Temperature Increase in Forming of

Advanced High-Strength Steels Effect

of Ram Speed Using a Servodrive Press

Ali FallahiarezoodarCenter for Precision Forming,

The Ohio State University,

339 Baker Systems,

9171 Neil Avenue,

Columbus, OH 43210

e-mail: fallahiarezoodar.1@osu.edu

Ruzgar PekerCenter for Precision Forming,

The Ohio State University,

339 Baker Systems,

9171 Neil Avenue,

Columbus, OH 43210

e-mail: peker.4@osu.edu

Taylan AltanCenter for Precision Forming,

The Ohio State University,

339 Baker Systems,

9171 Neil Avenue,

Columbus, OH 43210

e-mail: altan.1@osu.edu

In forming of advanced high-strength steel (AHSS), the tempera-ture increase at die/sheet interface affects the performance oflubricants and die wear. This study demonstrates that finite-element (FE) analysis, using commercially available software,can be used to estimate temperature increase in single as well asin multiple stroke operations. To obtain a reliable numerical pro-cess design, the knowledge of the thermal and mechanical proper-ties of the sheet as well as the tools is essential. Using U-channeldrawing the thermomechanical FE model has been validated bycomparing predictions with experimental results. The effect ofram speed and stroking rate (stroke per minute (SPM)) upon tem-perature increase in real productionlike operation have beeninvestigated. Deep drawing of CP800 and DP590 sheets in a ser-vodrive press, using an industrial scale die, has been studied.Thinning distribution and temperatures in the drawn part havebeen investigated in single and multiple forming operations. It isfound that temperatures may reach several 100 deg and affect thecoefficient of friction (COF). The values of COF under produc-tionlike conditions were compared to that obtained from labora-tory experiments. This study illustrates that in forming AHSS, (a)the temperature increase at the die/sheet interface is relativelyhigh and should be considered in process design stage, and (b)the lubricant performance is significantly affected by the ramspeed and sheet/die interface temperature during deformation.[DOI: 10.1115/1.4033996]

1 Introduction

Light-weight design has become essential for automotive indus-try due to its considerable economic and ecological benefits (i.e.,higher fuel economy and lower emission). High-strength andlightweight materials are introduced for a wide range of body

panels and structural parts. In addition to Al alloys, advanced andultrahigh-strength steels are increasingly used in automotive man-ufacturing. Therefore, to minimize the part failure and tool wear,press speed during part deformation often needs to be reduced.Electromechanical servodrive presses provide precision ram posi-tion and velocity control during the stroke. This allows improve-ment of drawability by only reducing the ram velocity during thepart deformation stage, without significantly reducing the produc-tion rate [1,2].

In sheet forming of AHSS, the heat generated due to frictionand plastic deformation can reach up to about 100–200 �Cdepending on process type and speed [3,4]. As reported by Farrenand Taylor [5], approximately 90% of the work done for the plas-tic deformation is converted into heat and the remaining 10% isused up in increasing the internal energy of the material [5]. Formaterial with temperature dependent flow stress, temperature raiseduring the deformation can affect the stiffness of the material andlowering the flow stress data. This softening can increase the realarea of contact resulting in higher adhesion. It has been reportedthat adhesion starts to rise at approximately 125–150 �C [6].

From the tribological point of view, temperature rise during thedeformation can affect the properties of the lubricant and influ-ence the lubrication performance. Organic based lubricants, whichoften used in sheet metal stamping, can be used below 250 �Cwithout significant thermal degradation [7]. Therefore, it is impor-tant to accurately determine the fraction of deformation workwhich goes to increasing the temperature of the material duringthe sheet forming process. There are a few studies that investi-gated the temperature generation in forming of AHSS and alumi-num alloys using numerical or experimental procedures [3,4,8,9].Kim et al. [3], investigated the shear fracture at die corner radiusin draw bending test of DP980 and they concluded that thedeformation-induced heating, in the order of 75 �C, has a domi-nant effect on the occurrence of shear failure [8].

