Workpiece surface integrity when slot milling -TiAl intermetallic alloy Hood, Richard; Aspinwall, David; Soo, Sein; Mantle, Andrew L.; Novovic, Donka DOI: 10.1016/j.cirp.2014.03.071 Citation for published version (Harvard): Hood, R, Aspinwall, DK, Soo, SL, Mantle, AL & Novovic, D 2014, 'Workpiece surface integrity when slot milling TiAl intermetallic alloy' CIRP Annals - Manufacturing Technology, vol 63, no. 1, pp. 53-56., 10.1016/j.cirp.2014.03.071 Link to publication on Research at Birmingham portal General rights When referring to this publication, please cite the published version. Copyright and associated moral rights for publications accessible in the public portal are retained by the authors and/or other copyright owners. 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Oct. 2014 CIRP Annals - Manufacturing Technology 63 (2014) 53–56 Contents lists available at ScienceDirect CIRP Annals - Manufacturing Technology jou rnal homep age : ht t p: // ees .e lse vi er . com /ci r p/ def a ult . asp Workpiece surface integrity when slot milling g-TiAl intermetallic alloy Richard Hood a, David K. Aspinwall (1)a,*, Sein Leung Soo (2)a, Andrew L. Mantle (3)b, Donka Novovic c a b c Machining Research Group, School of Mechanical Engineering, University of Birmingham, Edgbaston, Birmingham, UK Manufacturing Technology, Rolls-Royce plc, Derby, UK Turbines, Rolls-Royce plc, Derby, UK A R T I C L E I N F O A B S T R A C T Keywords: Titanium Surface integrity Slot milling Slot milling is presented as a potential manufacturing route for aerospace component feature production when machining g-TiAl intermetallic alloy Ti–45Al–2Mn–2Nb + 0.8 vol.% TiB2XD using 2 mm diameter AlTiN coated WC ball nose end milling cutters. When operating with flood cutting fluid at v = 88 m/min, f = 0.05 mm/tooth, d = 0.2 mm, maximum flank wear was 65 mm after 25 min. SEM micrographs of slot surfaces show re-deposited/adhered and smeared workpiece material to a length of 50 mm. Brittle fracture of the slot edges was restricted to <10 mm with sporadic top burr formation observed up to 20 mm. Cross sectional micrographs of the slot sidewalls showed bending of the lamellae limited to within 5 mm. ß 2014 CIRP. 1. Introduction Manufacturing components from gamma titanium aluminide (g-TiAl) intermetallic alloys with features that have ‘acceptable’ integrity is extremely difficult. This is mainly due to their low room temperature ductility, which is quoted as being <2%. Despite this, they are seen as potential replacement materials for nickel based superalloy turbine blades, not least because alloy density is typically 4 g/cm3,which is around half that of nickel alloys and their operating ceiling extends up to 850 8C [1]. Milling is likely to be an important operation in the manufacture of blades and a significant proportion of published machinability research on these alloys over the last decade has focused on this process [1–6]. Such studies have investigated the effects of operating parameters and conditions on tool life and surface integrity when milling g-TiAl alloys. The majority of cutters assessed however have been >6 mm in diameter [2–6], with the workpiece geometric shape simplified to rationalise testing, although Beranoagirre et al. [2] does detail the machining of a sample blade component but without disclosing the resulting workpiece surface integrity. Additionally, while publications on the micro milling of alpha/beta alloy Ti–6Al–4V can be found [7], equivalent data for g-TiAl in respect of micro/mesoscale blade feature production is sparse. The current research attempts to address these aspects. 2. Experimental work All milling tests were carried out on a Matsuura LX1, 3-axis high speed machining centre rated at 200–60,000 rpm, with a maximum * Corresponding author. http://dx.doi.org/10.1016/j.cirp.2014.03.071 0007-8506/ß 2014 CIRP. rapid feed rate of 90 m/min and a maximum spindle power of 5 kW. Cutting tools were held in MST Mizoguchi HSK shrink fit tool holders. Where required, Houghton Vaughan Hocut 3380 cutting fluid was supplied at a flow rate of 6 l/min. Trials in Phases 1 and 2 used two cast workpiece ingots which were 200 mm 100 mm 25 mm in size with composition, Ti–45Al– 2Mn–2Nb + 0.8 vol.