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Machinability and surface integrity of RR1000 nickel based superalloy
CIRP Annals - Manufacturing Technology 60 (2011) 89–92
Contents lists available at ScienceDirect
CIRP Annals - Manufacturing Technology jou rnal homep age : ht t p: // ees .e lse vi er. com/ci rp/ def a ult . asp
Machinability and surface integrity of RR1000 nickel based superalloy S.L. Soo a, R. Hood a, D.K. Aspinwall (1)a,*, W.E. Voice b, C. Sage b a b
Machining Research Group, School of Mechanical Engineering, University of Birmingham, Birmingham, UK Manufacturing Technology, Rolls-Royce Plc., UK
A R T I C L E I N F O
A B S T R A C T
Keywords: Nickel alloy Drilling Surface integrity
Alloy development for gas turbine components has produced materials able to maintain strength and integrity at operating temperatures up to 1050 8C. Next-generation RR1000 nickel-based superalloy reflects this philosophy, albeit at the expense of machinability. Experimental data for drilling showed flank wear when operating at 45 m/min to be <100 mm for a distance cut of 1800 mm (150 holes). Thrust forces measured 1600–1800 N. Re-deposited material on hole surfaces and drag/distortion was evident on cross-sectional micrographs. Roughness of end-milled specimens was <0.8 mm Ra with minimal damage using new tools, however significant burring/increased microhardness (150 HK0.05) and white layer formation occurred when employing worn tools. ß 2011 CIRP.
1. Introduction The word ‘superalloys’ defines a group or class of materials which are able to operate over an extended period of time at temperatures above 700 8C, without significant loss of their mechanical and physical properties, particularly strength and corrosion resistance. Typically such alloys employ a high proportion (weight %) of nickel, iron, cobalt and chromium in combination with up to 14 other elements and are regarded as amongst the most highly developed alloys in existence. One classification identifies alloys according to these major constituents such that materials are described as iron-base, cobalt-base and nickel-base, with the latter containing 40–75% Ni. The nickel-base alloys have outstanding oxidation resistance and exceed stainless steels in their high temperature strength [1]. Understandably material and processing costs are significant, irrespective of machining costs which are generally up to 5 that of more conventional alloys due in part to poor tool life, low productivity and the need for acceptable workpiece integrity. The development and growth in superalloys is due principally to the advent of the gas turbine engine in the late 1940s and the subsequent worldwide expansion of the aerospace/aeroengine industry. Application is however, not restricted solely to engines, whether for aircraft or land based power generation systems, encompassing as it does equipment for nuclear reactors, chemical and petrochemical installations, medical/prosthetic devices, rocket engine parts, cryogenic applications, etc. Currently some 50% of aeroengine weight is due to superalloy use, and such materials are particularly prevalent in the combustor and turbine sections where operating temperatures of 1100–1250 8C are typically experienced. It is unlikely that this percentage will increase due to parallel developments in lighter titanium materials with
* Corresponding author. 0007-8506/$ – see front matter ß 2011 CIRP. doi:10.1016/j.cirp.2011.03.094
increased temperature capability (up to 750/800 8C) and greater use of composites. Possible future use of intermetallics such as gTiAl for blades and stators in the HP compressor and blades in the LP turbine section which are currently made from heavier nickel superalloys, are estimated to provide significant weight savings. Despite such developments, the need to provide higher levels of engine fuel economy, reduced NOx emissions and noise, points to increased operating temperatures (higher compressor discharge and turbine entry temperatures) and consequently the use of more advanced nickel alloys, which are able to cope with higher temperatures and stresses than are currently experienced [2]. Fig. 1 details approximate operating temperatures and associated material usage in a typical aeroengine.
[()TD$FIG]
Fig. 1. Aeroengine temperatures and material use. Adapted from [3,4].
