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die steel
Journal of Materials Processing Technology 92±93 (1999) 293±301
Temperature measurement when high speed machining hardened mould/die steel R.C. Dewesa,*, E. Nga, K.S. Chuaa, P.G. Newtona, D.K. Aspinwalla,b a
School of Manufacturing and Mechanical Engineering, University of Birmingham, Birmingham, UK IRC in Materials for High Performance Applications, University of Birmingham, Birmingham, UK
b
Abstract When turning and face milling hardened (>30 HRC) mould/die steels, it is necessary to use conventional ceramic or polycrystalline cubic boron nitride (PCBN) tool materials. The use of high speed steel and cemented tungsten carbide (WC) tooling for this application is precluded due to their relatively low hot hardness values. In recent years, however, the mould and die industry has begun to utilise solid WC ball nose end mills for the high speed machining (HSM) of hardened steel cavities. Economic tool lives have been reported at high rotational speeds/feed rates. The paper details the use of 6 mm diameter solid, TiCN coated WC ball nose end mills for the HSM of a hot work tool steel, AISI H13, hardened to 52 HRC. Tests were performed on a Matsuura 20 000 rpm high speed machining centre and temperatures were measured using thermocouple and infrared techniques. Recorded tool /workpiece interface and chip temperatures were relatively low (200±4008C) and increased with higher cutting speed, when using worn rather than new tools and with the workpiece inclined at 608 to simulate ®nishing operations on the side of die cavities. # 1999 Elsevier Science S.A. All rights reserved. Keywords: High speed machining; Temperature measurement; Tool steel; Tungsten carbide
1. Literature survey
1.2. HSM cutting temperatures
1.1. High speed machining
Salomon [1] proposed that there was a peak cutting temperature at an intermediate cutting speed and that when cutting speed was increased from this point, there was a reduction in temperature, see Fig. 1(a). Since this claim, most of the literature has concluded that there is no corresponding reduction in temperature at higher cutting speeds. McGee [7] suggested that temperature increased with cutting speed up to a maximum which was equal to the melting point of the workpiece, see Fig. 1(b). No temperature reduction occurred at higher cutting speeds. This explains why there is no ®xed limit to the cutting speed when machining aluminium alloys (other than that imposed by machine tool considerations). The melting point of these alloys (up to 6608C) is lower than the temperature at which cemented carbide and ceramic tool materials begin to lose their strength and wear rapidly. Conversely, Trent [8] stated that ``it is in the cutting of iron, steel and other high melting point metals and alloys that the heat generated becomes a controlling factor''. Here the limit on cutting speed is a function of the cutting tools used.
High speed machining (HSM) was ®rst reported in 1931 by Salomon [1]. One de®nition is that the process ``involves machining at considerably higher cutting speeds and feed rates than those used in conventional machining'', however, it is most commonly used to describe end milling at high rotational speeds. The process has been adapted to a wide range of applications. In the aerospace sector, HSM is used to remove large volumes of aluminium quickly and to produce thin walled sections in wings [2±4]. One of the more recent applications of HSM is in the manufacture of moulds and dies from hardened tool steels [5,6]. Cavities can be produced from solid in the hardened state using HSM, rather than via the more traditional route: machining in the soft condition followed by electrical discharge machining (EDM), grinding and/or hand ®nishing. *Corresponding author. Tel.: +44-121-414-4175; fax: +44-121-4143541 E-mail address: [email protected] (R.C. Dewes)
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Fig. 1. Graphs of temperature vs. cutting speed: (a) Salomon's curve; (b) McGee's curve.
1.3. Temperature measurement techniques Several experimental methods involving radiation, thermocouple and metallography/microhardness techniques have been used to measure cutting temperatures. A standard thermocouple assembly embedded into a cutting tool or workpiece is known as an implanted, inserted or remote thermocouple [9,10]. On the other hand, if the tool and workpiece are used as the dissimilar metals of the thermocouple this is termed a tool/chip or tool/work thermocouple. The former is a low cost method which can be used to establish the distribution of temperatures at different points in a cutting tool by utilising a series of predetermined locations. Unfortunately, the presence of holes in the tool may alter the temperature distribution and it is not possible to directly measure the temperature at the tool/workpiece interface. The tool/chip thermocouple [11] gives the average temperature, not the maximum, at the interface. Drawbacks of the technique are that both workpiece and tool material must be electrically conductive, cutting ¯uid cannot be used, calibration is laborious, secondary voltages may occur and many tool/workpiece combinations do not form ideal thermocouples [8].
Another technique is to implant a suitable insulated thermocouple wire, Te¯on-coated constantan (55% copper, 45% nickel) wire for example, into a workpiece. When the workpiece is sheared during the machining process, the insulation is broken and an instantaneous hot junction is formed between the wire and the workpiece material in the cutting zone, see Fig. 2. An electromotive force (emf) is generated and assuming calibration of the system has been performed, an interface temperature can be calculated. The advantages of this method include easy calibration and use, and the fact that tool materials which are electrical insulators can be used. Unfortunately, the maximum temperature at the tool/chip interface is not always recorded and experimental error is caused by the variation in detection position along the cutting edge. These problems, however, can largely be overcome by test replications. This technique was used in the present work. A less invasive technique involves measurement of the thermal radiation emitted during the cutting process by using infrared (IR) sensitive photographic ®lm [13] or a pyrometer/infrared thermometer [14]. Fig. 3 shows the system used in the present work. Unfortunately, no cutting ¯uid can be used and the temperature measured is usually that of the top face of the chip and not the interface temperature
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Fig. 2. Machining sequence when using the implanted workpiece thermocouple [12].
Fig. 3. Agema Thermovision 900 system with IR camera (right) and system controller (left).
