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Yogesh K. Potdar

Alan T. Zehnder

Department of Theoretical

and Applied Mechanics,

Cornell University,

Ithaca, NY 14853

Measurements and Simulationsof Temperature and DeformationFields in Transient Metal Cutting

Advanced finite element software makes it possible to perform accurate simulations of

orthogonal metal cutting provided all input parameters such as material properties, fric-tion and material separation criteria are known. In principle, such properties can bedetermined by performing a series of cutting experiments and mechanical property tests,and then iterating the finite element simulations until acceptable agreement is reached.Cutting measurements have generally included only cutting forces and tool-chip tempera-tures. We hypothesize that by closely coupling simulations to conventional cutting forcemeasurements and with fine scale spatial and temporal experimental measurements oftemperature and strain fields, questions related to the choice of parameters in finiteelement simulations can be resolved. As a step towards that resolution a method for highresolution experimental measurements of temperature and strain fields is presented here.Temperatures of the workpiece and chip are measured during transient metal cutting overareas of 2727m and time scales of 200 ns by using infrared detectors. Three differentmaterials, 1018CR steel, Al6061-T6 and Ti-6Al-4V are tested. A grid method is used tomeasure deformations in steel with a spatial resolution of 50 m.DOI: 10.1115/1.1596571

Introduction

Metal cutting is one of the most important manufacturing tech-

nologies today. Because of the highly nonlinear nature of metalcutting and the complex coupling between deformation and tem-perature fields, a complete understanding of the mechanics ofmetal cutting is still lacking and is thus the topic of a great deal of

current research.Successful modeling of metal cutting is essential to develop

optimal cutting processes. A very large body of research has beenperformed related to metal cutting mechanics 1,2, starting fromtwo-dimensional analytical models 3,4 up to the present empha-sis on complex finite element simulations. Since the early work of

Usui and Shirakashi 5, considerable progress has been made infinite element FE modeling of metal cutting. Strenkowski andMoon 6, Lin and Lin 7, Xie et al. 8 and Marusich and Ortiz9, among others, have developed custom codes to model metalcutting. Others like Shet and Deng 10,11 and Lei et al. 12 usedgeneral purpose codes to simulate metal cutting. Various aspectsof the physics of cutting have been studied. For example, residualstresses in cutting were studied by Shih et al. 13,14. Obikawaet al. 15,16,17, El Hossainy et al. 18, Warnecke 19 and Ma-malis et al. 20 studied chip formation using FE simulations. Yenet al. 21 and Li et al. 22 used FE simulations to study toolwear.

In performing any FE study of metal cutting, one is faced with

a large number of input parameters. These include the materialmodel, which must include large deformation, high strain rate and

high temperature effects, the tool-chip friction model, which couldbe temperature, sliding speed and pressure dependent and the

separation criterion. In principle, such properties can be deter-mined by performing a series of cutting experiments and mechani-cal property tests, and then iterating parameters in the finite ele-ment simulations until acceptable agreement is reached. However,cutting measurements have generally included only cutting forcesand tool-chip temperatures. It is hypothesized that by closely cou-pling simulation to fine spatial and temporal scale experimental

measurements of temperature and strain fields, along with macro

scale measurements such as cutting forces, questions related to the

choice of parameters needed in finite element simulations of cut-

ting can be resolved.

Since temperature plays a key role in metal cutting, impacting

tool wear, cutting force and chip segmentation, there is a great

deal of research on measuring temperatures in metal cutting. For

example, Boothroyd 23 used infrared film to map out steadystate temperatures in turning. Chao et al. 24 used a PbS photo-conductive cell to measure temperatures on the clearance face of

the cutting tool. Prins 25 used a pyrometer to measure rake facetemperatures. Lezanski and Shaw 26 used the intrinsic thermo-

couple method to determine temperature at the tool-chip contact.Ay et al. 27 embedded fine gauge thermocouples in a cuttingtool to map out the temperature. Stephenson 28, Muller et al.29, MSaoubi et al. 30 and Davies et al. 31,32 used thermalvideo cameras to measure temperatures in the deformation zone

ahead of the cutting tool tip, temperatures at the chip-tool inter-

face, and temperatures of the cutting tool. A survey of the tempo-

ral and spatial resolutions of these measurements show that the

best temporal resolution is 50 s 25. The best spatial resolution,5 m is obtained in the thermal microscopy work of Davies et al.

Note that there have been no experiments to measure temperatures

at both high temporal and spatial resolution.

There have been considerable efforts in characterizing deforma-

tion in metal cutting by placing grids on the undeformed surface.

