 cycles to 700v

C)revealed that the crack length and crack density in gen-eral decrease with increasing hardness, see Fig. 9. Forall materials, no cracks larger than the evaluation cri-teria of about 5 lm were detected after 100 cycles to850

v

C.

3.5. Hardness after thermal fatigue

As expected, the surface hardness decreased with thenumber of thermal cycles for all materials and con-ditions, as exemplified by Fig. 10. It is seen that thesurface hardness is rapidly reduced initially, followedby a more gradual loss of hardness with the number ofcycles. The hardness levels and the thickness of the sof-tened surface layer were strongly dependent on thenumber of heat cycles and the maximum temperatureduring each cycle (see Fig. 11). Evidently, afterexposure to 600

v

C there is no notable effect on thehardness levels (Fig. 11a), whereas a considerablesoftening of the surface layer is detected after thermal

cycling to 700v

C (Fig. 11b) and 850v

C (Fig. 11c),respectively. For the 700

v

C test, the hardness decrea-ses gradually from a depth of about 0.5 mm towardsthe surface (Fig. 11b). Thermal cycling at 850

v

C resul-ted in a dramatic hardness reduction throughout thewhole specimen (see Fig. 11c), which also reveals ahardness maximum at about 0.3 mm depth for allsteels. Finally, the softening rate and the thickness ofthe softened surface layer is larger for the Hotvar steelthan for the QRO 90 steel, so that the hardness levelsgradually evens for increasing number of heat expo-sures (see Figs. 10 and 11b,c).

Fig. 9. Length and density of thermal fatigue cracks after 1000cycles to 700

v

C in air. One specimen of each material was tested.

Fig. 8. Length and density of thermal fatigue cracks after heat cyc-ling in argon atmosphere. (a) Maximum crack length vs. number ofcycles. (b) Mean crack length vs. number of cycles. (c) Crack densityvs. number of cycles. Each pile is the mean value of three or four speci-mens, except those for 100 cycles to 850

v

C and 1000 cycles to 700v

C,respectively, which are based on one specimen of each material. Theerror bars indicate the maximum and minimum recording.



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3.6. Residual stresses

Residual tensile stresses are gradually building up inthe surface layer during the thermal cycling, whereasthe residual shear stress is virtually unaffected (seeFig. 12). It is seen that the residual tangential and axialstresses increase rapidly from a compressive to a tensilestate, followed by a more gradual increase in tensilestresses with the number of heat exposures. Note thatthe tangential and axial stresses follow the same tend-ency. For the 600

v

C tests, the tensile stress state isreached after about 10 heat cycles, whereas that forthe 700

v

C experiments is obtained immediately after

the first cycle. Generally, the levels of the tensile stres-ses are higher after the 700

v

C tests than after those of600

v

C. The maximum scatter in residual stress at allnumber of cycles, except at the first where it is highest(about 6070%), is about 1040%.

4. Discussion

The laser speckle technique was shown to give thetrue surface strains (see Fig. 3 and Ref. [14]), and couldthus be used to obtain information necessary to verify

the mechanisms behind thermal fatigue. Please notethat this is possible just by recording the surface tem-perature and strain. Knowledge of the temperatureprofile below the surface is not needed!

4.1. Characteristics of the induction heating

Induction heating of steel using a frequency of3 MHz results in very fast heating of only a thin sur-face layer, of the order of 10 lm (skin-effect). Whenapplying this technique to the tool steel test rods, thethermal expansion of the surface material is retained by

Fig. 11. Hardness vs. depth after exposure to thermal cycling. (a) 20000 cycles to 600

v

C. (b) 10 000 cycles to 700v

C. (c) 1000 cycles to850

v

C.

Fig. 10. Surface hardness vs. number of heat cycles to 700v

C. (a)For the first 5000 cycles. (b) Close-up of the first 100 cycles (note thelinear scale for the number of cycles).



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the cooler bulk material (see Fig. 5). During the initialphase of the cooling, the surface contracts but the bulkmaterial still expands due to heat conduction. Thereby,the decrease in surface strain with temperature isdelayed (see Fig. 5).