From the perspective of publications on the machinability of superalloys, Inconel 718 has over several decades, been the main focus of attention. This is similarly true of the a/b alloy Ti–6Al–4V which has dominated machining publications dealing with titanium alloys. In fairness both have played significant roles in aeroengine development and continue to find application for blades and discs, however over recent years newer nickel
90
S.L. Soo et al. / CIRP Annals - Manufacturing Technology 60 (2011) 89–92
superalloys such as Udimet 720 and RR1000 together with titanium alloys such as Ti–6Al–2Sn–4Zr–6Mo have become increasingly important. Key texts by Bradley [1] and more recently Reed [5] list 100 of the more common wrought (polycrystalline) and cast superalloys, although a number of these have effectively been superseded, on the cast alloy front by improved 2nd, 3rd, and 4th generation compositions. For applications such as turbine discs some 25 current nickel based alloys are cited [5,6]. The wrought alloys encompass solid solution compositions and those capable of precipitation strengthening, while cast materials comprise conventionally cast, direct solidification and single crystal alloys. The increasing emphasis on powder metallurgy processing, followed by forging or alternatively hot isostatic pressing (HIP), in order to improve mechanical properties or achieve greater economy through near net shape manufacture, is an additional consideration. The following sections detail results of twist drilling and radius end milling of powder HIP-ed RR1000. As powder processed forged material, it is finding increased use for turbine discs, but is thought more difficult to machine than mainstream products and hence the new benchmark in relation to superalloy machinability. 2. Experimental work 2.1. Workpiece material format, tooling and equipment Workpiece material was initially supplied in the form of a bar with a bulk hardness of 450 HV30. Several discs measuring 90 mm diameter 12 mm thick as well as strips with dimensions of 100 mm 18 mm 12 mm (used for subsequent hole quality and integrity analysis) were cut from the bar for drilling trials while similarly, plates of 100 mm 32 mm 6 mm were produced for radius end milling tests. Specimens were held in suitably designed clamping fixtures and mounted on appropriate piezoelectric dynamometers for each process. Experiments were carried out on a FGC1000 flexible grinding centre with a maximum spindle rotational speed of 6000 rpm and power rating of 25 kW. All tests were performed wet using a water based mineral oil emulsion; Hocut 3380 at 7–10% concentration. When drilling an adaptor was retrofitted on the main spindle to enable through tool fluid delivery at a pressure of 50 bar and flow rate of 6 l/min. When end milling cutting fluid was supplied externally using a twin nozzle arrangement at 70 bar and 26 l/min. Drills were 8 mm diameter, twin fluted, TiAlN coated solid carbide Delta C ISO-S tools, while the end milling cutter employed was a 20 mm diameter indexable Coromill 300 unit employing 10 mm round, Al(TiN) coated carbide inserts. Tool wear was measured relative to the datum junction of the primary/secondary clearance (Fig. 2) using a microscope and table equipped with digital micrometers giving a resolution of 0.001 mm. Hole diameter was evaluated using a Talyrond 300, while surface roughness was measured with a Form Talysurf 120 L using the appropriate metrology standards. Microhardness measurements of the workpiece (depth profile), were taken using a suitable hardness testing machine at a load of 50 g and 15 s indent duration. For workpiece surface integrity evaluation, machined workpiece surface samples were hot mounted in edge retentive bakelite followed by grinding, polishing and etching according to standard procedures for nickel based superalloy materials. Selected machined workpiece surfaces and cross-sections were analysed using a scanning electron microscope (SEM) in order to assess the level of workpiece damage. 2.2. Experimental procedure and test parameters The effect of cutting speed when through hole drilling (12 mm deep) with a single peck cycle was investigated. Tests were performed at 30, 45 and 60 m/min (D1, D2 and D3 respectively). Feed rate was kept constant at 0.05 mm/tooth while tool runout was <5 mm with a 45 mm overhang. Maximum drill flank wear (VBmax) was measured on both flutes at appropriate intervals and