Many of these studies were performed by conducting cutting tests

under very slow speeds, e.g. Bitans and Brown 33 were able toobtain pictures of deformed grid initially printed on the surface

during cutting of wax at speed of 0.07 mm/s. Kufarev, as Zorev

34 describes in his book, studied the change in shape and theshift of the centers of the round prints of a diamond needle on a

polished surface. The tests were carried out for cutting of copper

at 0.3 mm/s. Palmer and Oxley 35 studied the trajectories andspeeds of movement of chip particles by cinemetography for cut-

ting speeds from 0.21 mm/s to 250 mm/s. They reported the thick-

ness of primary shear zone for SAE 1015 under these speeds.

Komanduri et al. 36 give a qualitative description of localizeddeformation by SEM imaging of segmented chips at cutting

Contributed by the Manufacturing Engineering Division for publication in the

JOURNAL OF MANUFACTURING SCIENCE AND ENGINEERING. Manuscript received

December 2001; Revised January 2003. Associate Editor: A. Shih.

Journal of Manufacturing Science and Engineering NOVEMBER 2003, Vol. 125 645Copyright 2003 by ASME

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speeds up to 3.6 m/s. In the current work we measure deforma-tions while cutting 1018 CR Steel at 4 m/s using a 50 m squaregrid.

Experimental Setup and Procedure

To design an experiment tractable by FEM simulations and inwhich our high resolution infrared optics are securely mounted, atransient, orthogonal cutting experiment was developed. In thissetup a cutting tool swings on the end of a 0.6 m long, 19 kg masspendulum 37. The pendulum supports, workpiece and the infra-red system are securely mounted to an optical table.

The maximum cutting speed, limited by the length of the pen-dulum, is 4.5 m/s. Materials that have been cut successfully in-clude 1018 CR steel, 3.125 mm thick, at depths up to 250 m,6061-T6 aluminum alloy up to 500 m, and Ti-6Al-4V, up to 150m depths of cut. The cutting tool was a commercially availableC6 grade carbide insert with 0 rake angle. The depth of cut is notconstant, however for the first 3 mm of cutting where we measuretemperatures, it does not change more than 4%. In this region ofinterest, the cutting speed drops by 2% during cutting. The cuttingforce is measured via strain gauges bonded to both sides of thecutting tool holder. To synchronize the position of the cutting toolwith the force and temperature signals, a photodiode-IR LED ar-rangement is used.

The temperature of the workpiece and the chip is measuredusing a linear array of 16 liquid nitrogen cooled indium anti-

monide InSb infrared IR detectors. Similar systems have beenused to measure temperature fields at the tips of dynamicallygrowing fractures, 38,39,40. IR radiation is focused onto thedetectors through a 3 reflective mirror system designed and fab-ricated at Cornell. The detectors are 81 m square with 100 mcenter to center spacing. Through the 3 lens, each detector fo-cuses on a 2727 m region on the workpiece. The total field ofview is 27494 m. As shown in Fig. 1, the detector array wasaligned parallel to the velocity of the cutting tool. Each detectorelement has its own preamplifier. Outputs from these are recordedin parallel at up to 10 million samples/sec per element using two4 channel, 12 bit, high speed Nicolet 4094 digitizers. Note thatalthough signals were sampled at intervals of 100 ns, the temporalresolution is limited to 200 ns due to the finite bandwidth of theIR amplifiers.

The detectors are calibrated by recording the IR signal on thedetectors when exposed to a range of known, uniform temperaturefields. For each material tested a separate calibration curve wasobtained. Since the temperatures measured in metal cutting arehigher, especially for Ti-6Al-4V, than the temperature up to whichcalibration can be carried out, a 2nd order polynomial is fitted tothe available data and is used to extrapolate temperatures beyondthe calibrated range. Similar techniques have been used for cali-bration by Guduru et al. 41 and Zehnder et al. 39,42. Note that

total radiant energy increases as T4, where T is the absolute tem-perature. However, due to the wavelength sensitivity of the IRdetectors and taking as target surface temperature in C, the

detector system output signal increases approximately as 2.The workpiece is small sheet, 30 mm20 mm and 3.25 mm

CRS and Al or 1.905 mm Ti-6-4 thick. It is held in place by a

rigid fixture that ensures accurate and repeatable location of thetest specimen with respect to the optical system. All specimens areprepared to a uniform surface finish by fine grain sand blasting,followed by polishing with 600 grit paper. The tool is adjusted inthe pendulum arm such that it can cut the desired depth of cut.This is achieved by using a set-up-specimen which is a depth ofcut shorter in height than the test specimens. The tool is adjustedso that it is just touching the top of set-up-specimen. Then forthe actual cutting experiment the setup specimen is replaced bya test specimen.