When heating to 850v

C, the heating rate is dramati-cally reduced above the Curie temperature of Fe(768

v

C) (see Table 1), as a result of the change in themagnetic properties of the material. This promotes thetemperature distribution through the specimen to evenout by heat conduction. Hence, thermal cycling byinduction heating above the Curie temperature does

not ideally generate the temperature profile representa-tive for die casting of e.g. brass [13]. Consequently, the850

v

C test in this investigation is not as representativefor die casting as those of 600 and 700

v

C, and is notincluded in the strain estimations below.

4.2. Mechanical surface conditions during thermalcycling

4.2.1. Deduction of mechanical surface strainSince the thermal strain of the surface layer is con-

strained during the thermal cycling, the surface

material will be exposed to cyclic stresses. The hypo-

thetical strains corresponding to these stresses aredefined as mechanical strains Emech. Crack nucleationand growth during thermal cycling is determined byfluctuations in Emech. Similarly, we define thermalstrains Eth from the thermal expansion coefficient of thetool material a(T) and the minimum temperatures Tminwithout any constraint as:

ethT aTT Tmin 2

In the calculations, a(T) is the nominal value for thesteels as the mean value between room temperatureand the temperature of interest.

Thus, emech at any part of the thermal cycle is poss-ible to deduce from the corresponding values of thesurface temperature cycle, thermal expansion coefficientof the tool material, and the measured surface strainEtot according to Ref. [15]:

emechT etotT ethT 3

Obviously, knowledge of the temperature distributionbelow the surface is not necessary.

The three types of surface strain during thermal cyc-ling can be deduced from results such as those ofFigs. 4 and 5, and using Eqs. (2) and (3) (see Fig. 13).Note that the surface layer is not totally constrained

Fig. 13. Typical surface strain response during thermal cycling(QRO 90 at 510 HV30, equilibrium temperature conditions). (a) Heatcycling to 600

v

C. (b) Heat cycling to 700v

C.Fig. 12. Residual tangential stress ru, axial stress rz, and shearstress s, respectively, in the surface layer vs. number of thermal cycles(QRO 90 at 510 HV30). (a) Heat cycling to 600

v

C. (b) Heat cyclingto 700

v

C.



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during the heat cycling because Etot is not representedby perfectly horizontal lines. The situation shouldrather be considered as a partially constrained case.

From mechanical surface strain loops such as thoseof Fig. 13, the minimum mechanical surface strain Eminand any residual mechanical surface strain at the mini-mum temperature eresidual of each cycle could be extrac-ted (see Fig. 14). For all tests, these two strains seem tobe independent of the number of thermal cycles. Forthe tests at 600 and 700

v

C, eresidual was about zero forall cycles, and emin was about 0.47% and 0.56%,respectively, for the two temperature levels at all equi-librium cycles. No differences were seen between thetested steels. These values of the mechanical surfacestrain can be used for numerical simulations andexperimental tests of thermal fatigue processes.

4.2.2. Estimated surface stress and risk of plastic strainSimplified, the elastic stress r(T) in the surface layer

at any temperature within the thermal cycles of the testrods can be estimated by Hookes law in plane stressusing the mechanical surface strain emech from thisinvestigation and Poissons ratio m(T) and modulus ofelasticity E(T) of the tool materials (isotropic and elas-ticideal-plastic) as:

rT ETemechT

1 mT4

Starting with a stress-free state at cyclic equilibriumtemperature and solving Eq. (4) with mT 0:3, andE(T) and emech(T) according to Figs. 2c and 13b,respectively, the stress on the surface layer during twothermal cycles varies as shown in Fig. 15. It is seen thatduring the first cycle, the compressive surface stressr(T) increases with temperature, and the surfacematerial behaves elastically until the stress reaches thetrue yield strength of the material ry(T) in compression(see Fig. 2f). Thus, the critical surface temperature Tcduring heating at which the material begins to deformplastically is the temperature when r(T) equals the

compressive yield strength ry(T). For QRO 90 at

510HV30 and Hotvar at 640 HV30, Tc was found to be

approximately 420 and 460v

C, respectively. There-

after, the surface material accumulates compressive

plastic strain until the maximum cycle temperature is

reached. The maximum plastic strain Ep is obtained at

Tmax and can be expressed by Eq. (5) below:

ep emechTmax emechTc

emechTmax ryTc1 mT

ETc5

During this plastic strain the stress level is determ-

ined by the compressive yield strength curve. During

the subsequent cooling, the surface stress is reversed

towards the tensile direction by the thermal contrac-

tion. It is seen in Fig. 15 that the developing tensile

stress r(T) during the cooling part of the cycle never

reaches the corresponding ry(T). The magnitude of the

tensile residual stress rresidual at Tmin 170v

C equals

the maximum rresidual and occurs at Tmin (Fig. 15). It is

Fig. 14. Mechanical surface strain vs. number of thermal cycles dur-ing heat cycling to 600 and 700

v

C, respectively (Hotvar 640 HV30).

Fig. 15. Surface stress vs. time during thermal cycling to 700v

C atequilibrium temperature conditions (QRO 90 at 510 HV30). r(T) isestimated using Eq. (4) and the true yield stress ry(T) is according toFig. 2f. The conventional yield strength Rp0.2 according to Fig. 2b,e isincluded for reference. (a) Two whole cycles. The residual stressrresidual after the first cycle is defined. The arrows indicate the devel-opment ofr(T). (b) Close-up of the first 10 s of the first thermal cycle(linear-logarithmic scale).



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deduced by Eq. (6) below:

rresidual ryTmax ETmaxemechTmax

1 mT6

Consequently, the residual tensile stress at Tmin is sim-ply the difference between the elastic stress that thematerial would have experienced at Tmax if there wasno plastic deformation (r(Tmax), Eq. (4)) and the yieldstress at Tmax.

Solving Eqs. (5) and (6) with m(T) and E(T) as above,and Emech(T) and ry(T) according to Figs. 13 and 2f,respectively, the values shown in Table 2 on the plasticsurface strain after the first thermal cycle and theresidual stress in the surface layer after each cycle wereobtained. Due to lack of true yield stress values (suchas those in Fig. 2f) for QRO 90 at 430 HV30, no esti-mations were made on this material. However, theplastic surface strain magnitude and the tensile residualstress in the surface layer is expected to be even higherfor this steel. Generally, the magnitude of the plasticstrain and the residual tensile stress increases morethan proportional to the maximum surface temperature

level, since the yield strength decreases at the same timeas r(Tmax) increases with temperature.

During the following thermal cycle, the maximumcompressive stress on the surface material is equal tothe compressive yield strength and, therefore, thematerial behaves elastic during the whole cycle. Thereason is the presence of residual tensile surface stress,which at the end of each cycle, except for the first one,is equal to that of the previous one. Consequently, thesurface material is already exposed to the worst mech-anical conditions during the first cycle.

On the other hand, if the surface material only

would experience elastic deformation during the firstcycle, there would be zero residual stress at the end ofthe first as well as the following cycles.

As seen in Fig. 15 and by comparing the data inTable 2 with Fig. 2f, the calculated residual stressesafter the 600 and 700

v

C cycles do not exceed the trueyield stress of the tool materials at 170

v

C. In spite ofthis fact, numerous cracks appear during the very firstfew cycles. However, the stress estimations are basedon simplified elasticideal-plastic materials and theinformation given in Fig. 2. Any thermally inducedalternations in yield stress or any presence of stress

concentrators has not been accounted for. Nor is thewell-known effect that ry increases with strain rate con-sidered. Locally, due to surface irregularities, the ten-sile stress during cooling may well exceed the yieldstress. It is indicated in Figs. 10 and 11 that the yieldstress values measured prior to testing also degrade sig-nificantly during thermal cycling, causing furtheraccumulation of plastic strain. This is confirmed by themeasured residual stresses after temperature cycling to600 and 700

v

C, respectively (see Fig. 12). Conse-quently, it is reasonable to believe that the surfacematerial will experience tensile stresses locally thatexceed the yield stress during thermal cycling to 600850

v

C after a certain number of cycles. This is sup-ported by the observation that thermal cracks areformed and propagate within the low-cycle fatiguerange (see Figs. 8 and 9).