averaged, with tests halted either when VBmax reached 0.2 mm or the supply of workpiece material was exhausted. Replicated hole diameter measurements were taken at several depths below the machined surface (1, 3, 5, 7, 9 and 11 mm), with the mean calculated for each depth and hole. Surface roughness (Ra) evaluation was performed along the feed direction at several locations of cross-sectioned holes (via wire EDM), with samples subsequently used for surface integrity analysis (microhardness and microstructure). Table 1 details operating parameters employed for the end milling work. Depth of cut per pass was fixed at 0.25 mm with each plate accommodating a slot depth of up to 4 mm (16 passes). Maximum flank wear was recorded for each cutting edge and averaged; the corresponding tool life criterion was 0.2 mm. Surface roughness profiles both parallel and perpendicular to the feed direction were taken at random locations within the middle two thirds of the slot, with the average for each direction calculated. Subsurface integrity examinations were performed on crosssections cut out from the centre area of selected slots. Table 1 Experimental array for radius end milling trials. Test number
Cutting speed (m/min)
Feed rate (mm/tooth)
M1 M2 M3 M4
25 50 25 50
0.05 0.05 0.10 0.10
3. Results and discussion 3.1. Drilling Tests performed at 30 and 45 m/min exhibited similarly high initial wear rates during the first 40 holes (70 mm average maximum flank wear), after which further progression became gradual and indeed VBmax was only 90 mm after 150 holes. In contrast, the trial with the highest cutting speed resulted in fracture at the outer corner radius of the drill after only 30 holes; see Fig. 2 for wear scar micrographs. In general, continuous, conically shaped helical swarf was produced. [()TD$FIG]
Fig. 2. Wear scar micrographs of drill cutting edges.
Thrust forces (Fz) did not exceed 1600 N for all 3 test conditions when the tool was new. With the onset of tool wear, Fz for Test D1 and D2 increased to approximately 1850 N and 1700 N respectively after 150 holes, while associated torque stabilised to between 5.25 and 5.40 Nm. Both thrust force and torque values when operating at 60 m/min increased exponentially in line with wear evolution and reached 1800 N and 6.75 Nm at the point of tool failure. Analysis of first holes drilled for Test D3 produced an oversize of 40 mm compared to 25 mm for Tests D1 and D2. In general, hole diameter decreased with increasing flank wear although all holes were still above the nominal size at the cessation of each test. Variation in cutting speed appeared to have no discernable effect on hole surface roughness when drills were new, with Ra values between 0.8 and 0.9 mm. These reduced as tests progressed, and in the case of Test D1, was 0.3 mm Ra at the end
S.L. Soo et al. / CIRP Annals - Manufacturing Technology 60 (2011) 89–92
of the trial. All samples exhibited a similar microhardness response, with a hardened layer extending to a depth of up to 100 mm and a maximum of 130 HK0.05 above the bulk value at the machined surface for new drills, see Fig. 3. No significant difference was detected between the feed and radial directions, although the hardness at hole surfaces produced using worn tools were 20 HK0.05 lower. [()TD$FIG]
91
The relatively low subsurface damage observed prompted a further test at 30 m/min to evaluate performance under dry conditions. Results detailed in Fig. 7 showed a white layer covering over 90% of the surface extending to a maximum depth of 15 mm in both radial and feed directions, which agreed with data reported by Kwong et al. [7] when drilling coarse grained RR1000 under equivalent parameters. Further deformation of the microstructure beneath the white layer was also seen, however this only extended 10 mm. Burr size increased considerably at both hole entry and exit, with the latter measuring up to 1 mm in width and 3 mm high. This would suggest that efficient cooling is critical to ensure that temperatures at the cutting zone are kept low in order to produce holes with acceptable surface integrity. [()TD$FIG]
Fig. 7. Surface damage when drilling without cutting fluid. Fig. 3. Microhardness depth profiles with new drills.