In a single experiment, the temperatures in a 27494 m re-gion can be mapped. With the help ofX, Yand Ztranslation stageson which the detectors are mounted, we can focus on any region

of interest. Referring to the coordinate system shown in Fig. 1, bykeeping Y position fixed, we can change Z positions to see tem-perature evolution at various heights along the depth of cut. Thus,in a given set of experiments, we can scan all the region ofinterest.

In some sense, each experiment is different, but we have tested

the repeatability of the experiments extensively and forces, chipmorphology and temperature signals for identical experimentshave been found to be very repeatable. For example, the forces forall tests of a given material and given depth of cut are within 10percent of each other. IR signals obtained in identical tests, look-ing at the same region on the workpiece did not differ more than57 mV and the maximum IR signal obtained while cuttingAl6061-T6 was found to be within 6872 mV. It may be worthnoting that experiments are repeatable only within the precision ofthe fixtures involved; the height can be adjusted only within 50m, the optical system is refocused from one series of tests to theother and the least count of micrometer used for this is 10 m. Inaddition to cutting forces and temperature measurements, the de-formation of the chip is measured with high spatial resolution in aselected set of experiments.

Finite Element Analysis

Since the objective of the research was to explore the couplingof simulations and experiments, the depth of cut, cutting speed,and choice of materials in the simulations was dictated by theexperiments. 1018 CR steel was chosen as the primary materialfor testing and simulation. Cutting of two other materials,Al6061-T6 and Ti-6Al-4V, were simulated once the methodologywas established for 1018 CR steel. It should be noted that themethodology employed here is quite general. Using this approachShet et al. 10,11 simulated AISI 4340 at cutting speeds up to 2.5m/s and a range of rake angles between 15 deg30 deg and pre-dicted residual stresses in AISI 1020 steel. All analyses and mod-eling were performed using ABAQUS Standard, a commercial,general purpose finite element code 43.

A schematic diagram of the FE model is shown in Fig. 2a. Arigid cutting tool moves towards the left with a constant velocity,v, and removes a layer of workpiece material. The thickness oflayer to be removed is decided a priori, and it is assumed thatmaterial separates along the line of cut, i.e. the path traced by thecutting tool tip. The tool is assumed to be sharp. To facilitate thismaterial separation three contact pairs are defined in the FEmodel. Pair 1 is between the chip and rake face of the tool, pair 2is defined between the bottom surface of the chip layer and thesurface of the workpiece along the line of separation, pair 3 isbetween the workpiece and the tool tip, Fig. 2a. Fixed displace-ment boundary conditions are assumed along edges DC, CB, andAB.

Fig. 1 Schematic of orthogonal cutting showing a typical re-gion on which IR detectors are focused in a single experiment.

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A typical finite element mesh used in this work is illustrated inFig. 2b. The model consists of 3470 nodes in the workpiece and80 nodes in the tool. The height of the chip layer is equal to thedesired depth of cut 150 or 250 m. The characteristic length ofelements in the chip layer is 14% of the depth of cut. Plane strainelements are used. The sharp, 0 deg rake angle, C6-carbide inserttool used in the experiments was modeled with zero tool-tip radiusand effectively made rigid by using an artificially high elasticmodulus. Heat conduction into the tool is accounted for. The ini-tial 64 deg inclination of the chip layer elements is used to alle-viate numerical problems due to distortion of the elements as theyseparate from the workpiece and slide along the rake face of thetool. An initial chip separation is adopted in order to achieve a

smooth transition of the cutting process in the beginning whentool first comes in contact with the workpiece. Without this initialchip separation, it was not possible to obtain the convergence.

Constitutive Models

The thermal and mechanical properties of both the workpieceand the tool are critical to the thermomechanically coupled analy-sis. Inelastic deformation of the workpiece materials is modeledwith an overstress power law model, written as 44

pD 0

1 n

, for 0 (1)

p0, for 0 (2)

where p is the effective plastic strain rate,

is the current flowstress, and 0 is the static flow stress. This rate dependent powerlaw is highly suitable for high strain rate applications at finitestrains. The values of D and n, Table 1, are temperature indepen-dent and are obtained for the three materials as best fits to theavailable high strain rate data 4447. The temperature and straindependence in the model comes from the dependence of the cur-rent flow stress on strain and temperature, i.e. 0f(,p) ,where is the temperature, and p is the effective plastic strain.These values are taken from Refs. 48,49. It is assumed that 90%of the plastic work is dissipated thermally 42.