Fig. 4. Sandvik 1020, 6 mm diameter TiCN coated carbide cutter.
because the tool/workpiece interface is obscured. Errors can also result as a consequence of using an incorrect emisivity value. When using high speed steel or WC tools with an iron/ silicon binder, phase transformations in the tool material can be used to deduce the temperature to which the tool has been subjected and produce isothermal maps [15]. Optical microscopy is utilised to compare sections of the cutting tool with standard specimens. The disadvantages of this method include the limited types of cutting tool materials that can be used and the manufacture of standard specimens which must be very carefully prepared. The aim of the present work was to determine cutting temperatures when high speed ball nose end milling hardened die steel, in order to explain the successful use of tungsten carbide (WC) cutters for this operation. Temperatures were measured using the workpiece/Te¯on-coated constantan thermocouple and infrared techniques previously outlined.
coated with titanium carbonitride, TiCN). Consequently, 6 mm diameter 1020 tools, shown in Fig. 4, were used for the present work. Cutters had four ¯utes with a positive helix angle of 308. A tool overhang from the collet of 30 mm was employed. The workpiece material used was AISI H13 hot work tool steel at a hardness of 52 HRC. It was obtained with a chemical composition of 0.38% C, 1.00% Si, 0.34% Mn, 5.00% Cr, 1.3% Mo and Fe balance. Not surprisingly, AISI H13 is not a standard thermocouple material categorised under British Standard BS1041, therefore, a calibration bar was designed, see Fig. 5. For the machining trials, a workpiece block 80mm55mm35 mm was electrical discharge wire machined into two separate pieces as shown in Fig. 6. The design allowed insertion of thermocouple wires in the gap while giving suf®ciently high clamping force to hold the assembly in place during machining. Two diameters (0.075 and 0.5 mm) of Te¯on-coated constantan wire were obtained from Omega Engineering,
2. Experimental 2.1. Tooling, workpiece materials and machine tools Previous HSM machinability work on hardened steels [16], highlighted the suitability of Sandvik Coromant 1020 (formerly MC45) solid carbide ball nose end mills (PVD
Fig. 5. Calibration bar.
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Fig. 6. Workpiece block.
UK. There are a number of reasons why it is desirable to use thinner wires in the measuring junction during machining. They are poorer conductors of heat which reduces the tendency of heat loss to the environment. Secondly, it is critical to cut the wire cleanly to induce a suf®ciently high thermal emf for measurement. The wire diameter also determines the gap width in between the two parts of the workpiece. A large wire diameter would result in a larger gap causing an interruption in cutting and possibly reduced temperature measurements. Prior to the machining tests, the workpiece and thermocouple wires were prepared. Two 600 mm lengths of 0.5 mm diameter constantan wire were utilised for longer connections between the measuring apparatus and workpiece. One of the wires was silver soldered to the side of the workpiece and the other was soldered to a 40 mm length of 0.075 mm diameter wire. The thinner wire was formed into a sinusoidal wave shape, 10 mm high with 2 mm protruding out of the face of workpiece. Fig. 7 shows the workpiece with the two halves separated in order that the looped wire can be seen. The wire was bonded to the male half of the workpiece using adhesive tape. A Cincinnati milling machine (maximum spindle speed 1500 rpm) and a Matsuura FX-5 CNC controlled high speed machining centre (maximum spindle speed 20 000 rpm) were used for the work, see Fig. 8. The latter machine
Fig. 7. AISI H13 workpiece with constantan wire in place.
Fig. 8. Matsuura FX-5 machining centre.
had a maximum power of 15 kW and a continuously variable feed rate up to 15 m/min. 2.2. Equipment ± thermocouple technique A Fluke 8050A digital multimeter, set at its lowest DC voltage range of 0±200 mV, was used for voltage measurement during initial calibration work. The advantage of this was that it displayed voltage values instantaneously, thus saving time. Unfortunately, the refresh rate on the LCD screen of 0.25 s was a limitation. A Gould 100 Ms/s oscilloscope (DSO) 420 was used for some calibration and the majority of temperature evaluation work. Its advantages included high accuracy, high sampling rate, rapid screen refresh rate and a capability to produce hard copies using its pen plotter. When using the HSM set-up at the operating parameters detailed, the thermal emf generated lasted for a very short period of time, typically 40 ms. By using the oscilloscope, the input signal could be monitored continuously, so that the probability of not capturing a short duration thermal emf signal was minimised. A Kane-May KM1420 temperature data logger was used for the calibration work. This was speci®c to the K-type thermocouple standard (nickel±chromium and nickel±aluminium), as in BS1041. The logger had a measuring range from ÿ100 to 12008C with an accuracy of 0.28C and a response time of <1 s. A stainless steel sheathed K-type probe was connected to the logger and was used to measure the temperature at the measuring junction of the calibration bar. The data in the logger was transferred to a personal computer using the equipment manufacturer's software TERM2UK. The computer enabled the storage of data ®les in ASCII format and allowed the generation of temperature versus time plots on the screen. In order to perform thermocouple calibration, it was necessary for the measuring or hot junction of the calibration bar to be heated while the other end, the reference or cold junction remained at room temperature. Because of this, the furnace calibration method reported by Braiden [17] was not suitable. The induction heating method [18] was also
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machine was not equipped with in®nitely variable table feeds. By looping the thermocouple wire in the manner shown in Fig. 7, each test was effectively replicated a number of times. Tests on the Matsuura machining centre involved both the thermocouple and infrared techniques, see Table 2. All tests were performed using the thermocouple technique, however, due to limited equipment availability, the infrared technique was limited to tests 5 ± 10. Resources did not allow a full factorial experiment to be performed, therefore, it was felt that most information could be gained in the time available by investigating four variables: cutting speed (tests 1 ± 3), tool wear (tests 3 ± 10), angle of the workpiece (tests 3 ± 6) and axial/radial depth of cut. In order to simulate the actual machining of a mould or die cavity, the workpiece was clamped both horizontally and at 608 to the horizontal while maintaining a constant cutting speed by varying the rotational speed. Fig. 9 illustrates the interface cutting speed at workpiece angles 08 and 608 when using a 0.5 mm depth of cut. Table 2 classi®es roughing, semi-®nishing and ®nishing cuts depending on the depth of cut applied. During all tests, dry cutting was employed and feed was ®xed at 0.1 mm/ tooth.
rejected. It was decided that the most suitable heating method was to use an oxygen±acetylene torch. 2.3. Equipment ± infrared technique Equipment consisted of an IR camera and a system controller, see Fig. 3, borrowed from the UK's Engineering and Physical Sciences Research Council (EPSRC). The controller was a dedicated microcomputer set up speci®cally for use with the camera. It consisted of a processor unit, VGA video monitor, mouse and keyboard with specialised system controls for image handling. The software was menu driven and was contained in a windows environment. The camera was designed to operate at wavelengths in the infrared spectrum (8±12 m). The temperature ranges selected for this work were 0±2508C and 100±6008C. The emissivity level used was 0.82. Interchangeable lenses were available for the use with the scanner. In this case, a 408x208 resolution lens was used which had a minimum focal length of 0.3 m. 2.4. Machining parameters In the tests on the Cincinnati milling machine, only the thermocouple technique was used. Fixed parameters included the angle of workpiece (08), down/climb milling, dry machining, use of a tool with 0.3 mm maximum ¯ank wear and an axial/radial depth of cut of 0.5 mm. The effect of feed per tooth and cutting speed were evaluated, see Table 1. Feed per tooth values are approximate because the
2.5. Experimental procedure ± thermocouple technique The calibration experiment involved a number of steps. The 0.075 mm diameter thermocouple wire was silver soldered into hole A on the calibration bar, see Fig. 5. The bar
Table 1 Test parameters on the Cincinnati milling machine Test
Cutting speed (m/min)
Feed per tooth (mm/tooth)
Spindle speed (rpm)
Feed rate (mm/min)
a b c d
16 21 28 21
0.10 0.10 0.10 0.05
870 1140 1500 1140
420 420 570 220 Fig. 9. Effect of workpiece angle on cutting speed.