SEM analysis showed re-deposited material present on the surfaces of every cross-sectioned hole assessed, with the size of adhered chips varying from a few microns to 100 mm, see Fig. 4. An increase in smeared material was also observed towards the top section of holes however there was no obvious trend between the size/incidence of re-deposited material and flank wear level or operating parameters. Burring occurred at both hole entry and exit although this was relatively moderate as even with worn tools, the maximum burr width/height was <100 mm, see Fig. 5. [()TD$FIG]
Fig. 4. SEM analysis of hole surfaces.
Corresponding micrographs of etched samples highlighted instances of apparent white layers on most surfaces however these were generally less than 2 mm thick, see Fig. 6(a). In addition, distorted layers/surface drag to a depth of up to 4 mm were visible mainly on radial sections and are illustrated in Fig. 6(b). [()TD$FIG]
3.2. Radius end milling All tests resulted in heavy initial flank wear of up to 60 mm within the first two minutes of cutting except for Test M1, which achieved the longest tool life of 120 min. Doubling of cutting speed to 50 m/min (while maintaining feed at 0.05 mm/tooth) reduced machining time to 53 min. Conversely, increasing the feed rate from 0.05 to 0.1 mm/tooth (while keeping cutting speed constant at 25 m/min), only reduced tool life to 71 min. When milling using the highest parameters (Test M4), the tool was worn after only 27 min. However, none of the inserts failed catastrophically, all experienced progressive uniform flank wear over the duration of the experiments. It is estimated that an overall tool life of 8 h is possible when operating at the lowest parameters, assuming 4 usable cutting edges per round insert (rotating 908 after reaching tool life criterion). Surface roughness (Ra) of the machined slots was generally low over the cutting parameters tested, which ranged between 0.2 mm and 0.4 mm when measured perpendicular to the feed and from 0.27 mm to 0.55 mm along the feed direction. Tool condition did not appear to have a significant influence on roughness, particularly at the lowest operating regime. Fig. 8 shows microhardness plots measured parallel to the feed direction of the surfaces machined with new and worn cutting tools. With the former, a strain hardened region up to 50 HK0.05 above the bulk hardness of 460 HK0.05 was recorded. In contrast, profiles for worn tooling show an increase of up to 150 HK0.05, due to the greater deformation caused by the blunter tools. Similar results were seen for samples in the perpendicular direction, although the
[()TD$FIG]
[()TD$FIG]
Fig. 5. Burr formation at hole entry and exit in Test D3 using worn drills.
Fig. 6. Microstructural micrographs of hole radial cross-sections.
Fig. 8. Microhardness depth profiles parallel to feed direction.
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S.L. Soo et al. / CIRP Annals - Manufacturing Technology 60 (2011) 89–92
hardness produced by new tools was 100 HK0.5 higher at the machined surface. All hardened layers extended to a depth of 50 mm from the workpiece surface. No discernable trend could be seen relative to an increase in cutting speed or feed rate, with all profiles showing almost identical maximum hardness levels. Assessment of milled specimens showed burrs visible at both entry and exit locations as well as along the sides of the slot when using worn tools. These varied in height from 25 to 150 mm and width from 16 to 50 mm. In general, higher operating parameters caused greater burring of the workpiece, while use of new inserts prevented such defects from forming on the side walls. SEM micrographs of the machined surfaces revealed limited damage, although small (50 mm) deposits of adhered material and laps were present over the bottom of slots irrespective of the operating parameters used or tool condition, see Fig. 9. [()TD$FIG]
Fig. 9. Micrographs of end milled slot surfaces.