% TiB2. These had a grain size of 200 mm and bulk hardness of 330–350 HV and were surface ground on all sides. Phase 3 work involved machining of representative cast component features on similar but proprietary workpiece material and only involved finishing operations. Tool wear was measured using a Wild microscope and a toolmakers table with digital micrometre heads giving a resolution of 0.001 mm. Images of tool wear were taken using a Canon digital SLR camera. Workpiece surface roughness was measured using a Taylor Hobson Form Talysurf 120L and an Alicona IF G4 microscope system. For workpiece surface integrity evaluation, selected samples were hot mounted in edge retentive bakelite then prepared using a grinder/polisher. Workpiece surfaces and cross-sections were analysed using a Leica optical microscope running Buehler Optimet software or a JEOL 6060 scanning electron microscope (SEM) to assess the level of damage. Residual stress measurements were undertaken using the blind hole drilling technique. Difficulties can be encountered when using X-ray diffraction techniques with g-TiAl due to the material’s tetragonal phase and high levels of lattice strain, leading to broad, overlapping reflections (line broadening) [4]. Each target site was prepared by light abrasion (two passes of 400 grade SiC paper) followed by thorough degreasing (acetone). Strain gauge rosettes type EA-06062RE-120 (Vishay Precision), were bonded to the workpiece surface using Loctite 407 adhesive. Each rosette was drilled using a PC-controlled 3-axis drilling machine. Fig. 1 shows the experimental setup. 54 R. Hood et al. / CIRP Annals - Manufacturing Technology 63 (2014) 53–56 Table 1 Operating parameters and conditions for Phase 1 tests. Fig. 1. Schematic of residual stress measurement. Depth increments were set at 32 mm (four times), 64 mm (four times) and 128 mm (eight times), giving a completed hole depth of 1.41 mm for the determination of stresses up to 1024 mm below the machined surface. Results from the individual target gauges were obtained in the form of relaxed strains, which were subsequently converted to stress values. Fig. 2 shows schematics/images of the coated cutting tools supplied by Seco Tools (UK); WC XMG (WC 90% Co 10%, grain size 0.8 mm). Six different tool formats were used associated with 5 basic types (denoted Type 1 to Type 5), all of which were 2 mm in diameter and available as stock items (except Type 5). As limited data was available for small diameter end milling of g-TiAl alloys, it was decided to test a variety of possible tools. The Type 1 tool was a 2 flute ball nose end mill with a Mega 64 (AlTiN-Monolayer) coating. The cutting rake was 08, helix angle was 178 and the cutting/flute length was 2 mm. Two different tool subtypes; 1(a) and 1(b) were used with the former having a maximum cutting depth of 4 mm while for the latter it was 10 mm. Fig. 2. Cutting tool types. Tool Type 2 was similarly a 2 flute ball nose end mill with the Mega coating (AlTiN) but having a cutting rake and helix angle of 108 and 308 respectively, while the cutting length and maximum cutting depth was 9 mm. The Type 3 cutter was a 2 flute high feed end mill with a Mega coating having a corresponding cutting length of 0.15 mm and maximum cutting depth of 6 mm. Tool Type 4 was a 4 flute end mill with a Mega coating. The cutting rake was 28, helix angle was 208 while the cutting length and maximum cutting depth were 2.5 and 5 mm respectively. The Type 5 tool was a bespoke design produced specifically for the current work, which was based on the Type 1 ball nose end mill with a cutting length and maximum cutting depth of 9 mm. This extended cutting/flute length was designed to reduce material smearing. Experimental work involved 3 phases, which employed down milling unless otherwise stated. In Phase 1, research was aimed at determining tool life and workpiece integrity performance when machining the sidewall of a slot under various finishing conditions. Table 1 details the variable test parameters utilised with the radial depth of cut/stepover fixed at 0.2 mm. The levels chosen were based on tool manufacturer recommendations and previous research using larger diameter tools [4]. Tests 1 and 2 were performed using Type 1(a) tools to compare dry and flood coolant operation. Tests 3 and 4 were carried out to investigate the effect of tool geometry (Type 1(b) vs. 2) on tool life, with operating parameters selected