Table 2 Test parameters on the Matsuura machining centre
Roughing cut
Roughing cut Semi-finishing cut Finishing cut
Test
Maximum flank wear (mm)
Angle of workpiece (degree)
Cutting speed (m/min)
Spindle speed (rpm)
Axial depth of cut (mm)
Radial depth of cut (mm)
Feed rate (mm/min)
1 2 3 4 5 6 7 8 9 10
0 0 0 0.3 0 0.3 0 0.3 0 0.1
0 0 0 0 60 60 60 60 60 60
100 150 200 200 200 200 200 200 200 200
9597 14 396 19 194 19 194 10 610 10 741 10 741 10 741 10 741 10 741
0.5 0.5 0.5 0.5 0.5 0.5 0.2 0.2 0.2 0.2
0.5 0.5 0.5 0.5 0.5 0.5 0.2 0.2 0.1 0.1
3839 5759 7678 7678 4244 4244 4296 4296 4296 4296
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was clamped in place and the thermocouple wires from the measuring and reference junctions were connected to the multimeter. At the measuring junction, the K-type stainless steel probe was inserted into hole B, see Fig. 5, and connected to the data logger. The measuring junction was heated with the oxygen±acetylene torch to 7008C. The voltage on the multimeter was recorded and the bar allowed to cool. During cooling, at ®xed temperature intervals, the emf value was recorded. The values were plotted on a graph of voltage versus temperature. A linear relationship between voltage and temperature was established and the equation of the line determined. The calibration procedure was repeated ®ve times and ®ve calibration lines superimposed. The equation of a line of best ®t was determined. One calibration test using the oscilloscope rather than multimeter was carried out in order to con®rm that both sets of equipment gave similar results. When calibration was complete, initial tests were undertaken on the Cincinnati milling machine. The workpiece/ thermocouple junction was set-up as discussed previously. The assembly was clamped in a vice with plywood used to insulate it from the machine to prevent electrical interference. Machining cuts were taken across the workpiece using the parameters given in Table 2 until the protruding thermocouple wire was cut and a signal triggered in the oscilloscope. The voltage was printed using the pen plotter. When the entire face of the workpiece had been machined and a number of temperature readings obtained, the workpiece was re-prepared. The voltages were analysed and converted to temperatures. A mean temperature value was calculated and taken as the interface temperature for the test. All subsequent work was performed on the Matsuura machining centre. Each time a new workpiece was clamped in the vice, it was levelled to 30 m. Trigonometric calculations were used to check the level of the workpiece at 608. The voltage and timebase settings on the oscilloscope were modi®ed depending on the cutting parameters, in order to produce reliable results. A CNC program was executed which resulted in the cutter machining across the workpiece, retracting above the workpiece and returning to the start of the next cut and applying the radial depth of cut. Each time a wire was machined through, a signal occurred and the voltage trace was recorded and printed out as with the tests on the Cincinnati machine. 2.6. Experimental procedure ± infrared technique The Agema Thermovision 900 system was set up, the temperature range selected and the level and span of the image adjusted to produce a recognisable image. The IR camera was ®lled with liquid nitrogen, positioned at a safe distance from the workpiece to avoid any possibility of damage by the swarf and connected to the system controller. The automatic image storing sequence (1 s intervals) was initiated, see Fig. 10. The machining sequence was set up as described previously with the cutting
Fig. 10. A thermal image captured when machining workpiece at 608.
parameters given in Table 2. After the sequence was recorded, the images were recalled and analysed by using the ``spotmeters'' in the software. When placed over the image, these gave a temperature value at a particular point. The maximum chip temperature was therefore recorded for each test. 3. Results and discussion 3.1. Calibration of the thermocouple The ®ve sets of calibration results using the multimeter were combined to give a line of best ®t, see Fig. 11 for graph and equation of the line. Also shown are the results and equation obtained when using the oscilloscope. The difference in readings between the two measuring devices was only 0.14% at 20 mV. Therefore, the equation from the oscilloscope was used to calculate the temperature from the induced thermal emf for the machining tests. 3.2. Temperature results The effect of cutting speed on interface temperature, as measured using the thermocouple technique can be seen in Fig. 12 which shows the results of Cincinnati tests a ± c and Matsuura tests 1 ± 3. Mean temperature values are shown. An analysis indicated that mean and median temperatures fell within 108C in all cases. It can be seen that over the range of cutting speeds tested, the interface temperature increased in a manner similar to that found by McGee when machining aluminium [7]. This is in line with conventional metal cutting theory [19]. A machine with higher spindle speeds would be required to fully test Salomon's ®ndings that temperature reduces at higher cutting speeds [1]. Although not shown graphically, reducing the feed rate from 0.10 to 0.05 mm/tooth at a cutting speed of 21 m/min
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Fig. 11. Calibration results and equations for multimeter and oscilloscope.
gave a very slight reduction in temperature from 1108C to 1078C. The next stage of the work was to test the effect of workpiece angle and tool wear on interface temperature, see Fig. 13. There was an increase in temperature when machining with the workpiece at 608 rather than 08. This can
be explained because at 08 the cutting speeds on the ball nose end mill ranged from zero in the centre to 200 m/min at the periphery, while at 608 the range is from 173 to 200 m/min, see Fig. 9. As a result, higher average cutting speeds occurred in the cutting zone and therefore, higher temperatures would be expected. An effect which is likely to have prevented temperatures at 08 from being even lower was the nature of the ball nose cutter. At 608, all four teeth of the cutter would have done an equal amount of cutting, however, at 08, two of the teeth would have done very little work due to the fact that they stop short of the centre of the tool. The effective feed per tooth would have approached 0.2 mm/ tooth which would be expected to contribute to higher temperatures. When using a tool with 0.3 mm of ¯ank wear, temperatures increased. With a worn tool, more surface contact between the tool and workpiece occurred and the increase in sliding friction and rubbing caused higher temperatures. Figs. 14 and 15 show the effect of axial and radial depth of cut on interface and chip temperatures, respectively. Roughing, semi-®nishing and ®nishing cuts are indicated as R, SF
Fig. 13. The effect of workpiece angle and tool wear on interface temperature.