In terms of subsurface integrity, minimal deformation was seen when employing new inserts. Contrary to this, white layers of up to 15 mm deep were observed on sections perpendicular to the feed direction produced in Test M1 using worn tooling, see Fig. 10(a). Traces of re-deposited material up to 5 mm deep were also found in this region. Specimens of sections in the parallel direction and Test M2 showed only limited damage with no visible white layers, see Fig. 10(b). Tests M3 and M4 at higher feed rates produced no
[()TD$FIG]
evidence of white layer or re-deposited material even with worn tools, see Fig. 10(c) and (d). An explanation for the poor surface condition seen in M1 is the increased contact time between the tool and workpiece material due to the mix of low cutting speed and feed rate. 4. Conclusions Tool performance was comparable when drilling RR1000 with low (30 m/min) or intermediate (45 m/min) cutting speeds as both achieved 150 holes without exceeding 100 mm flank wear. All holes inspected showed hardened layers to a depth of 100 mm with limited surface anomalies (adhered chips, material drag, etc.) and moderate burring at entry and exit, irrespective of machining conditions. Results from the dry drilling trial revealed relatively thick white layers present over the hole subsurface and indicated that use of cutting fluid (preferably through spindle) is vital to ensure acceptable hole integrity when drilling this material. Despite excellent tool life when milling at low cutting speed and feed rate, white layers were detected in some of the surfaces assessed. These however were not present when utilising a higher feed rate level (0.1 mm/tooth) in addition to minimal subsurface damage even with worn tools. Slot Ra did not exceed 0.55 mm while only minor flaws in the form of re-deposited chips and laps were observed in the majority of specimens analysed. Compared with Inconel 718, results for RR1000 were similar however the pace of cutting tool developments may have countered machinability performance. Acknowledgments We would like to thank Hardinge-Bridgeport for use of the FGC1000 machine at Birmingham, Sandvik Coromant for the supply of tooling and Kistler Instruments for dynamometry.
References
Fig. 10. Sample cross-sectional SEM micrographs of milled surfaces.
[1] Bradley EF (1988) Superalloys: A Technical Guide. ASM International . (Consulting Ed). [2] Voice W, Hardy M, Rugg D (2005) Powder Prospects ‘Set for Takeoff’ as Production Methods Change. Metal Powder Report 60(4):8–13. [3] Eblinger J, Helm D (2004) Titanium in Aero-Engines, Ti-2003 Sci. Technol.. Proceedings of the 10th World Conference on Titanium, vol. V, July 13–18, Hamburg, Germany, Wiley-VCH. pp. 2845–2852. [4] Rolls-Royce Plc.. (1993) Schematic ETG 37219C );(3). (A3). [5] Reed RC (2006) The Superalloys: Fundamentals and Applications. Cambridge University Press. [6] Furrer D, Fecht H (1999) Ni-based Superalloys for Turbine Discs. Journal of Metals 51:14–17. [7] Kwong J, Axinte DA, Withers PJ (2009) The Sensitivity of Ni-based Superalloy to Hole Making Operations: Influence of Process Parameters on Subsurface Damage and Residual Stress. Journal of Materials Processing Technology 209:3968–3977.
Contents lists available at ScienceDirect
CIRP Annals - Manufacturing Technology jou rnal homep age : ht t p: // ees .e lse vi er. com/ci rp/ def a ult . asp
Machinability and surface integrity of RR1000 nickel based superalloy S.L. Soo a, R. Hood a, D.K. Aspinwall (1)a,*, W.E. Voice b, C. Sage b a b
Machining Research Group, School of Mechanical Engineering, University of Birmingham, Birmingham, UK Manufacturing Technology, Rolls-Royce Plc., UK
A R T I C L E I N F O
A B S T R A C T
Keywords: Nickel alloy Drilling Surface integrity
Alloy development for gas turbine components has produced materials able to maintain strength and integrity at operating temperatures up to 1050 8C. Next-generation RR1000 nickel-based superalloy reflects this philosophy, albeit at the expense of machinability. Experimental data for drilling showed flank wear when operating at 45 m/min to be <100 mm for a distance cut of 1800 mm (150 holes). Thrust forces measured 1600–1800 N. Re-deposited material on hole surfaces and drag/distortion was evident on cross-sectional micrographs. Roughness of end-milled specimens was <0.8 mm Ra with minimal damage using new tools, however significant burring/increased microhardness (150 HK0.05) and white layer formation occurred when employing worn tools. ß 2011 CIRP.