based on results from Tests 1 and 2. Tests 5 and 6 investigated the effect of cutting speed to determine the maximum possible value before tool failure/heavy wear occurred using Type 3 cutters. Test Tool type Cutting speed (m/min) 1 2 3 4 5 6 7 8 9 10 1(a) 1(a) 1(b) 2 3 3 4 4 4 4 55 55 88 88 55–160 160 88 88 88 88 Feed rate (mm/tooth) Axial depth of cut (mm) Cutting environment 0.01 0.01 0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.05 0.1 0.1 0.2 0.2 0.2 0.2 0.1–1.0 0.05 0.2 0.05 Dry Wet Dry Dry Dry Dry Dry Wet Wet Wet Test 7 involved tool Type 4 to investigate the maximum depth of cut before tool breakage. Here, varying axial depths of cut from 0.1 up to 1.0 mm were employed with the maximum machining depth of 1.0 mm. Test 8 provided an assessment of cycle time and compared workpiece surface integrity when using three different milling strategies namely single direction rasters; one providing down milling, the other up milling and conventional raster operation giving alternate up and down milling scenarios. Operating parameters were similar to those in Test 7 but with flood coolant and a 0.05 mm axial depth of cut. Residual stress measurements were undertaken on surfaces machined using Test 8 parameters with a raster milling strategy along with a 0.2 (Test 9) and 0.05 mm (Test 10) depth of cut. Financial constraints limited additional measurements, the two tests chosen giving ‘preferred’ parameters for workpiece integrity in Phase 1 work. Phase 2 work involved the milling of an ‘end slot’ from solid (roughing). The slot was 2.5 mm wide, approximately 12 mm long and 9 mm deep. Several tests (not all reported here) were performed using cutting parameters based on Phase 1 results to determine their effect on tool life and workpiece integrity. Fig. 3 shows a schematic with cutter paths developed using Delcam PowerMill software (Fig. 3(a)) and a section of the machined workpiece (Fig. 3(b)). A profile cutter path was used. Tests 1 and 2 compared the number of end slots machined against flank wear for Types 1(b) and 3 tools using a cutting speed of 88 m/min and a depth of cut of 0.05 mm with flood coolant. The feed rate for the Type 1(b) tool was 0.05 mm/tooth while for the high feed end mill (Type 3) it was and 0.1 mm/tooth. Fig. 3. Phase 2 workpiece (a) CAM model, (b) workpiece. As outlined previously, Phase 3 research involved machining of representative features on ten cast workpieces (finishing only). Commercial restrictions preclude detailing the exact geometry however workpiece surface integrity data is presented. Machining parameters were based on the preferred values from Phases 1 and 2, including a cutting speed of 88 m/min, a stepover of 0.2 mm and flood coolant. Variable parameters were feed rates of 0.025 and 0.05 mm/tooth along with axial depths of cut of 0.05, 0.1 and 0.2 mm. A new Type 5 tool was used for each workpiece. 3. Results and discussion Phase 1 experimental results for Tests 1, 2, 3, 4 and 6 are shown in Fig. 4. Tool life for Tests 1 and 2 was generally good with a R. Hood et al. / CIRP Annals - Manufacturing Technology 63 (2014) 53–56 Fig. 4. Tool life curve for Phase 1 sidewall milling. machining time in excess of 100 min using Test 1 operating parameters. Plotted data are the averages for the maximum flank wear values on individual teeth. Results for Tests 1 and 2 were comparable until a machining time of 60 min after which Test 2 (involving use of flood coolant) showed a higher wear. Given the limited literature available on slot milling of titanium alloys with small diameter tools, the parameter levels selected for these tests were conservative. Test 3 showed a higher rate of tool wear, yet despite the 60% increase in cutting speed (55–88 m/min), the average maximum flank wear level was only 60 mm after 25 min. In contrast, the cutter in Test 4 fractured on first contact with the workpiece, which is not an uncommon occurrence when micro milling titanium alloys [7]. The rake angle for the Type 2 cutter used in this test was +108 compared with 08 for Test 1–3 tools. Priarone et al. [3] describes a similar result with a 50% increase in tool life when comparing tooling with a radial rake angle of +128 and +48. Test 5 showed a maximum cutting speed of 160 m/min was achievable and therefore Test 6 was performed as a life trial using this value. Tool life