In order to determine the temperature, cost, and time could begreatly reduced by using computer simulations rather than runningexperiments with a try-out press in the die shop. FE simulationswere used for temperature prediction in metal forming process[3,4,8,9].

In published literature, the temperature at the part/tool interfaceis predicted after a single forming operation. However, it is impor-tant to estimate how the temperature increases after several num-bers of forming operations, especially in forming of AHSSmaterials. In the current study, first a simulation of U-channeldrawing is conducted to validate the methodology and assump-tions used in the thermomechanical simulation model, by compar-ing the predictions with experimental results available in theliterature. Then, a similar thermomechanical FE model for deepdrawing of industrial scale nonsymmetric panel is developed. Thetemperature rise in sheet and tools is determined for two differentAHSS materials (1.4 mm CP800 and 1.4 mm DP590). The simula-tion results are validated with experimental results in terms ofthinning distribution. Effect of COF and lubrication performanceon temperature generation at tool/sheet interface is investigated.Simulations of multiforming operation of U-channel forming anddeep drawing process were conducted to investigate the effect ofthe ram speed (SPM) on temperature rise at the tool surface afterseveral consecutive forming operations.

2 Material Model and Element Type

In the present study, three different AHSS materials, complexphase grade steel CP800 (1.4 mm), and dual phase grade steelsDP590 (1.4 mm), and DP780 (2 mm), are examined. Tensile prop-erties for DP780 were obtained from literature [4], and for CP800and DP590 tensile tests, in rolling direction, were conducted atHonda R&D Department of Honda America Raymond, OH. Hill48 plasticity law with isotropic material model was used for simu-lating the plastic behavior of the sheet. The flow stress for CP800and DP590 materials were obtained from the viscous pressure

Manuscript received January 15, 2016; final manuscript received June 21, 2016;published online July 19, 2016. Assoc. Editor: Matteo Strano.

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bulge (VPB) test and are shown in Fig. 1. The detail informationabout VPB test can be found in previous publications [1,2,10].The flow stress data obtained from the VPB test then were extrap-olated up to true strain values equal to 1 using the Hollomon’sequation (r¼ k en). The calculated K and n values, along with theother properties obtained from the tensile test, are summarized inTable 1.

Thermal properties of the sheet and the tools are selected basedon the data available in the literature [4,6,8]. Due to unavailabilityof the thermal data for the materials used in this study, the thermalproperties for sheet material are based on the properties of low-carbon steel. Using the actual thermal properties for each specificmaterial will increase the accuracy of the predictions. However, inthe current study, the predicted temperature for U-channel draw-ing is compared with available experimental results and it isobserved that the predicted temperature for given boundary condi-tions and process parameters is in a good agreement with experi-mental measurements. Therefore, the prediction error for usingthe thermal properties of low-carbon steel in forming AHSS mate-rials can be considered to be negligible. For the tools, the relevantthermal properties for the tool steel were utilized using the dataavailable in the PAM-STAMP database and Ref. [4]. The thermal dataused for the sheet and the tools are summarized in Table 2. Thedissipation factor, a, describes the percentage of internal workwhich will be converted to temperature. b shows the fraction offriction work dissipation converted into heat. In simulations, it isassumed that the heat generated due to the friction is equally dis-tributed into the tools and sheet surface.

It is known that the surface heat transfer coefficient (SHTC) isa function of gap and pressure between the sheet and the toolsurfaces. However, due to simplicity, this parameter was consid-ered as a constant value. In simulation of multiforming operations,the heat convection between the tool surface and the air, a factorthat influences temperature reduction during the die opening andtransformation stages, is also considered. The initial temperatureof the sheet and the tools were considered 20 �C in allsimulations.

The sheet material was simulated using shell elements with fiveintegration points through thickness direction. The initial elementsize was selected based on the minimum sliding radius in each dieset. In simulations of deep drawing process “mesh refinement”option was activated as suggested by PAM-STAMP V15.1 manual.