Fig. 14. Effect of axial and radial depth of cut on interface temperature.
Fig. 12. Effect of cutting speed on interface temperature.
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4. Conclusions
Fig. 15. Effect of axial and radial depth of cut on chip temperature.
and F. In all but one case, chip temperatures are much lower than interface temperatures. Values measured by the infrared technique were typically one third those from the thermocouple. It is widely accepted that during metal cutting, the majority of heat (80%) is transferred to the chip, therefore, it could be argued that this is a larger difference than would be expected. A possible explanation is that with the IR camera and lens used, the hot zone was relatively small compared to the total area scanned, see Fig. 10. This, combined with limitations in the software which would allow the use of ``spotmeters'' but not output a maximum temperature over an area, meant that there was no guarantee that the hottest point was measured. It is considered that the thermocouple technique was intrinsically superior for this application and gave more accurate, absolute temperature values. The differences in temperature between the three types of cut are as would be expected, with smaller depths of cut giving signi®cantly lower temperatures. This would validate the use of small depths of cut during the HSM of hardened steels in order to minimise cutter wear. With the roughing cut, there is a large difference between the temperatures generated with the new and worn tools, a smaller difference when semi-®nishing and an even smaller range when ®nishing. This can be explained by reference to the extent of ¯ank wear on the worn tools, see Table 2, higher levels of ¯ank wear giving a bigger difference in temperature. Currently, EDM is widely used for the production of moulds and dies from hardened steels. Due to the high temperatures which occur, typically 10 000 ± 20 0008C, surface integrity problems such as heat affected zones, white layers and tensile residual stresses are found on machined components. With the temperatures measured using the thermocouple technique (<4008C), it would be expected that surface integrity problems would be minimised when using HSM, as the tempering temperature of AISI H13 (5508C) is not exceeded. In parallel work, some of which has been published [20], the authors found that HSM gave superior surface integrity to EDM.
1. Interface temperatures measured using the thermocouple technique when machining with the workpiece at 08 were 198±3018C. Higher temperatures, 247±3858C, occurred at a 608 workpiece angle. 2. Temperature increased with cutting speed and no reduction at higher speeds occurred. This is contrary to Salomon's theory and agrees with McGee's findings. Feed per tooth did not have a large effect on temperature. 3. Machining with a worn tool generated higher temperatures than when using new tools. 4. The IR technique indicated lower temperatures than the thermocouple method with values of 68±3908C. Heat losses to the tool and workpiece and drawbacks with the IR technique explain the lower temperatures measured. 5. The relatively low cutting temperatures measured explain why WC products can be successfully used for the HSM of hardened steels. Acknowledgements We would like to thank Prof. A.A. Ball, Head of the School of Manufacturing and Mechanical Engineering, and Prof. M.H. Loretto, Director of the IRC in Materials for High Performance Applications, for provision of facilities and funding. We are grateful to Messrs. R. Fasham, W. Hewitt, P. Thornton and J. Wedderburn for their assistance with the experimental work. Thanks also go to the Control, Design and Production Directorate of EPSRC, De Beers Industrial Diamond Division (Pty) Ltd., Dynacast (UK) Ltd., Kieninger Tooling Ltd., Matsuura Machinery plc, Sandvik Coromant (UK) Ltd., United Engineering Forgings Ltd., and WH Smith and Sons (Tools) Ltd., for additional funding and support. We also wish to express our gratitude to the EPSRC for lending the infrared equipment.
References [1] C. Salomon, Verfahren zur bearbeitung von metallen oder bei einer bearbeitung durch schneidende werkzeuge sich ahnlich verhaltenden werkstoffen, German Patent no. 523594, (1931). [2] G.L. Criger, High speed machining in production, SAMPE Quarterly, 12 April 1981. [3] J. Tlusty, High speed milling, in: Proceedings of the 6th IMEC, Osaka, Japan, (1994), pp. 35±42. [4] J.R. Coleman, No myth high speed machining, Manuf. Eng. 61 (Oct. 1992). [5] J. Gough, High speed machining for toolmaking applications, Prec. Toolmaker 8(3) (1990) 154±157. [6] H. Schulz, High speed machining - some of the latest developments, Metalworking World 3 (1994) 32±36. [7] F.J. McGee, High speed machining - study: methods for aluminium workpieces, Am. Machinist (1979) 121±126. [8] E.M. Trent, Metal cutting, 3rd edition, Butterworth, London, ISBN 0750610689, (1991).
R.C. Dewes et al. / Journal of Materials Processing Technology 92±93 (1999) 293±301 [9] M.P. Groover, G.E. Kane, A continuing study in the determination of temperatures in metal cutting using remote thermocouples, J. Eng. Ind. (Trans. ASME) 93(2) (1971) 603±608. [10] A.H. Redford, B. Mills, S. Akhtar, Temperature - tool life relationships for resulphurised low carbon free machining steels, CIRP Annals 25(1) (1976) 89±91. [11] P. Lezanski, M.C. Shaw, Tool face temperatures in high speed milling, J. Eng. Ind. (Trans. ASME) 112(2) (1990) 132±135. [12] W. Chen, A report of the research on surface integrity of hardened steel following face milling, School of Manuf. and Mech. Eng., University of Birmingham, UK, (1994). [13] G. Boothroyd, Temperatures in orthogonal metal cutting, Proc. Instn. Mech. Engrs. 177(29) (1963) 789±802. [14] J.P. Kottenstette, Measuring tool-chip interface temperatures, J. Eng. Ind. (Trans. ASME) 108(2) (1986) 101±104. [15] P.A. Dearnley, E.M. Trent, Wear mechanisms of coated carbide tools, Metals Technology, 9 (1982) 60±75.