Temperature dependence of initial yield stress under static load-ing and Youngs modulus are illustrated in Figs. 3a, b. Initial

yield stresses are 1018 CRS: 0478 MPa, Al 6061: 0276 MPa, and Ti-6-4: 01050 MPa. Room temperatureYoungs moduli for these materials are 1018 CRS: 210 GPa, Al6061: 70 GPa and Ti-6-4: 104 GPa. Figure 3c shows hardeningunder static loading at room temperature. Flow stress as a functionof strain level and temperature is given as a tabular input for FEprogram. Figure 3d shows the data used along with the best fitcurves. The temperature dependence of the Youngs modulusworkpiece only and of the thermal properties workpiece andtool are taken from standard references 48,49,50,51,52 and areaccounted for in all simulations.

Chip Separation Criterion. In this study two chip separationcriteria are explored. The first is based on a critical stress criterion.In this criterion nodes along the cutting path or between contactpair 2 and 3 in Fig. 2 separate, or debond, when the tensile stressat a specified distance ahead of that pair reaches a critical value.Limitations in the software used allow this critical value to dependon temperature, but not on strain rate or other parameters. With

this in mind, the critical distance was chosen to be on the order of2 or 3 element lengths 70100 m. At this distance, the strainrates are modest and one can justify the use of quasistatic strengthvalues in the separation criterion. It should be noted that if thisdistance is chosen to be too large, the critical stress is neverreached and the analysis does not converge as the elements adja-cent to the tool keep severely deforming and there is no materialseparation along the line of cut. Using very small distances resultsin physically unrealistic, long cracks ahead of the tool. The values

of distances and critical stresses chosen for different materials are:1018 CR steel, critical stress of 648 MPa at 100 m, Al6061,critical stress of 350 MPa at 100 m, Ti-6Al-4V, critical stress of1300 MPa at 20C and 450 MPa at 600C, at 75 m.

The second procedure uses a critical distance criterion in whichseparation occurs at a fixed distance ahead of the cutting tool tip.In this study we used a critical distance of one element length 35m for 1018 CR steel and Al6061-T6 with 250 m depth of cut,and 21 m for Ti-6Al-4V with 150 m depth of cut.

The goal here is not to claim that chip separation occurs by agiven physical mechanism but to discover if any of the availablecriteria describe the process sufficiently well. Cutting simulationsusing the two criteria were conducted with identical meshes andmaterial properties. We observe that forces are consistently higherin the analyses using the critical distance criterion. The maximum

temperatures reached are also higher in the critical distance crite-rion. However the higher temperatures occur in a small region, onthe order of one element size near the rake face. In particular, it isnoted that the critical stress criterion links separation directly tothe material properties. This means that such a criterion is morelikely to capture inhomogeneous material deformation, includingformation of shear bands. Note also that this criterion results inseparation extending some distance ahead of cutting tool. This iscontrary to more commonly believed scenario that material sepa-ration occurs only as the tool reaches a material point and there isno crack of variable length ahead of cutting tool. We suspect thatany separation criterion based on material properties in FE mod-eling will result in a region of separation ahead of the tool and thatthe length of this region will fluctuate during transient cutting,particularly when there is localized deformation.

If a critical distance criterion is used, it is assumed that therewill not be a crack extending ahead of tool tip, however, thevalue of critical distance is an issue. Note that the critical distancecan be made a function of cutting tool tip radius, however, in thiswork, the cutting tool was always assumed to be sharp, i.e. zerotool tip radius and the critical distance was taken to be one ele-ment size. Refinement of the mesh and hence choice of differentdistances could lead to different results.

Simulations performed using the critical distance criterion didnot show the strain localizations which indicate the onset of seg-mented chip formation. However, strain localization is observed inthe simulations using the critical stress criterion. This is illustratedin Fig. 4 which shows localized higher values of equivalent plastic

Fig. 2 a Schematic of FE model showing contact pairs, bfinite element mesh of the workpiece, chip layer and cuttingtool.

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strains forming periodic bands on the chip. Lacking adaptive re-meshing, ABAQUS Standard is not capable of simulating seg-mented chip formation. However, the presence of localized bands

of deformation in Figs. 4a, b indicates the onset of segmentedchip formation for both Al 6061 and Ti-6-4, consistent with theexperiments in which segmented chips are formed for cuttingthese materials at 4.3 m/s.