This option allows the software to more accurately calculate theflow of the material by automatically reducing the element size atlocations where severe deformation occurs. The elastic deflectionof the tools is neglected. The tools are simulated as surfaces (notvolume) to reduce the simulation time. Six thermal points wereconsidered in the thickness direction for the tools. The predictionerror due to use of surface element instead of solid element is notinvestigated in this study.

3 U-Channel Drawing

3.1 Simulation Setup and Validation of ThermomechanicalFE Model. Three-dimensional coupled thermomechanical FEmodel is developed to analyze the U-channel forming of DP780(2 mm). Figure 2 shows the geometrical parameters of the die andspecimen. The commercial software PAM-STAMP-V15.1 withexplicit solver was used. The die geometry and tooling setup issimilar to what was used in experiments to be able to validate theaccuracy of the simulation model and results by comparing withexperimental results [4]. The model consists of die, punch, blank-holder, ejector pin, and sheet. The punch was fixed in all threedirections. Blankholder and die were fixed in X and Y directionsand allowed to move in the Z direction which was considered asthe forming direction. 27 kN constant force was applied on theblankholder and the ejector pin. The distance between the die andthe blankholder was checked in several forming strokes to ensure

Fig. 1 Flow stress data for 1.4 mm CP800 and DP590 obtainedfrom the VPB test

Table 1 Mechanical properties of sheet materials used in this study. Data for DP780 is from Ref. [4].

Tensile test Bulge test

Material Yield stress (MPa) Tensile stress (MPa) Young’s modulus Strength coefficient (K) (MPa) Strain hardening exponent (n)

DP780 (2 mm) 587 884 210 (assumed) 1332 0.11DP590 (1.4 mm) 355 600 180 1311 0.27CP800 (1.4 mm) 770 900 230 1325 0.11

Table 2 Thermal properties used in the thermomechanical FEsimulation. Thermal properties for sheet are for low-carbonsteel and are from previous publication by Pereira and Rolfe [4].

Tools Sheet

Density (kg/mm3) 7.7� 10�6 7.9� 10�6

Heat conductivity (J/s m �C) 22 52Heat convection with air (W/m2 �C) 50 —Specific heat capacity (J/kg �C) 460 480SHTC (J/s m2 �C) 20 20Expansion coefficient (1/ �C) 12� 10�6 12� 10�6

Friction energy heat dissipation factor (aÞ 0.9 —Friction heat distribution factor (b) 0.5 —In elastic heat fraction — 0.9Initial temperature (�C) 20 20

Fig. 2 Schematic cross section of the simulation setup andthe geometrical parameters used for simulation of U-channelforming [4]

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the closure of the blankholder. The die movement is specified byspeed versus time curve. The same speed pattern used in theexperiment was applied to the die representing the motion of thesingle-action mechanical press with 1 SPM ram speed [4]. Basedon the geometry of the press, the crank rotational speed and thedrawing depth used in the experiment, the deformation starts atabout 8 mm/s and reduces to 0 mm/s at the end of 40 mm formingstroke. The total forming duration was about 9 s. The temperatureat the die/sheet interface was predicted and compared with experi-mental results.

3.2 Results of U-Channel Drawing and Validation ofThermomechanical FE Model. Several assumptions were madein the simulations due to unavailability of the required data. Theseassumptions, discussed below, can limit the accuracy of the simu-lation results: (a) the accurate thermal properties of the sheetmaterial used in this study were not available. So, the propertiesof low-carbon steel were used from literature; (b) the heat transfercoefficient which determines the amount of the heat transferredfrom sheet to the dies was considered as a constant value. How-ever, as determined experimentally, this value is a function of thegap and pressure between the sheet and the die [9,11]; (c) thesheet and the tools are simulated as surface elements and thermalintegration points through the thickness were considered to ana-lyze the heat conduction inside the tools; (d) a constant value ofthe COF is considered in all simulations and it is known that theCOF is not constant in all locations in the die during stamping.