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[16] R.C. Dewes, D.K. Aspinwall, High speed machining of hardened tool steel using cemented carbide tooling, in: Proceedings of the Third International Conference on Progress of Cutting and Grinding, Osaka, Japan, vol. 3, (1996) pp. 217±222. [17] P.M. Braiden, The calibration of tool/work thermocouples, Proc. in: Proceedings of the Eighth International MTDR Conference, Manchester, UK, Perganon, Oxford, (1967), pp. 653±666. [18] G. Barrow, A review of experimental and theoretical techniques for assessing cutting temperature, CIRP Annals 22(2) (1973) 203. [19] B.T. Chao, K.J. Trigger, An analytical evaluation of metal cutting temperatures, Trans. ASME, 73, (1951) p. 57. [20] R.C. Dewes, D.K. Aspinwall et al., Tool wear and surface integrity observations during the high speed milling of hardened die steel, in: Proceedings of the International Conference on Design and Production of Dies and Moulds, Istanbul, Turkey, (1997).
Temperature measurement when high speed machining hardened mould/die steel R.C. Dewesa,*, E. Nga, K.S. Chuaa, P.G. Newtona, D.K. Aspinwalla,b a
School of Manufacturing and Mechanical Engineering, University of Birmingham, Birmingham, UK IRC in Materials for High Performance Applications, University of Birmingham, Birmingham, UK
b
Abstract When turning and face milling hardened (>30 HRC) mould/die steels, it is necessary to use conventional ceramic or polycrystalline cubic boron nitride (PCBN) tool materials. The use of high speed steel and cemented tungsten carbide (WC) tooling for this application is precluded due to their relatively low hot hardness values. In recent years, however, the mould and die industry has begun to utilise solid WC ball nose end mills for the high speed machining (HSM) of hardened steel cavities. Economic tool lives have been reported at high rotational speeds/feed rates. The paper details the use of 6 mm diameter solid, TiCN coated WC ball nose end mills for the HSM of a hot work tool steel, AISI H13, hardened to 52 HRC. Tests were performed on a Matsuura 20 000 rpm high speed machining centre and temperatures were measured using thermocouple and infrared techniques. Recorded tool /workpiece interface and chip temperatures were relatively low (200±4008C) and increased with higher cutting speed, when using worn rather than new tools and with the workpiece inclined at 608 to simulate ®nishing operations on the side of die cavities. # 1999 Elsevier Science S.A. All rights reserved. Keywords: High speed machining; Temperature measurement; Tool steel; Tungsten carbide
1. Literature survey
1.2. HSM cutting temperatures
1.1. High speed machining
Salomon [1] proposed that there was a peak cutting temperature at an intermediate cutting speed and that when cutting speed was increased from this point, there was a reduction in temperature, see Fig. 1(a). Since this claim, most of the literature has concluded that there is no corresponding reduction in temperature at higher cutting speeds. McGee [7] suggested that temperature increased with cutting speed up to a maximum which was equal to the melting point of the workpiece, see Fig. 1(b). No temperature reduction occurred at higher cutting speeds. This explains why there is no ®xed limit to the cutting speed when machining aluminium alloys (other than that imposed by machine tool considerations). The melting point of these alloys (up to 6608C) is lower than the temperature at which cemented carbide and ceramic tool materials begin to lose their strength and wear rapidly. Conversely, Trent [8] stated that ``it is in the cutting of iron, steel and other high melting point metals and alloys that the heat generated becomes a controlling factor''. Here the limit on cutting speed is a function of the cutting tools used.
High speed machining (HSM) was ®rst reported in 1931 by Salomon [1]. One de®nition is that the process ``involves machining at considerably higher cutting speeds and feed rates than those used in conventional machining'', however, it is most commonly used to describe end milling at high rotational speeds. The process has been adapted to a wide range of applications. In the aerospace sector, HSM is used to remove large volumes of aluminium quickly and to produce thin walled sections in wings [2±4]. One of the more recent applications of HSM is in the manufacture of moulds and dies from hardened tool steels [5,6]. Cavities can be produced from solid in the hardened state using HSM, rather than via the more traditional route: machining in the soft condition followed by electrical discharge machining (EDM), grinding and/or hand ®nishing. *Corresponding author. Tel.: +44-121-414-4175; fax: +44-121-4143541 E-mail address: [email protected] (R.C. Dewes)
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Fig. 1. Graphs of temperature vs. cutting speed: (a) Salomon's curve; (b) McGee's curve.
1.3. Temperature measurement techniques Several experimental methods involving radiation, thermocouple and metallography/microhardness techniques have been used to measure cutting temperatures. A standard thermocouple assembly embedded into a cutting tool or workpiece is known as an implanted, inserted or remote thermocouple [9,10]. On the other hand, if the tool and workpiece are used as the dissimilar metals of the thermocouple this is termed a tool/chip or tool/work thermocouple. The former is a low cost method which can be used to establish the distribution of temperatures at different points in a cutting tool by utilising a series of predetermined locations. Unfortunately, the presence of holes in the tool may alter the temperature distribution and it is not possible to directly measure the temperature at the tool/workpiece interface. The tool/chip thermocouple [11] gives the average temperature, not the maximum, at the interface. Drawbacks of the technique are that both workpiece and tool material must be electrically conductive, cutting ¯uid cannot be used, calibration is laborious, secondary voltages may occur and many tool/workpiece combinations do not form ideal thermocouples [8].
Another technique is to implant a suitable insulated thermocouple wire, Te¯on-coated constantan (55% copper, 45% nickel) wire for example, into a workpiece. When the workpiece is sheared during the machining process, the insulation is broken and an instantaneous hot junction is formed between the wire and the workpiece material in the cutting zone, see Fig. 2. An electromotive force (emf) is generated and assuming calibration of the system has been performed, an interface temperature can be calculated. The advantages of this method include easy calibration and use, and the fact that tool materials which are electrical insulators can be used. Unfortunately, the maximum temperature at the tool/chip interface is not always recorded and experimental error is caused by the variation in detection position along the cutting edge. These problems, however, can largely be overcome by test replications. This technique was used in the present work. A less invasive technique involves measurement of the thermal radiation emitted during the cutting process by using infrared (IR) sensitive photographic ®lm [13] or a pyrometer/infrared thermometer [14]. Fig. 3 shows the system used in the present work. Unfortunately, no cutting ¯uid can be used and the temperature measured is usually that of the top face of the chip and not the interface temperature
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Fig. 2. Machining sequence when using the implanted workpiece thermocouple [12].
Fig. 3. Agema Thermovision 900 system with IR camera (right) and system controller (left).