It is also noted that for Al6061-T6, plastic strain contours ap-pear to emanate from the tool tip, while for Ti-6Al-4V the plasticdeformation extends a small distance ahead of the tool tip on theorder of 50 m. A similar trend is observed for temperature fieldsin Ti-6Al-4V, i.e. they also seem to extend some distance ahead ofthe cutting tool in the simulations using critical stress criterion. Itwas observed during the experiments that temperature fields inTi-6Al-4V extend significantly ahead of the cutting tool. The cor-relation between plastic strain and temperature fields is expected,especially in the region where plastic strains are the only source ofheat generation.

Thermal and Frictional Contact Between Chip and RakeFace. Friction along the tool-chip interface plays a critical rolein the metal cutting process since it directly affects cutting force,

power and amount of energy dissipated at the chip-tool interface.In this study, it is modeled using the modified Coulomb frictionlaw for the contact pair 1 in Fig. 2a. Let be the chip shear

stress at a contact point along the tool-chip interface and p be thenormal pressure at the same point. This law states that slip occurswhen c and sliding occurs when c . The critical frictionstress is given by

cminp ,th (3)

where is the friction coefficient and th is a threshold valuerelated to material failure. It should be noted that the conventionalCoulomb friction law is recovered if th is set to infinity. For thisstudy, th was chosen to be the shear yield stress of the workpiecematerials at the anticipated tool-chip interface temperature. Thesewere taken to be: c388 MPa for 1018 CRS, 170 MPa for Al6061, and 310 MPa for Ti-6Al-4V. The friction coefficients

0.1 and 0.3 are explored. As discussed later, 0.3 appearsto provide the best comparison between experimental and FE re-

sults, considering temperature, contact length and cutting force.The thermal contact on the tool-chip interface is another impor-tant parameter in the simulation. If there is perfect thermal con-tact, the temperatures across the chip-rake face interface are con-tinuous. This seems to be a reasonable assumption in steady statecutting, however it is not clear if this assumption is valid fortransient cutting. If thermal contact is poor, then even for lowcoefficients of friction, temperatures occur close to the rake face.As the coefficient of friction is increased, the maximum tempera-ture increases. If thermal contact is high, then the maximum tem-peratures reached are reduced significantly. In the case of lowercoefficients of friction, the location of occurrence of maximumtemperature shifts substantially away from the rake face and into

Fig. 3 a Variation of initial yield stress with temperature. b variation of elastic moduluswith temperature. c strain hardening under quasistatic loading. d dependence of flowstress on strain rate, data points for Al6061 45 and Ti-6Al-4V 46 are from experiments andthose for steel are from Shawki-Clifton model 44. Lines show best fit curves obtainedusing parameters in Table 1.

Table 1 Best fit values of D and n

Material D, 1/s n

Al 6061 1.63105 1.75

1018 CR steel 5.13109 10.95

Ti-6Al-4V 2.23104 1.53


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the chip. As the coefficient of friction is increased, the location ofmaximum temperature moves closer to the rake face.

This is illustrated in Fig. 5, which shows how the maximumtemperature reached in FE simulation for three different materialsis affected by gap conductance, k, across chip-rake face interface,where the gap conductance is defined as

qkAB, (4)

and q is heat flux, W/m2, k is the gap conductance, W/m2K andA and B are temperatures on the two sides of the gap. Figure 6

shows how the temperature field changes when k varies from 104

to 107 W/m2K. It is noted that in the case of higher k, maximumtemperatures are lower and occur in the chip at a distance of theorder of 2550 m away from the rake face rather than at the

chip-rake face interface. This is due to the large heat flux out ofthe chip into the cold cutting tool when thermal contact is high.In the context of Coulomb friction, thermal contact conditions

and the coefficient of friction need not be related directly. More-over, it is not clear if Coulomb friction is the right assumption forthe conditions prevailing on the rake face. For example Moufkiet al. 53 suggest that the coefficient of friction may depend ontemperature. It may also be a function of contact pressure. Onecan conjecture that thermal contact is a function of contact pres-sure and would improve with increased contact pressure. One canalso conjecture that if coefficient of friction were to be a functionof contact pressure, it would increase with increased contact pres-sure. In this scenario, higher contact pressure would result in in-

Fig. 5 Variation of maximum temperature with gap conduc-tance for 1018 CR steel, Al6061-T6, and Ti-6Al-4V. FE simula-tions using critical distance criterion.