Figure 3 shows the predicted maximum temperature generatedat the die/sheet interface compared to the experimental results.Simulation results are in good agreement with experimental meas-urements from Ref. [4]. The maximum predicted temperature atthe die/sheet interface goes up to about 45 �C (25 �C temperaturerise) after about 3 s of drawing (about 22 mm drawing depth) anddrops to about 33 �C at the end of deformation. Similar trend wasobserved in experiments. The temperature rise in terms of valueand trend follows the experimental measurement very well.Therefore, the prediction error as a result of using thermal proper-ties of low-carbon steel instead of the real values for the specificsheet material used in the experiment can be considered to be neg-ligible. Also, it can be concluded that the methodology and theboundary conditions used to develop the thermomechanical FEmodels in this study can predict the temperatures close to thatencountered in the real stamping operations.

Some inconsistency of the simulation results compared to theexperiment can be observed at about 2–3 s after the deformationstarts. Since the temperature and pressure can affect the lubrica-tion condition and COF, assuming a constant value of COF insimulation causes some errors in predictions. However, due tounavailability of data for variable COF in function of pressure and

temperature for the lubricant used in the study, it was not possibleto use a variable COF in the simulations.

4 Deep Drawing of Nonsymmetrical Industrial Size

Panel

4.1 Experimental Procedure. Figure 4 shows a schematic ofthe die geometry used for deep drawing process in this study. Thedie was manufactured by SHILOH Industries, Valley City, OH.1.4 mm CP800 and DP590 sheet materials were used in theexperiments. The tensile test and the bulge test results showedDP590 is more formable than CP800, by comparing the uniformelongation and bulge height. Therefore, the total drawing depths70 mm and 48 mm were selected for DP590 and CP800, respec-tively. Rectangular sheets (600� 395 mm for CP800, and640� 435 mm for DP590) were used. In case of DP590, the cor-ners of the sheet were chamfered by 100� 75 mm to reduce thepossibility of wrinkle and improve the drawing process.

The sheets samples were cleaned by acetone and then lubri-cated in both sides. Cup drawing test (CDT) as a regular lubrica-tion test in sheet metal stamping [3,12] was used for theevaluation of lubricants. Ten different lubricants suggested bylubricant companies, for deep drawing of AHSS, were tested. Incup drawing lubrication test, sheets are lubricated on both sidesand then formed up to a predefined drawing depth using a constantblankholder force. The tests were performed at about 30 mm/sram speed using 160-ton hydraulic press. Then, the flange perime-ter of the cups was measured and the lubricants were ranked basedon the length of the flange perimeter. The flange length, as a func-tion of lubrication performance, is smaller when the cup draw-inis more and it is an indication of better lubrication performance.The selected lubricant used in this study, polymer synthetic, wasprovided by IRMCO Company, Evanston, IL, and was one of thebest ranked lubricants from the CDTs.

A 300-ton Aida servodrive press with 25-ton servohydrauliccushion was used. This press provides accurate control of ram dis-placement/speed and blankholder force during the deformationprocess. The press speed, at the stroke position when the punchtouched the sheet, was about 75 mm/s. A constant blankholder forcewas used. Preliminary simulation results showed that 200 kN and250 kN constant blankholder forces were sufficient to maintain clo-sure of the blankholder and avoid creation of wrinkles for the sheetsize used to form DP590 and CP800, respectively. Three samplesfrom each material were used in the tests. However, due to the diffi-culty of cutting the materials for thickness measurement, only twosamples from each material were measured.

4.2 FE Model of Deep Drawing Process and Predictionof Thinning. Three-dimensional thermomechanical simulationmodel of deep drawing process for CP800 and DP590 with

Fig. 3 Max. temperature predicted by FE simulation and meas-ured experimentally at die/sheet interface during the U-channeldrawing of 2 mm DP780. The experimental results are fromRef. [4].