Fig. 4. Sandvik 1020, 6 mm diameter TiCN coated carbide cutter.
because the tool/workpiece interface is obscured. Errors can also result as a consequence of using an incorrect emisivity value. When using high speed steel or WC tools with an iron/ silicon binder, phase transformations in the tool material can be used to deduce the temperature to which the tool has been subjected and produce isothermal maps [15]. Optical microscopy is utilised to compare sections of the cutting tool with standard specimens. The disadvantages of this method include the limited types of cutting tool materials that can be used and the manufacture of standard specimens which must be very carefully prepared. The aim of the present work was to determine cutting temperatures when high speed ball nose end milling hardened die steel, in order to explain the successful use of tungsten carbide (WC) cutters for this operation. Temperatures were measured using the workpiece/Te¯on-coated constantan thermocouple and infrared techniques previously outlined.
coated with titanium carbonitride, TiCN). Consequently, 6 mm diameter 1020 tools, shown in Fig. 4, were used for the present work. Cutters had four ¯utes with a positive helix angle of 308. A tool overhang from the collet of 30 mm was employed. The workpiece material used was AISI H13 hot work tool steel at a hardness of 52 HRC. It was obtained with a chemical composition of 0.38% C, 1.00% Si, 0.34% Mn, 5.00% Cr, 1.3% Mo and Fe balance. Not surprisingly, AISI H13 is not a standard thermocouple material categorised under British Standard BS1041, therefore, a calibration bar was designed, see Fig. 5. For the machining trials, a workpiece block 80mm55mm35 mm was electrical discharge wire machined into two separate pieces as shown in Fig. 6. The design allowed insertion of thermocouple wires in the gap while giving suf®ciently high clamping force to hold the assembly in place during machining. Two diameters (0.075 and 0.5 mm) of Te¯on-coated constantan wire were obtained from Omega Engineering,
2. Experimental 2.1. Tooling, workpiece materials and machine tools Previous HSM machinability work on hardened steels [16], highlighted the suitability of Sandvik Coromant 1020 (formerly MC45) solid carbide ball nose end mills (PVD
Fig. 5. Calibration bar.
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Fig. 6. Workpiece block.
UK. There are a number of reasons why it is desirable to use thinner wires in the measuring junction during machining. They are poorer conductors of heat which reduces the tendency of heat loss to the environment. Secondly, it is critical to cut the wire cleanly to induce a suf®ciently high thermal emf for measurement. The wire diameter also determines the gap width in between the two parts of the workpiece. A large wire diameter would result in a larger gap causing an interruption in cutting and possibly reduced temperature measurements. Prior to the machining tests, the workpiece and thermocouple wires were prepared. Two 600 mm lengths of 0.5 mm diameter constantan wire were utilised for longer connections between the measuring apparatus and workpiece. One of the wires was silver soldered to the side of the workpiece and the other was soldered to a 40 mm length of 0.075 mm diameter wire. The thinner wire was formed into a sinusoidal wave shape, 10 mm high with 2 mm protruding out of the face of workpiece. Fig. 7 shows the workpiece with the two halves separated in order that the looped wire can be seen. The wire was bonded to the male half of the workpiece using adhesive tape. A Cincinnati milling machine (maximum spindle speed 1500 rpm) and a Matsuura FX-5 CNC controlled high speed machining centre (maximum spindle speed 20 000 rpm) were used for the work, see Fig. 8. The latter machine
Fig. 7. AISI H13 workpiece with constantan wire in place.
Fig. 8. Matsuura FX-5 machining centre.
had a maximum power of 15 kW and a continuously variable feed rate up to 15 m/min. 2.2. Equipment ± thermocouple technique A Fluke 8050A digital multimeter, set at its lowest DC voltage range of 0±200 mV, was used for voltage measurement during initial calibration work. The advantage of this was that it displayed voltage values instantaneously, thus saving time. Unfortunately, the refresh rate on the LCD screen of 0.25 s was a limitation. A Gould 100 Ms/s oscilloscope (DSO) 420 was used for some calibration and the majority of temperature evaluation work. Its advantages included high accuracy, high sampling rate, rapid screen refresh rate and a capability to produce hard copies using its pen plotter. When using the HSM set-up at the operating parameters detailed, the thermal emf generated lasted for a very short period of time, typically 40 ms. By using the oscilloscope, the input signal could be monitored continuously, so that the probability of not capturing a short duration thermal emf signal was minimised. A Kane-May KM1420 temperature data logger was used for the calibration work. This was speci®c to the K-type thermocouple standard (nickel±chromium and nickel±aluminium), as in BS1041. The logger had a measuring range from ÿ100 to 12008C with an accuracy of 0.28C and a response time of <1 s. A stainless steel sheathed K-type probe was connected to the logger and was used to measure the temperature at the measuring junction of the calibration bar. The data in the logger was transferred to a personal computer using the equipment manufacturer's software TERM2UK. The computer enabled the storage of data ®les in ASCII format and allowed the generation of temperature versus time plots on the screen. In order to perform thermocouple calibration, it was necessary for the measuring or hot junction of the calibration bar to be heated while the other end, the reference or cold junction remained at room temperature. Because of this, the furnace calibration method reported by Braiden [17] was not suitable. The induction heating method [18] was also
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machine was not equipped with in®nitely variable table feeds. By looping the thermocouple wire in the manner shown in Fig. 7, each test was effectively replicated a number of times. Tests on the Matsuura machining centre involved both the thermocouple and infrared techniques, see Table 2. All tests were performed using the thermocouple technique, however, due to limited equipment availability, the infrared technique was limited to tests 5 ± 10. Resources did not allow a full factorial experiment to be performed, therefore, it was felt that most information could be gained in the time available by investigating four variables: cutting speed (tests 1 ± 3), tool wear (tests 3 ± 10), angle of the workpiece (tests 3 ± 6) and axial/radial depth of cut. In order to simulate the actual machining of a mould or die cavity, the workpiece was clamped both horizontally and at 608 to the horizontal while maintaining a constant cutting speed by varying the rotational speed. Fig. 9 illustrates the interface cutting speed at workpiece angles 08 and 608 when using a 0.5 mm depth of cut. Table 2 classi®es roughing, semi-®nishing and ®nishing cuts depending on the depth of cut applied. During all tests, dry cutting was employed and feed was ®xed at 0.1 mm/ tooth.