Fig. 6 Change in FE using critical distance and 0.3 tem-perature field C with gap conductance k while cutting 1018

CR steel a low conductance, k104 Wm2"K b high conduc-

tance, k107 Wm2"K.

Fig. 4 FE simulated cutting at 4.3 ms, 0.3. The contourplot shows equivalent plastic strains. Localized regions ofplastic strain form bands on the chip surface, indicating onsetof segmented chip formation a Al6061-T6, d250 m b Ti-6Al-4V, d150 m.


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creased temperatures due to greater friction but would result inreduced temperatures due to better thermal contact and greaterheat conduction.

Experimental Results and Comparison to FE Simula-

tions

Temperature Measurement. Figure 1 shows schematicallythe field of view of the IR detectors and the coordinate frame usedin the following discussion. Experimental data was obtained for1018 CR steel, Al6061-T6 and Ti-6Al-4V. Figure 7 shows the

temperatures from different elements of the IR array system for asingle experiment cutting 1018 CR steel at 4.3 m/s, 250 m depthof cut. In this case, the line of view of the IR array is 250 mabove the line of cut; i.e. the view is just along the free surface ofthe workpiece. Detector element number 16 sees the cutting firstand detector number 3 sees the cutting last, e.g. referring to Fig. 1,detector 16 is focused on point A and detector 3 is focused on B.Since this is a transient cutting experiment, we note that the tem-perature increases as cutting progresses. The maximum tempera-ture at any point, such as A in Fig. 1, is recorded when cutting toolis very near to A. The signal on the detector drops off as thecutting tool passes through its field of view. Figure 7 shows thetemperature until the maximum temperature is reached. Figure8a shows the comparison of results from the FE study withexperimental result for temperature evolution at a point 1.55 mmfrom the beginning of cutting and 250 m above the line of cut.We note that the FE model tracks the temperature rise reasonablywell. The cutting experiments are repeated by moving the detec-tors along the Z axis. Figure 9a shows maximum temperaturesobtained during a set of nine experiments and is compared withFE results for 0.1 and 0.3. In this case, the peak tempera-ture rise 250 m above the tool tip is about 330C.

Figure 10 shows the IR signals obtained on the same detectorelement in five of these repeat experiments in which detectors arefocused at 0, 150, 250, 300, and 400 m for 1018 CRS and Al6061, and 0, 50, 100, 150, 200 m for Ti-6-4, above the line ofcut. Let us call these locations P, Q, R, S and T respectively. TheIR signals before calibration are shown to give an illustration ofwhat the raw data look like. Note the excellent agreement for thetime when the maximum signal is recorded. This correspondsfairly closely with the time when the tools rake face has just

reached the region being seen by the detector. Thus, we note thatthe highest temperatures occur close to the rake face as expected.Note that there is a spread of about 5-6 s in the time duration

when the signals shown reach maximum value. This is attributedto experimental scatter, since these signals are taken from fivedifferent experiments. The tool travels 2025 m in 56 s. Theposition detection system shows an uncertainty of about 25 m inthe repeated experiments. Given these limitations, along with theunderstanding that IR systems resolution is 2727 m, it is as-sumed in this study that all the maximum values of the signal arereached at the same time instant when the tools rake face has justreached the point of interest.

Figure 10 illustrates how the rise time for temperature at agiven point depends on its height. This is consistent with the shearplane model of cutting, where the shear plane extends away fromthe tool tip towards the free surface. Thus we note that the tem-

Fig. 7 Experimental data while cutting 1018 CRS. Detector ar-ray is focused along the original free surface of the workpiece250 m above the line of cut.

Fig. 8 Temperatures, experimental and FE comparison of evo-lution of temperature at a point 1.55 mm from start of cut. a1018 CR steel b Al6061-T6 c Ti-6Al-4V.


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perature at point P along the line of cut rises last since this pointsees the deformation and temperature rise just as tool tip reachesvery close to it. However, at points Q, R, S and T above the lineof cut the temperature rises sooner. For points S and T whichwere not in the original workpiece material, the beginning of tem-perature rise is an indication of when cut chip reaches that point.The thickness of the cut chip can be calculated as the time dura-tion for the rise at such point multiplied by the cutting tool veloc-ity. Given our uncertainty about tool position, cutting velocity andrelative position of maximum temperature with position of rakeface, this is an approximate calculation. For 1018 CRS, points S

and T show a rise time about 100 s, indicating 100 s

4.3 m/s430 m of chip thickness. This is close to measuredchip thickness of 400 m.