Fig. 4 Schematic of the nonsymmetric industrial scale dieused in this study for deep drawing process (die set built byShiloh Industries)

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1.4 mm thickness was developed using the die geometry shown inFig. 4. The model consists of die, punch, blankholder, pad, andsheet. The boundary conditions are set similar to what was used inthe experiments. The punch was fixed in all three directions.Blankholder, pad, and die were fixed in X and Y directions andallowed to move in the axial Z direction (forming direction). Aconstant 250 and 200 kN blankholder force were applied on theopposite direction of the die movement for CP800 and DP590,respectively. A 2 kN force was also applied on the pad to avoidthe separation of the sheet from the top of the punch. The toolswere modeled using surface elements and considered as rigidbody. Similar thermal and mechanical properties as described inSec. 2 for the tools and sheets are used.

As mentioned in Sec. 4.1, due to the different formability ofDP590 and CP800, different sheet sizes and drawing depths wereused for each of these materials in the experiments. Therefore, asimulation model, with drawing depth and blankholder forceaccording to the experimental setup, is developed for eachmaterial.

In order to be able to investigate the effect of the material prop-erties on temperature rise at the tool/sheet interface, simulationwith similar forming conditions and sheet dimensions used forCP800 was also conducted for DP590. In this way, the only differ-ence that provides different temperature predictions in the simula-tion models is the sheer material properties.

In order to determine the COF from the CDT test, several simu-lations of cup drawing process with different COF values wereconducted with similar boundary conditions used in the lubrica-tion evaluation tests. The COF value which provides the similarflange length as measured from the experiment is selected as theCOF for that lubricant, COF¼ 0.08. Then, this calculated COFvalue is used as initial COF in simulations of deep drawing pro-cess. The predicted load–stroke curves from the simulations withCOF¼ 0.08 were compared with that obtained from the experi-ments. It was observed that the predicted load–stroke curves wereslightly lower than experimental results. Therefore, the initialCOF (0.08) was modified and increased to the value (COF¼ 0.12)that the predicted load–stroke curves match with experiments.

4.3 Results of Deep Drawing Process Using NonsymmetricalDie Geometry and Comparison With Experimental Results.Similar assumptions, mentioned in Sec. 3.2, are also consideredfor simulations of deep drawing process. However, as discussed inSec. 3.2, these assumptions are not significantly affecting theaccuracy of simulation results in terms of temperature prediction.Comparison of simulation results with experimental measure-ments in terms of both temperature prediction and thinning distri-bution shows that the FE model and the boundary conditions usedin the current study are reasonably reliable.

4.3.1 Prediction of Thinning Distribution and Validation ofthe FE Model. Thinning distribution along the curvilinear lengthat the corner of the panel where the maximum thinning is pre-dicted in simulation, Fig. 5, is compared with experimental meas-urements, Figs. 6 and 7 for CP800 and DP590, respectively. ForCP800, the maximum predicted thinning on the part is about 13%,Fig. 5. However, the maximum thinning value calculated alongthe curvilinear line is about 9%, Fig. 6. The reason is that the cur-vilinear line where the thinning values are measured in experi-mental samples is not passing through the element which showsthe maximum thinning in simulation. For CP800, results showedthe predicted thinning is in the same range of the thinning mea-surement from the experiments except at location number 4. Thenominal sheet thickness was about 1.4 mm and the real initialthickness was not measured before the test. A small differencebetween the nominal and real thickness (even about 0.05 mm) fora sheet in range of 1.4 mm thickness cause about 3% different inthinning measurement. The differences between the predictionand experimental thinning measurements in this study are lessthan 3%.

Similar results are illustrated in Fig. 7 for DP590 formed up to70 mm drawing depth using 200 kN blankholder force. For thismaterial also, the maximum difference between the predicted thin-ning value and the experimental measurement is about 4%.