rejected. It was decided that the most suitable heating method was to use an oxygen±acetylene torch. 2.3. Equipment ± infrared technique Equipment consisted of an IR camera and a system controller, see Fig. 3, borrowed from the UK's Engineering and Physical Sciences Research Council (EPSRC). The controller was a dedicated microcomputer set up speci®cally for use with the camera. It consisted of a processor unit, VGA video monitor, mouse and keyboard with specialised system controls for image handling. The software was menu driven and was contained in a windows environment. The camera was designed to operate at wavelengths in the infrared spectrum (8±12 m). The temperature ranges selected for this work were 0±2508C and 100±6008C. The emissivity level used was 0.82. Interchangeable lenses were available for the use with the scanner. In this case, a 408x208 resolution lens was used which had a minimum focal length of 0.3 m. 2.4. Machining parameters In the tests on the Cincinnati milling machine, only the thermocouple technique was used. Fixed parameters included the angle of workpiece (08), down/climb milling, dry machining, use of a tool with 0.3 mm maximum ¯ank wear and an axial/radial depth of cut of 0.5 mm. The effect of feed per tooth and cutting speed were evaluated, see Table 1. Feed per tooth values are approximate because the
2.5. Experimental procedure ± thermocouple technique The calibration experiment involved a number of steps. The 0.075 mm diameter thermocouple wire was silver soldered into hole A on the calibration bar, see Fig. 5. The bar
Table 1 Test parameters on the Cincinnati milling machine Test
Cutting speed (m/min)
Feed per tooth (mm/tooth)
Spindle speed (rpm)
Feed rate (mm/min)
a b c d
16 21 28 21
0.10 0.10 0.10 0.05
870 1140 1500 1140
420 420 570 220 Fig. 9. Effect of workpiece angle on cutting speed.
Table 2 Test parameters on the Matsuura machining centre
Roughing cut
Roughing cut Semi-finishing cut Finishing cut
Test
Maximum flank wear (mm)
Angle of workpiece (degree)
Cutting speed (m/min)
Spindle speed (rpm)
Axial depth of cut (mm)
Radial depth of cut (mm)
Feed rate (mm/min)
1 2 3 4 5 6 7 8 9 10
0 0 0 0.3 0 0.3 0 0.3 0 0.1
0 0 0 0 60 60 60 60 60 60
100 150 200 200 200 200 200 200 200 200
9597 14 396 19 194 19 194 10 610 10 741 10 741 10 741 10 741 10 741
0.5 0.5 0.5 0.5 0.5 0.5 0.2 0.2 0.2 0.2
0.5 0.5 0.5 0.5 0.5 0.5 0.2 0.2 0.1 0.1
3839 5759 7678 7678 4244 4244 4296 4296 4296 4296
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was clamped in place and the thermocouple wires from the measuring and reference junctions were connected to the multimeter. At the measuring junction, the K-type stainless steel probe was inserted into hole B, see Fig. 5, and connected to the data logger. The measuring junction was heated with the oxygen±acetylene torch to 7008C. The voltage on the multimeter was recorded and the bar allowed to cool. During cooling, at ®xed temperature intervals, the emf value was recorded. The values were plotted on a graph of voltage versus temperature. A linear relationship between voltage and temperature was established and the equation of the line determined. The calibration procedure was repeated ®ve times and ®ve calibration lines superimposed. The equation of a line of best ®t was determined. One calibration test using the oscilloscope rather than multimeter was carried out in order to con®rm that both sets of equipment gave similar results. When calibration was complete, initial tests were undertaken on the Cincinnati milling machine. The workpiece/ thermocouple junction was set-up as discussed previously. The assembly was clamped in a vice with plywood used to insulate it from the machine to prevent electrical interference. Machining cuts were taken across the workpiece using the parameters given in Table 2 until the protruding thermocouple wire was cut and a signal triggered in the oscilloscope. The voltage was printed using the pen plotter. When the entire face of the workpiece had been machined and a number of temperature readings obtained, the workpiece was re-prepared. The voltages were analysed and converted to temperatures. A mean temperature value was calculated and taken as the interface temperature for the test. All subsequent work was performed on the Matsuura machining centre. Each time a new workpiece was clamped in the vice, it was levelled to 30 m. Trigonometric calculations were used to check the level of the workpiece at 608. The voltage and timebase settings on the oscilloscope were modi®ed depending on the cutting parameters, in order to produce reliable results. A CNC program was executed which resulted in the cutter machining across the workpiece, retracting above the workpiece and returning to the start of the next cut and applying the radial depth of cut. Each time a wire was machined through, a signal occurred and the voltage trace was recorded and printed out as with the tests on the Cincinnati machine. 2.6. Experimental procedure ± infrared technique The Agema Thermovision 900 system was set up, the temperature range selected and the level and span of the image adjusted to produce a recognisable image. The IR camera was ®lled with liquid nitrogen, positioned at a safe distance from the workpiece to avoid any possibility of damage by the swarf and connected to the system controller. The automatic image storing sequence (1 s intervals) was initiated, see Fig. 10. The machining sequence was set up as described previously with the cutting
Fig. 10. A thermal image captured when machining workpiece at 608.
parameters given in Table 2. After the sequence was recorded, the images were recalled and analysed by using the ``spotmeters'' in the software. When placed over the image, these gave a temperature value at a particular point. The maximum chip temperature was therefore recorded for each test. 3. Results and discussion 3.1. Calibration of the thermocouple The ®ve sets of calibration results using the multimeter were combined to give a line of best ®t, see Fig. 11 for graph and equation of the line. Also shown are the results and equation obtained when using the oscilloscope. The difference in readings between the two measuring devices was only 0.14% at 20 mV. Therefore, the equation from the oscilloscope was used to calculate the temperature from the induced thermal emf for the machining tests. 3.2. Temperature results The effect of cutting speed on interface temperature, as measured using the thermocouple technique can be seen in Fig. 12 which shows the results of Cincinnati tests a ± c and Matsuura tests 1 ± 3. Mean temperature values are shown. An analysis indicated that mean and median temperatures fell within 108C in all cases. It can be seen that over the range of cutting speeds tested, the interface temperature increased in a manner similar to that found by McGee when machining aluminium [7]. This is in line with conventional metal cutting theory [19]. A machine with higher spindle speeds would be required to fully test Salomon's ®ndings that temperature reduces at higher cutting speeds [1]. Although not shown graphically, reducing the feed rate from 0.10 to 0.05 mm/tooth at a cutting speed of 21 m/min
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Fig. 11. Calibration results and equations for multimeter and oscilloscope.
gave a very slight reduction in temperature from 1108C to 1078C. The next stage of the work was to test the effect of workpiece angle and tool wear on interface temperature, see Fig. 13. There was an increase in temperature when machining with the workpiece at 608 rather than 08. This can
be explained because at 08 the cutting speeds on the ball nose end mill ranged from zero in the centre to 200 m/min at the periphery, while at 608 the range is from 173 to 200 m/min, see Fig. 9. As a result, higher average cutting speeds occurred in the cutting zone and therefore, higher temperatures would be expected. An effect which is likely to have prevented temperatures at 08 from being even lower was the nature of the ball nose cutter. At 608, all four teeth of the cutter would have done an equal amount of cutting, however, at 08, two of the teeth would have done very little work due to the fact that they stop short of the centre of the tool. The effective feed per tooth would have approached 0.2 mm/ tooth which would be expected to contribute to higher temperatures. When using a tool with 0.3 mm of ¯ank wear, temperatures increased. With a worn tool, more surface contact between the tool and workpiece occurred and the increase in sliding friction and rubbing caused higher temperatures. Figs. 14 and 15 show the effect of axial and radial depth of cut on interface and chip temperatures, respectively. Roughing, semi-®nishing and ®nishing cuts are indicated as R, SF
Fig. 13. The effect of workpiece angle and tool wear on interface temperature.