Converting the data from the above set of experiments to tem-perature using the detector calibrations, the temperature distribu-tion ahead of the cutting tool can be visualized. Figure 11bshows the temperature distribution in 1018 CRS ahead of the cut-ting tool after 350 s of cutting. It is compared with the FEcontour plot obtained from analysis based on the critical stresscriterion with 0.3 in Fig. 11a. The FE temperature fieldagrees well 100 m and above the line of cut. The FE temperaturefield close to tool tip just above the line of cut decreases rapidly inmagnitude whereas experimental contours show a more gradualchange. Although the FE analysis predicts higher maximum tem-

Fig. 9 Max Temperatures: Experimental and FE using criticaldistance criterion vs. height a 1018 CR steel b Al6061-T6 cTi-6Al-4V.

Fig. 10 IR signals at different heights for a 1018 CR steel bAl6061-T6 c Ti-6Al-4V.


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peratures, they occur within 50 m of the rake face. In general,FE contours are more closely packed than the experimentallymeasured contours. Further experimentation, possibly with a twodimensional array of detectors, 54 combined with advances inFE simulations would help resolve these discrepancies.

With similar sets of experiments, temperature fields are ob-tained for Al6061-T6 and Ti-6Al-4V. The temperature rise forAl6061 is less than that for 1018CR steel, the maximum tempera-tures being of the order of 220C. The signals at different heightsare illustrated in Fig. 10b. Figure 12 shows the temperature dis-

tribution ahead of the cutting tool after 350 s of cutting. It is

compared with an FE contour plot obtained from analysis based

on critical distance criterion with 0.3 shown in Fig. 12b. TheFE temperature field agrees well at 100 m and above the line of

cut. Similar to the observations for steel, the FE temperature field

close to the cutting tool tip just above the line of cut decreases

rapidly in magnitude whereas experimental contours show a more

gradual change. Also, FE predicts higher maximum temperatures

within 30 m of the rake face.

Temperature rise in Ti-6Al-4V is much higher than Al6061-T6

and 1018 CR steel. Maximum temperatures up to about 800C are

observed. Figure 10c shows IR signals obtained on a single de-tector as the tests were repeated by changing the position of the

line of detectors along Z axis. Figure 13 shows the temperature

distribution ahead of the cutting tool after 150 s of cutting. It is

compared with FE contour plots obtained from analysis based on

critical distance criterion with 0.3 in Fig. 13b and FE basedon critical stress and 0.1 and 0.3 in Figs. 13c, d respec-tively. It is noted, as stated in the earlier section, that the critical

distance criterion shows slightly higher maximum temperatures.

Also, Fig. 13c, d demonstrate how an increased coefficient offriction increases maximum temperatures, especially near the rake

face.

In all materials the FE temperature field shows higher tempera-

tures near the rake face and the contours are more closely packed.

In other words, experiments predict more diffused temperature

field than FE analysis. FE predicts higher maximum temperatures,but again they occur only within 50 m of the tools rake face.

Both of these differences between the FE and experimental results

reflect the finite spatial resolution of the temperature measure-

ments. Note that temperature is measured on the surface; the tool-

chip interface may be hotter. Note also that the temperatures could

be considerably higher under steady state conditions where the

initially cool cutting tool cannot act as a heat sink.

Contact Length. After a series of tests, the cutting tool insert

shows a burnished region where the chip slides over the rake

surface. Table 2 shows a comparison of contact lengths from ex-

periments and FE analyses for different materials. The table, along

with comparison of the FE temperature fields to the experiments

suggests that 0.3 is a reasonable choice for the friction

coefficient.

Fig. 11 Comparison of experimental and FE using critical

stress criterion and 0.3 temperature field ahead of cuttingtool while cutting 1018 CR steel at 4.3 ms, 250 m depth of cut,350 s after the beginning of cut.

Fig. 12 Comparison of experimental and FE using critical distance criterion and 0.3 tem-perature field ahead of cutting tool while cutting Al6061-T6 at 4.3 ms, 250 m depth of cut, 350s after the beginning of cut.


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Cutting Force. Dividing the cutting force by the depth of cutand thickness of the sample, the average measured cutting forceswere 1018 CRS: 1120 MPa, Al 6061: 640 MPa and Ti-6-4: 1670MPa. The cutting force from the FE simulations did not stronglydepend on the coefficient of friction. The cutting force was higherwhen the critical distance criterion was used. With the criticalstress criterion and using 0.3, the FE cutting force was 1018CRS: 1200 MPa, Al 6061: 660 MPa and Ti-6-4: 2070 MPa. TheFE and experimental cutting force results agree well for the CRSand Al. However, for Ti-6-4, the FE cutting force is considerablyhigher than the experiments. This could indicate that the strainrate hardening used for Ti-6-4 was too large or that the thermalsoftening was insufficient or that the separation criterion used toolarge a stress value.