4.3.2 Effect of the COF on Temperature Rise at Die/SheetInterface. Once the accuracy of the simulation results are vali-dated by comparing the thinning predictions with experimentalmeasurements, temperature rise at the tool/sheet interface is inves-tigated for CP800 and DP590 used in this study. The COF used inthe deep drawing simulation is initially determined by CDT andthen modified by comparing the load–stroke curves of the simula-tions and experiments, as described in Secs. 4.1 and 4.2. The rea-son for calculating different COF in deep drawing process

Fig. 5 Measurement of the thinning distribution along a curvi-linear length of 1.4 mm CP800 panel formed up to 48 mm. Thick-ness of the six locations along the cutting line are measured inexperimental sample and the thinning values are comparedwith simulation results for approximately the same locations (a)simulation prediction and (b) the formed panel and the loca-tions of measurements.

Fig. 6 Thinning percentage versus curvilinear length at thecorner of the formed panel, shown in Fig. 5, material CP800, ini-tial thickness 1.4 mm, Drawing depth 48 mm, and blankholderforce 250 kN

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compared to the CDT can be described due to different pressuresand temperatures at die/sheet interface in these two differentforming processes. The CDT was performed at about 30 mm/sram speed but the deep drawing process was conducted at about75 mm/s. In the deep drawing process, the forming speed is abouttwo times faster and the die corner radius is smaller (10 mm com-pared to 16 mm) compared to the CDT. Smaller die corner radiusand more complexity of the die cavity compared to the cup draw-ing cause higher strains at the part and consequently increase thedeformation-induced heating. Also, the faster ram speed in deepdrawing experiments reduces the time of heat distribute throughthe tools by conduction. So, the temperature at the tool/sheet inter-face was higher in deep drawing process compared to CDT. Thishigher temperature can affect the lubrication performance andincrease the COF. Also, the different pressure between the toolsand the sheet in the CDT and the deep drawing operations can beanother reason for different COF values. As reported in severalprevious publications [7,13,14], assuming a constant COF forwhole model is not accurate. FE model with variable COF as afunction of temperature, and pressure can increase the accuracy ofthe simulation results. However, variable COF values as functionsof pressure and temperature are not available at this moment.

Temperature rise at the die/sheet interface is predicted for twodifferent COF values (0.08 and 0.12) during the forming of 1.4CP800 up to 48 mm. Temperature rise versus stroke is shown inFig. 8. As expected, the maximum temperature for lower COF islower. Changing the COF from 0.08 to 0.12 increases the maxi-mum temperature at die/sheet interface at 48 mm stroke fromabout 80 �C to 91 �C. Therefore, using accurate COF is necessaryfor accurate prediction of temperature in the simulation of sheetmetal stamping.

4.3.3 Effect of the Material Properties on Temperature Gen-eration. It is known that the deformation-induced heating is pro-portional to the deformation energy, i.e., the area below the flowstress curve of the material. The flow stress data developed byVPB test for the materials used in this study are shown in Fig. 1.

The nominal thickness for both materials (CP800 and DP590)is 1.4 mm. Therefore, for a given forming conditions (COF, sheetsize, blankholder force, drawing depth, and forming speed), thetemperature generated at the part for CP800 must be higher. Thepredicted maximum temperature is approximately 33% higher forCP800 than for DP590. Comparison of the tensile strengths ofboth materials shows that the ultimate tensile stress for CP800 isabout 30% higher than DP590. These results show the effect ofmaterial strength on temperature generation during the deep draw-ing process. Also, for higher strength material, the forming loadand contact pressure in forming is higher and it is known that theCOF is a function of pressure. Therefore, the COF value is differ-ent during the forming of different materials with differentstrengths and consequently, the temperature generation due tofriction work is different. However, since a constant COF is con-sidered in the simulations the effect of the material properties onfriction work is not considered in this study.

The FE simulation of deep drawing process for DP590 with theconditions used in the experiment (640� 435 mm rectangularsheet, 200 kN blankholder force, and 70 mm drawing depth) isalso conducted, to predict the temperature in the part and tool/sheet interface at similar conditions to what the material wastested. The location of the maximum temperature at 70 mm strokewas in the same location shown in Fig. 9, but the value reached upto about 130 �C. The maximum temperature calculated at the die/sheet interface at this stroke was about 110 �C.