However, from Fig. 12 it is evident that the surfacelayer material initially exhibits a compressive residual

stress of the order of 500 MPa. This stress should besuperimposed to the initial elastic stress during heatingas given in Fig. 15. The consequence is that the surfacematerial will reach the compressive yield stress earlierduring the first cycle, and accumulate a larger amountof compressive plastic strain until the maximum cycletemperature is reached. However, the residual stressafter the first and consecutive cycles is not influenced.It is also obvious from Fig. 12 that the magnitude ofthe measured residual stresses is lower than calculatedones (see Table 2). The reason, except those mentioned,is the fact that the X-ray measurements give average

values of the stress within a surface layer of a fewmicrometres. The stress in the superficial surface layerwhich was calculated is probably much higher. Thelimited depth resolution of XRD also explains theobservation that no tensile residual stresses areobserved after the first 600

v

C cycle (see Fig. 12a).Similarly, if the surface layer initially had been in a

tensile residual stress state, the sign and magnitude ofthe residual stress after the first heat cycle could alsobeen given by Table 2, provided that the initial tensilestress did not completely prevent plastic deformationduring heating.

The conclusion is that an initial tensile residual stress

state above a certain magnitude in the tool materialsurface layer may delay crack initiation, whereas initialcompressive residual stresses of any level would facili-tate crack formation.

In the above calculations, it was assumed that thespecimens were preheated to the minimum equilibriumtemperature of about 170

v

C and that steady-state tem-perature conditions prevailed from the first cycle itself.However, the initial specimen temperature is 60

v

C,and from Fig. 4b it is obvious that it takes about fourcycles to obtain cyclic equilibrium temperature con-ditions with constant minimum and maximum surface

Table 2Estimated plastic strains [%] and residual stresses [MPa] after thermalcycling with preheating (Tmin 170

v

C)

Material ep=rresidual

Tmax 600v

C Tmax 700v

C

QRO 90 at 510 HV30 0.20/600 0.30/880

Hotvar at 640 HV30 0.15/500 0.26/830



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temperature. For the 600 and 700v

C tests, these con-ditions are obtained with an approximately constanttemperature range (TmaxTmin), whereas the range forthe 850

v

C experiments decreases slightly during thefirst cycles.

The fact that the mechanical surface strain is relatedto the temperature range (see Figs. 13 and 4b), impliesthat emech(T) for the first 600 and 700

v

C cycle shouldbe approximately equal to that at equilibrium con-

ditions. Solving Eq. (4) as above and using the tem-perature values for the first cycle, the critical surfacetemperature for plastic compression Tc was found to beapproximately 330 and 385

v

C, respectively, for QRO90 at 510HV30 and Hotvar at 640 HV30. SolvingEqs. (5) and (6) as above and using Tmax according toFig. 4b, the values given in Table 3 on the plastic sur-face strain and the residual stress in the surface layerafter the first thermal cycle without preheating (initial

temperature 60v

C) were obtained. Again, no estima-tions on QRO 90 at 430 HV30 could be made.

Obviously, preheating of the specimens to the mini-mum equilibrium temperature has a tendency to makethe mechanical conditions during the first cycle worsethan without preheating. However, since the cyclicequilibrium temperature conditions are obtained withinthe first few cycles (see Fig. 4b), there would probablynot be any notable effect on the thermal cracking.

4.3. Thermal fatigue cracking

When exposing the tested tool steels to thermal cyc-ling to 600850

v

C, it takes the order of


13/13

. The surface strain recordings offered by the laserspeckle technique, proved to yield sufficient infor-mation for calculation of surface strains and stressesresponsible for the thermal fatigue of the surfacelayer, without knowledge of the temperature profilebelow the surface.

. With the tool steels tested, low-cycle fatigue occursfor the tests with Tmax 600 and 700

v

C even thoughthe calculated tensile stress after cooling does notexceed the yield stress that the material exhibitsprior to testing. The reason is the gradual softeningof the tool steels during the thermal cycling, and thepresence of stress concentrators.