was however extremely short, with a machining time of only 30 s for an average maximum flank wear level of 200 mm. Wear scar photographs from Test 1 and Test 6 at cessation are also shown in Fig. 4. Tool breakage did not occur even when using a depth of cut of 1 mm with the Type 4 tool in Test 7 and five passes were achieved before the test was stopped. Test 8 showed only marginal differences in surface/subsurface damage after machining in either a single raster up or down milling mode, or when raster milling with alternate up/down operation. No cracking was observed and bending of lamellae was restricted to <4 mm. Consequently, the reduced cycle time associated with conventional raster operation was deemed preferable. Residual stress depth profiles for Tests 9 and 10 (new tools) are detailed in Fig. 5. Both trials showed similar trends with a maximum compressive stress of between 400 and 527 MPa at the surface, which became neutral at a depth of 450 mm. The lower depth of cut for Test 10 produced a reduction (25%) in residual stress for each measurement. In both tests, the residual stress parallel to the feed was 10% higher (more compressive) than that measured perpendicular to the feed direction. Measurement Fig. 5. Residual stress depth profile measurements. 55 uncertainty was 85 MPa at a depth of 16 mm, 19 MPa at a depth of 512 mm and 23 MPa at a depth of 1024 mm. These errors consist of a random strain uncertainty of 3 me and a linear error of 7% of the total stress level [8]. Published residual stress data for ball end milling g-TiAl using 6 mm diameter cutters, details compressive residual stresses ranging in magnitude from 600 to 1200 MPa [4]. The lower levels with the current work reflect the less arduous operating parameters used, yet both sets of results concur that the stress became neutral at a distance of 0.3–0.5 mm below the workpiece surface. In general, cutting temperatures when ball nose end milling g-TiAl alloys are low, of the order of 200–300 8C for a cutting speed 70–120 m/min, feed rate 0.12 mm/tooth and depth of cut (axial and radial) 0.2–0.5 mm [5]. Operating regimes subject to low cutting temperatures coupled with a high strain hardening rate, typically produce compressive stresses which are likely to produce longer fatigue life as the peak tensile load at the surface would be reduced and surface cracks would tend to close [4]. Workpiece surface roughness (Ra) measurements performed on Test 9 and 10 samples showed a range in values from 0.8 to 1.2mm. Phase 2 work machining slots from solid as detailed in Fig. 3, produced mixed results. For the high feed tool (Type 3), 25 slots were machined before the average maximum flank wear level of 200 mm was reached, while the ball nose end mill (Type 1(b)) produced 10 slots before tool edge fracture occurred and the test halted. Fig. 6 shows the wear progression with a wear scar photograph for the ball nose end mill trial at test cessation. Fig. 6. Tool life curve for Phase 2 solid slot milling. Images of the machined workpiece surface from Phase 2 are shown in Fig. 7. The surfaces were covered extensively in smeared/ adhered workpiece material with a length of up to 50 mm, see Fig. 7(a and b) for the slot mid point and closed end respectively. Possible cracks with a length of up to 100 mm, were also found as shown in Fig. 7(c). It was often difficult to ascertain if these were cracks in the bulk workpiece or adhered material. Additionally edge fracture was observed, see Fig. 7(d), despite the low level of operating parameters, edge profiling would be required in practice Fig. 7. SEM images of the Phase 2 machined workpiece surface. 56 R. Hood et al. / CIRP Annals - Manufacturing Technology 63 (2014) 53–56 to minimise this. Results showed considerably worse surface integrity than found in Phase 1. Scanning electron microscope images of the cast workpiece surfaces machined as part of Phase 3 testing showed surface defects on nearly every cast feature. These were less severe than those detailed for Phase 2 work and were consistent with Phase 1 finishing investigations. The damage consisted of tearing of the surface and fracture of the milled/cast surface interfaces as shown in Fig. 8. In general, the size and incidence of the tearing increased as the level of operating parameters increased. As with the residual stress