Currently, metal manufacturers and automotive companies areinvestigating the development and production of materials withtensile strength of about 1500 MPa. Based on the simulationresults presented in this study, the temperature at the part for theseultrahigh-strength materials can reach to up to 200 �C. This hightemperature can affect the lubrication performance, the materialproperties, and the surface quality of the part and the tooling.

5 Multistroke Forming Operations

5.1 Simulation Setup. Similar methodology used for the sin-gle forming operation was used to simulate the multiforming oper-ation of the U-channel-forming and the deep drawing processes.The temperature rise at the die/sheet interface was predicted

Fig. 7 Thinning percentage versus curvilinear length at thecorner of the formed panel, Fig. 5, material DP590, initial thick-ness 1.4 mm, drawing depth 70 mm, and blankholder force200 kN

Fig. 8 Max. temperature rise at die/sheet interface for two val-ues of COF, during the deep drawing of 1.4 CP800 with 250 kNblankholder force and 75 mm/s ram speed. (Die geometry isshown in Fig. 4).

Fig. 9 Temperature predicted at panels formed up to 48 mmdepth with 250 kN blankholder force and 75 mm/s ram speed;(top) 1.4 mm CP800, (bottom) 1.4 mm DP590

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immediately after when the deformation stage was finished andalso before starting the next steps.

For U-channel drawing, two different forming speeds, 5 SPMand 30 SPM were considered. The ram (punch) speed during de-formation was calculated using the well-known equations forslider-crank mechanism of the press [14,15].

The information, used in the simulation of multistroke formingoperations of U-channel drawing was as follows:

� For press speed at 5 SPM, the ram speed at contact was41 mm/s, the duration of forming was 2 s, and the transfertime between two strokes was 10 s.

� For press speed at 30 SPM, the corresponding values were250 mm/s (ram speed at contact), 0.32 s (duration of form-ing), and 2 s (transfer time between two consecutivestrokes).

The material data for 1.4 mm CP800 is only used in simulationof multistroke deep drawing operations. Crank diameter of theservopress used in the experiment, 400 mm, was utilized to createthe ram speed versus stroke curve. As a particular characteristic ofthe servodrive presses, the selected ram speed versus stroke,Fig. 10, allows the press to move at 6 SPM before and after touch-ing the material while during the material deformation the ramspeed remains constant, 75 mm/s. The total duration from the startof one forming operation to the next was considered about 10 s (6SPM). The forming duration was about 0.64 s (75 mm/s and48 mm drawing depth) and 9.36 s was considered for die openingand part transformation.

5.2 Results and Discussion. Servodrive presses, compared tomechanical presses of similar capacity, can be run 50–100%faster, in terms of SPM. However, it is not completely clear howtemperature at the tool sheet interface increases at multiformingsteps at different forming speeds. Therefore, in the current studythe simulation of multiforming operations of U-channel drawingand deep drawing were investigated as two examples.

Figure 11 shows the increase of maximum temperature pre-dicted at die/sheet interface for 2 mm DP780 after several opera-tions of U-channel drawing with two different forming speed (5and 30 SPM). At 5 SPM ram speed, the maximum temperature atdie/sheet interface increased from 41 �C to 46 �C after eight opera-tions. At 30 SPM, maximum temperature increased from 100 �Cto 117 �C after nine operations. However, the temperature rise atdie/sheet interface does not increase significantly (less than 1 �C)after nine operations and reach almost steady-state condition. Asexpected, using faster forming speed and transfer time, the tem-perature at the die/sheet increased significantly. Faster formingoperation reduces the time of heat transfer from the tools surface

into the air, and therefore, the temperature rise after several opera-tions was more significant when using a faster ram speed.