Deformation Measurement. A fine gauge grid method wasdeveloped to measure deformation in orthogonal metal cutting.The goal is to develop a simple, inexpensive method which can beused to measure deformation fields with high spatial resolutionunder a wide range of cutting conditions and material choices.

A 50 m grid pattern was deposited onto the polished polishedsurface of a 1018 CR steel workpiece using a TEM grid as a

mask. The sample was cut at a speed of 4.0 m/s, and a depth of500 m, deeper than the depth used for the temperature measure-ments so that a large number of grid points would be availablerelative to the depth of cut. The cut chip is collected and thedeformed grid pattern on this chip is then imaged using scanningelectron microscopy, Fig. 14a.

The image is processed by outlining the edges of the deformedgrid and calculating the centroids. A best fit ellipse is found foreach deformed contour, as shown in Fig. 14b along with thedirections of the major and minor axes of the best fit ellipses. Ifthe original circle of radius r deforms into an ellipse with majorand minor axes a and b respectively, then, a/r and b/r are theprincipal stretches. Principal stretches obtained from three suchrows of grids are plotted in Fig. 14c. It is noted that principalstretches decrease as one moves away from the rake face. Highvalues of principal stretches in row B indicate the presence of ashear band. In Fig. 14c, principal stretches are plotted and com-pared with FE results (0.1, critical distance criterion. Sincethe FE results show very similar values of deformations in theregion detailed in Fig. 14b, there is only one line showing FE

results in the Fig. 14c. Simulations with 0.3 show very highvalues of maximum stretch, up to 7, very near the rake face, but atabout 150 m from the rake face, the values fall to the almostsame level as those for 0.1. Experimental measurements showless deformation that the simulations. This is expected, given thatany experimental method averages deformation over a finiteregion.

Summary and Conclusions

Temperature fields in transient orthogonal metal cutting havebeen measured experimentally with a spatial resolution of 2727 m and temporal resolution of 200 ns. This is accomplished

Fig. 13 Comparison of experimental and FE temperature field ahead of cutting tool while cuttingTi-6Al-4V at 4.3 ms, 150 m depth of cut, 150 s after the beginning of cut a experimental b FEcritical distance, 0.3 c FE critical stress, 0.1, d FE critical stress, 0.3

Table 2 Comparison of contact length on the rake face, in m,for different materials and friction coefficients. FE resultsbased on critical stress separation criterion.

Material Experimental FE, 0.1 FE, 0.3

1018 CR Steel 500 345 475Al6061-T6 585 475 530Ti-6Al-4V 440 300 450


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by building a system consisting of a linear array of 16 high speedInSb IR detectors and a 3 optical system. Three materials, 1018

CR steel, Al6061-T6 and Ti-6Al-4V have been cut at 4.3 m/s. Themaximum measured temperatures in these cases were 330C,220C and 800C for the conditions tested. The maximum tem-peratures occur very near the rake face of the tool and some dis-tance, on the order of a depth of cut up from the tool tip. Thetemperature field ahead of the cutting tool after a certain durationof cutting, 350 s for 1018 CR steel and Al6061-T6, and 150 sfor Ti-6Al-4V, is mapped by carrying out a series of experiments.The IR system, due to its finite resolution, may not be able todetect temperatures occurring on length scales smaller than 25m. The performance of the magnifying lens and the design of IRdetectors will be a major engineering challenge in development ofany higher resolution systems. A grid method has been used to

measure deformations occurring during cutting of 1018 CR steelat 4 m/s. Principal stretches of up to 3 are measured. Finite ele-ment simulations were performed using a general purpose code. Itis shown that friction, chip separation criterion and the extent ofthermal contact can play a significant role in defining temperatureand deformation fields. The novel experimental methods pre-sented here provide an opportunity to gather local data, that whenused in conjunction with simulations could lead to realistic com-putational cutting models.

Acknowledgments

The authors acknowledge the support of the National ScienceFoundation CMS 9700698. This work made use of the CornellCenter for Material Research Shared Experimental Facilities, sup-ported through the National Science Foundation Materials Re-search Science and Engineering Centers Program DMR-0079992. Particular acknowledgement is made of the use of theResearch Computing Facility and of the Materials Facility. Thefirst author wishes to thank Ms. Nidhi Sawhney for useful discus-sions on image processing in Matlab.

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