. Preheating of the tool does not improve the resist-ance to thermal fatigue cracking.

. An initial tensile residual stress state of a certainmagnitude in the tool material surface may delaycrack initiation, whereas an initial compressiveresidual stress of any level facilitates crack forma-

tion.. The hardness ranking between the materials was

maintained throughout the tests, even though allsteels suffered a considerable softening.

. Unfortunately, the induction induced heating to 850v

C is disturbed by the Curie transition at 768v

C,which dramatically reduces the heating rate abovethat temperature. Thus, the results generated cannotbe directly transferred to the thermal operationsaimed at.

Acknowledgements

The authors are grateful to Uddeholm Tooling AB,Tour & Andersson AB and Bodycote Heat TreatmentAB. The financial support from the Swedish Knowl-edge Foundation is also acknowledged. Special thanksto Mr Lars Carlsson at Karlstad University for assist-ing at the verification tests of the laser speckle arrange-ment.

References

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[2] Davis JR, editor. ASM speciality handbook, tool materials.Materials Park OH: ASM International; 1995. p. 251.

[3] Allsop DF, Kennedy D. Pressure diecasting, Part 2: the tech-nology of the casting and the die. Oxford: Pergamon Press Ltd;1983.

[4] Danzer R, Sturm F, Schindler A, Zleppnig W. Thermal fatiguecracks in pressure die casting dies. Gisserei-Praxis 1983;19(20):287.

[5] Kovrigin VA, Starokozhev BS, Yurasov SA. Effect of oxidizingprocesses on crazing of die-casting molds. Met Sci Heat Treat1980;22:688.

[6] Nehrenberg AE. Metallurgical aspects of heat checking in brassdie casting dies. Die Castings 1949; Oct/Nov: 30.

[7] Zleppnig W, Danzer R, Fischer FD, Maurer KL. Influence ofthe structure and the temperaturefield on the formation andpropagation of thermal fatigue cracks. Proceedings of the 6thBiennial European Conference on Fracture, Amsterdam, 1986.p. 1139.

[8] Schindler AM, Danzer RB. Advanced materials for die castingand extrusion applicationdie life in die casting. Proceedings ofthe International Conference Step into the 90s, 3. Queensland,1986. p. 905.

[9] Yamaguchi I. A laser-speckle strain gauge. J Phys E Sci Instrum

1981;14:1270.[10] Yamaguchi I. Speckle displacement and decorrelation in the dif-

fraction and image fields for small object deformation. Opt Acta1981;28:1359.

[11] Yamaguchi I, Takemori T, Kobayashi K. Stabilized and acceler-ated speckle strain gauge. Opt Eng 1993;32:618.

[12] Yamaguchi I. Encoder and strain gauge using laser speckle. OptLasers Eng 1989;11:223.

[13] Persson A, Hogmark S, Bergstrom J. Temperature profiles andconditions for thermal fatigue cracking in brass die casting.J Mater Process Technol [in press]

[14] Velay V, Persson A, Bernhart G, Bergstrom J, Penazzi L. Ther-mal fatigue of a tool steel: experiment and numerical simulation.Proceedings of the 6th International Tooling Conference.Karlstad, 2002. p. 667.

[15] Sehitoglu H. ASM handbook, vol. 19. Materials Park (OH):ASM International; 1996. p. 527.

[16] Lieurade HP, Dias A, Giusti J, Wisniewski A. Experimentalsimulation and theoretical modelling of crack initiation andpropagation due to thermal cycling. High Temp Technol1990;8:137.

[17] Hockanadel PW, Edwards GR, Maguire MC, Baldwin MD. Theeffect of microstructure on the thermal fatigue resistance ofinvestment cast and wrought AISI H-13 hot work die steel.Transactions of the 18th International Die Casting Congress andExposition, Indianapolis, 1995. p. 343.

[18] Baron HG, Bloomfield BS. Resistance to thermal stress fatigueof some steels, heat-resisting alloys and cast irons. J Iron SteelInst 1961;197:223.

[19] Norstrom L-A. Performance of hot-work tool steels. ScandJ Metall 1982;11:33.


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