1. Introduction The word ‘superalloys’ defines a group or class of materials which are able to operate over an extended period of time at temperatures above 700 8C, without significant loss of their mechanical and physical properties, particularly strength and corrosion resistance. Typically such alloys employ a high proportion (weight %) of nickel, iron, cobalt and chromium in combination with up to 14 other elements and are regarded as amongst the most highly developed alloys in existence. One classification identifies alloys according to these major constituents such that materials are described as iron-base, cobalt-base and nickel-base, with the latter containing 40–75% Ni. The nickel-base alloys have outstanding oxidation resistance and exceed stainless steels in their high temperature strength [1]. Understandably material and processing costs are significant, irrespective of machining costs which are generally up to 5 that of more conventional alloys due in part to poor tool life, low productivity and the need for acceptable workpiece integrity. The development and growth in superalloys is due principally to the advent of the gas turbine engine in the late 1940s and the subsequent worldwide expansion of the aerospace/aeroengine industry. Application is however, not restricted solely to engines, whether for aircraft or land based power generation systems, encompassing as it does equipment for nuclear reactors, chemical and petrochemical installations, medical/prosthetic devices, rocket engine parts, cryogenic applications, etc. Currently some 50% of aeroengine weight is due to superalloy use, and such materials are particularly prevalent in the combustor and turbine sections where operating temperatures of 1100–1250 8C are typically experienced. It is unlikely that this percentage will increase due to parallel developments in lighter titanium materials with
* Corresponding author. 0007-8506/$ – see front matter ß 2011 CIRP. doi:10.1016/j.cirp.2011.03.094
increased temperature capability (up to 750/800 8C) and greater use of composites. Possible future use of intermetallics such as gTiAl for blades and stators in the HP compressor and blades in the LP turbine section which are currently made from heavier nickel superalloys, are estimated to provide significant weight savings. Despite such developments, the need to provide higher levels of engine fuel economy, reduced NOx emissions and noise, points to increased operating temperatures (higher compressor discharge and turbine entry temperatures) and consequently the use of more advanced nickel alloys, which are able to cope with higher temperatures and stresses than are currently experienced [2]. Fig. 1 details approximate operating temperatures and associated material usage in a typical aeroengine.
[()TD$FIG]
Fig. 1. Aeroengine temperatures and material use. Adapted from [3,4].
From the perspective of publications on the machinability of superalloys, Inconel 718 has over several decades, been the main focus of attention. This is similarly true of the a/b alloy Ti–6Al–4V which has dominated machining publications dealing with titanium alloys. In fairness both have played significant roles in aeroengine development and continue to find application for blades and discs, however over recent years newer nickel
90
S.L. Soo et al. / CIRP Annals - Manufacturing Technology 60 (2011) 89–92
superalloys such as Udimet 720 and RR1000 together with titanium alloys such as Ti–6Al–2Sn–4Zr–6Mo have become increasingly important. Key texts by Bradley [1] and more recently Reed [5] list 100 of the more common wrought (polycrystalline) and cast superalloys, although a number of these have effectively been superseded, on the cast alloy front by improved 2nd, 3rd, and 4th generation compositions. For applications such as turbine discs some 25 current nickel based alloys are cited [5,6]. The wrought alloys encompass solid solution compositions and those capable of precipitation strengthening, while cast materials comprise conventionally cast, direct solidification and single crystal alloys. The increasing emphasis on powder metallurgy processing, followed by forging or alternatively hot isostatic pressing (HIP), in order to improve mechanical properties or achieve greater economy through near net shape manufacture, is an additional consideration. The following sections detail results of twist drilling and radius end milling of powder HIP-ed RR1000. As powder processed forged material, it is finding increased use for turbine discs, but is thought more difficult to machine than mainstream products and hence the new benchmark in relation to superalloy machinability. 