Fig. 14. Effect of axial and radial depth of cut on interface temperature.
Fig. 12. Effect of cutting speed on interface temperature.
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4. Conclusions
Fig. 15. Effect of axial and radial depth of cut on chip temperature.
and F. In all but one case, chip temperatures are much lower than interface temperatures. Values measured by the infrared technique were typically one third those from the thermocouple. It is widely accepted that during metal cutting, the majority of heat (80%) is transferred to the chip, therefore, it could be argued that this is a larger difference than would be expected. A possible explanation is that with the IR camera and lens used, the hot zone was relatively small compared to the total area scanned, see Fig. 10. This, combined with limitations in the software which would allow the use of ``spotmeters'' but not output a maximum temperature over an area, meant that there was no guarantee that the hottest point was measured. It is considered that the thermocouple technique was intrinsically superior for this application and gave more accurate, absolute temperature values. The differences in temperature between the three types of cut are as would be expected, with smaller depths of cut giving signi®cantly lower temperatures. This would validate the use of small depths of cut during the HSM of hardened steels in order to minimise cutter wear. With the roughing cut, there is a large difference between the temperatures generated with the new and worn tools, a smaller difference when semi-®nishing and an even smaller range when ®nishing. This can be explained by reference to the extent of ¯ank wear on the worn tools, see Table 2, higher levels of ¯ank wear giving a bigger difference in temperature. Currently, EDM is widely used for the production of moulds and dies from hardened steels. Due to the high temperatures which occur, typically 10 000 ± 20 0008C, surface integrity problems such as heat affected zones, white layers and tensile residual stresses are found on machined components. With the temperatures measured using the thermocouple technique (<4008C), it would be expected that surface integrity problems would be minimised when using HSM, as the tempering temperature of AISI H13 (5508C) is not exceeded. In parallel work, some of which has been published [20], the authors found that HSM gave superior surface integrity to EDM.
1. Interface temperatures measured using the thermocouple technique when machining with the workpiece at 08 were 198±3018C. Higher temperatures, 247±3858C, occurred at a 608 workpiece angle. 2. Temperature increased with cutting speed and no reduction at higher speeds occurred. This is contrary to Salomon's theory and agrees with McGee's findings. Feed per tooth did not have a large effect on temperature. 3. Machining with a worn tool generated higher temperatures than when using new tools. 4. The IR technique indicated lower temperatures than the thermocouple method with values of 68±3908C. Heat losses to the tool and workpiece and drawbacks with the IR technique explain the lower temperatures measured. 5. The relatively low cutting temperatures measured explain why WC products can be successfully used for the HSM of hardened steels. Acknowledgements We would like to thank Prof. A.A. Ball, Head of the School of Manufacturing and Mechanical Engineering, and Prof. M.H. Loretto, Director of the IRC in Materials for High Performance Applications, for provision of facilities and funding. We are grateful to Messrs. R. Fasham, W. Hewitt, P. Thornton and J. Wedderburn for their assistance with the experimental work. Thanks also go to the Control, Design and Production Directorate of EPSRC, De Beers Industrial Diamond Division (Pty) Ltd., Dynacast (UK) Ltd., Kieninger Tooling Ltd., Matsuura Machinery plc, Sandvik Coromant (UK) Ltd., United Engineering Forgings Ltd., and WH Smith and Sons (Tools) Ltd., for additional funding and support. We also wish to express our gratitude to the EPSRC for lending the infrared equipment.
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R.C. Dewes et al. / Journal of Materials Processing Technology 92±93 (1999) 293±301 [9] M.P. Groover, G.E. Kane, A continuing study in the determination of temperatures in metal cutting using remote thermocouples, J. Eng. Ind. (Trans. ASME) 93(2) (1971) 603±608. [10] A.H. Redford, B. Mills, S. Akhtar, Temperature - tool life relationships for resulphurised low carbon free machining steels, CIRP Annals 25(1) (1976) 89±91. [11] P. Lezanski, M.C. Shaw, Tool face temperatures in high speed milling, J. Eng. Ind. (Trans. ASME) 112(2) (1990) 132±135. [12] W. Chen, A report of the research on surface integrity of hardened steel following face milling, School of Manuf. and Mech. Eng., University of Birmingham, UK, (1994). [13] G. Boothroyd, Temperatures in orthogonal metal cutting, Proc. Instn. Mech. Engrs. 177(29) (1963) 789±802. [14] J.P. Kottenstette, Measuring tool-chip interface temperatures, J. Eng. Ind. (Trans. ASME) 108(2) (1986) 101±104. [15] P.A. Dearnley, E.M. Trent, Wear mechanisms of coated carbide tools, Metals Technology, 9 (1982) 60±75.
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[16] R.C. Dewes, D.K. Aspinwall, High speed machining of hardened tool steel using cemented carbide tooling, in: Proceedings of the Third International Conference on Progress of Cutting and Grinding, Osaka, Japan, vol. 3, (1996) pp. 217±222. [17] P.M. Braiden, The calibration of tool/work thermocouples, Proc. in: Proceedings of the Eighth International MTDR Conference, Manchester, UK, Perganon, Oxford, (1967), pp. 653±666. [18] G. Barrow, A review of experimental and theoretical techniques for assessing cutting temperature, CIRP Annals 22(2) (1973) 203. [19] B.T. Chao, K.J. Trigger, An analytical evaluation of metal cutting temperatures, Trans. ASME, 73, (1951) p. 57. [20] R.C. Dewes, D.K. Aspinwall et al., Tool wear and surface integrity observations during the high speed milling of hardened die steel, in: Proceedings of the International Conference on Design and Production of Dies and Moulds, Istanbul, Turkey, (1997).