Figure 12 shows the max temperature rise at die/sheet interfacefor several forming operations of deep drawing process for CP800at 6 SPM production rate. The maximum temperature increased6 �C after seven forming operations. The temperature rise fromthe sixth to the seventh operation is about 2 �C and the tempera-ture rise does not reach steady-state condition after seven opera-tions. However, the main objective of this study is to show howtemperature rises during the deformation of AHSS. The calcula-tion of the maximum temperature when steady-state conditionreached is out of the scope of this study. Therefore, to save com-putation time, the model was developed to predict the temperatureonly up to seven strokes.

Fig. 10 Ram speed versus stroke curve used in 48 mm deepdrawing of CP800 with Aida 300-ton servodrive press and 25-ton servocushion, (crank diameter 5 400 mm), stroke 5 0 is thetop dead center and stroke 5 400 is when the ram is at bottomdead center

Fig. 11 Calculated temperature rise at die/sheet interface in U-channel drawing with: (a) 5 SPM and (b) 30 SPM forming speed,F: forming stage; T: transfer stage

Fig. 12 Temperature rise at die/sheet interface at deep drawingof 1.4 mm CP800 in consecutive multiple forming, F: formingstage; T: transfer stage

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Below is a summary of the results obtained from the simulationof both U-channel forming and deep drawing:

(a) the die/sheet interface temperature increases with increas-ing deformation speed, as indicated by SPM

(b) the die surface temperature continues to increase slightlyduring multiple successive forming operations, until a nearsteady-state surface temperature is reached

6 Summary and Conclusions

The objective of the study, discussed in this paper, was tounderstand and explain how die/sheet interface temperatures canbe predicted for given forming conditions. The knowledge of theinterface temperatures are expected to help in evaluating lubricantperformance and, possibly, in understanding die wear and gallingissues.

Two forming examples were studied: (1) U-channel drawing of2 mm thick DP780 and (2) deep drawing of DP590 (1.4 mm) andCP800 (1.4 mm) in a nearly rectangular die. FE models weredeveloped for both applications to predict die/sheet interface tem-peratures. In case of U-channel drawing, the predictions werecompared with experimental data from the literature and reasona-ble agreement between predictions and measurements were found.In case of deep drawing, the thinning in the part and temperatureswere estimated. The effect of press speed and temperatures uponfriction was discussed. Forming in single as well as in multiple-consecutive press strokes was considered.

This study illustrated that it is possible to estimate temperaturesat the die/sheet interface, using FE simulation. The temperatureprediction and thinning distribution were in good agreement withexperimental results. In U-channel drawing, the temperatureincreases with increasing forming stroke and then decreases to-ward the bottom dead center. The location where the temperatureat the tool/sheet interface starts to decrease is dependent of the ge-ometry and the forming speed. On the other hand, in deep drawingoperations, the temperature in the part as well as the die/sheetinterface keeps increasing with forming stroke. The COF used inthe FE simulation can significantly affect the prediction of thetemperature. Therefore, for accurate temperature prediction, it isnecessary to use a COF value that is closed to the COF valuewhich is present in the experiment.

The information presented in the current study helps to predictthe temperature for a specific forming process for given processconditions. Using this information, the effect of temperature uponlubricant performance in stamping can be better understood sothat for a given application, lubricants may be selected, based onquantitative information. Furthermore, prediction of interface

temperatures will also lead to understanding and possible predic-tion of tool wear in stamping.

Acknowledgment

Tensile data for some of the materials used in this study(CP800 and DP590) were provided by Honda R&D (Mr. JimDykeman). The deep drawing tests were conducted at Aida-America (Mr. Shrini Patil) using the die set, manufactured atShiloh Industries (Mr. Cliff Hoschouer). Financial support wasprovided by the Center for Precision Forming of The Ohio StateUniversity. These supports are gratefully acknowledged by theauthors.

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Journal of Manufacturing Science and Engineering SEPTEMBER 2016, Vol. 138 / 094503-7

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