results, the level of workpiece surface damage was somewhat less than that reported in previous research with larger cutters and higher material removal rates [6], particularly with respect to fracture/pullout. Having said this, a large amount of smeared/adhered workpiece material was observed in the SEM images, which may have obscured any surface damage, see Fig. 8(a and d). similar cracks were seen. It is likely therefore that they were a byproduct associated with the casting process. Not surprisingly, tests performed under more arduous conditions which included a depth of cut of 0.1 mm and feed rate of 0.05 mm/tooth, showed greater bending of the lamellae and an ‘arc’ shaped crack (6 mm deep and 80 mm wide) was observed as shown in Fig. 9(c). Cross-sectional analysis of the cast/milled surface interfaces showed mixed results. For some cast features there was limited damage (<10 mm), whereas for others more damage was observed. Sporadic top burr formation as shown in Fig. 9(d) occurred on several cross-sections with dimensions of up to 20 mm. Burr formation when machining brittle materials like gTiAl is relatively rare but presents an additional problem requiring an edge profiling solution. However, the sporadic nature of the formation is likely to preclude the use of a prediction system to modify the design to avoid burrs [9]. 4. Conclusions Fig. 8. SEM images of the machined surface of the cast workpieces. The majority of workpiece cross-sections produced following machining of the cast features, also showed some degree of damage; see Fig. 9. Where visible this consisted of bending of the lamellae to an average depth of <5 mm and a maximum depth of 15 mm as illustrated in Fig. 9(b). Cracks (100 mm long and 4 mm wide) between lamellae were observed in two locations within 0.5 mm of each other on workpiece 3 (from the batch of 10). This was the only time that cracks of this nature were observed, indeed, in two further replications using identical operating levels no Fig. 9. Cross-sectional micrographs for Phase 3 investigation. Workpiece surface integrity characteristics including residual stress data, have been established for the slot milling of intermetallic alloy Ti–45Al–2Mn–2Nb + 0.8 vol.% TiB2 when using small diameter end milling tools. Such data is a prerequisite in identifying appropriate strategies for the manufacture of aeroengine components. Residual stress measurement using the blind hole drilling technique showed compressive residual stresses of 527 MPa which became neutral at a depth of 450 mm. Slot surfaces showed re-deposited/adhered and smeared workpiece material to a length of 50 mm. Brittle fracture of the slot edges was restricted to <10 mm with sporadic top burr formation observed up to 20 mm. Slot sidewall micrographs detailed bending of the lamellae limited to within 5 mm when using finishing conditions. Slots machined from solid showed the worst surface integrity of all the features machined. Acknowledgments We would like to thank Rolls-Royce plc and Seco Tools (UK) Ltd. for both financial and technical support. References [1] Aspinwall DK, Dewes RC, Mantle AL (2005) The Machining of g-TiAl Intermetallic Alloys. CIRP Annals – Manufacturing Technology 54(1):99–104. [2] Beranoagirre A, Olvera D, Lopez de Lacalle LN (2012) Milling of Gamma Titanium Aluminium Alloys. International Journal of Advanced Manufacturing Technology 62:83–88. [3] Priarone PC, Rizzuti S, Settineri L, Vergnano G (2012) Effects of Cutting Angle, Edge Preparation and Nano-Structured Coating on Milling Performance of a Gamma Titanium Aluminide. Journal of Materials Processing Technology 212:2619–2628. [4] Mantle AL, Aspinwall DK (2001) Surface Integrity of a High Speed Milled Gamma Titanium Aluminide. Journal of Materials Processing Technology 118:143–150. [5] Aspinwall DK, Mantle AL, Chan WK, Hood R, Soo SL (2013) Cutting Temperatures when Ball Nose End Milling g-TiAl Intermetallic Alloys. CIRP Annals – Manufacturing Technology 62(1):75–78. [6] Hood R, Aspinwall DK, Sage C, Voice W (2013) High Speed Ball Nose End Milling of g-TiAl Alloys. Intermetallics 32:284–291. [7] Thepsonthi T, Ozel T (2013) Experimental and Finite Element Simulation based Investigations on Micro-Milling Ti–6Al–4V Titanium Alloy: Effects of cBN Coating on Tool Wear. 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