2. Experimental work 2.1. Workpiece material format, tooling and equipment Workpiece material was initially supplied in the form of a bar with a bulk hardness of 450 HV30. Several discs measuring 90 mm diameter 12 mm thick as well as strips with dimensions of 100 mm 18 mm 12 mm (used for subsequent hole quality and integrity analysis) were cut from the bar for drilling trials while similarly, plates of 100 mm 32 mm 6 mm were produced for radius end milling tests. Specimens were held in suitably designed clamping fixtures and mounted on appropriate piezoelectric dynamometers for each process. Experiments were carried out on a FGC1000 flexible grinding centre with a maximum spindle rotational speed of 6000 rpm and power rating of 25 kW. All tests were performed wet using a water based mineral oil emulsion; Hocut 3380 at 7–10% concentration. When drilling an adaptor was retrofitted on the main spindle to enable through tool fluid delivery at a pressure of 50 bar and flow rate of 6 l/min. When end milling cutting fluid was supplied externally using a twin nozzle arrangement at 70 bar and 26 l/min. Drills were 8 mm diameter, twin fluted, TiAlN coated solid carbide Delta C ISO-S tools, while the end milling cutter employed was a 20 mm diameter indexable Coromill 300 unit employing 10 mm round, Al(TiN) coated carbide inserts. Tool wear was measured relative to the datum junction of the primary/secondary clearance (Fig. 2) using a microscope and table equipped with digital micrometers giving a resolution of 0.001 mm. Hole diameter was evaluated using a Talyrond 300, while surface roughness was measured with a Form Talysurf 120 L using the appropriate metrology standards. Microhardness measurements of the workpiece (depth profile), were taken using a suitable hardness testing machine at a load of 50 g and 15 s indent duration. For workpiece surface integrity evaluation, machined workpiece surface samples were hot mounted in edge retentive bakelite followed by grinding, polishing and etching according to standard procedures for nickel based superalloy materials. Selected machined workpiece surfaces and cross-sections were analysed using a scanning electron microscope (SEM) in order to assess the level of workpiece damage. 2.2. Experimental procedure and test parameters The effect of cutting speed when through hole drilling (12 mm deep) with a single peck cycle was investigated. Tests were performed at 30, 45 and 60 m/min (D1, D2 and D3 respectively). Feed rate was kept constant at 0.05 mm/tooth while tool runout was <5 mm with a 45 mm overhang. Maximum drill flank wear (VBmax) was measured on both flutes at appropriate intervals and
averaged, with tests halted either when VBmax reached 0.2 mm or the supply of workpiece material was exhausted. Replicated hole diameter measurements were taken at several depths below the machined surface (1, 3, 5, 7, 9 and 11 mm), with the mean calculated for each depth and hole. Surface roughness (Ra) evaluation was performed along the feed direction at several locations of cross-sectioned holes (via wire EDM), with samples subsequently used for surface integrity analysis (microhardness and microstructure). Table 1 details operating parameters employed for the end milling work. Depth of cut per pass was fixed at 0.25 mm with each plate accommodating a slot depth of up to 4 mm (16 passes). Maximum flank wear was recorded for each cutting edge and averaged; the corresponding tool life criterion was 0.2 mm. Surface roughness profiles both parallel and perpendicular to the feed direction were taken at random locations within the middle two thirds of the slot, with the average for each direction calculated. Subsurface integrity examinations were performed on crosssections cut out from the centre area of selected slots. Table 1 Experimental array for radius end milling trials. Test number
Cutting speed (m/min)
Feed rate (mm/tooth)
M1 M2 M3 M4
25 50 25 50
0.05 0.05 0.10 0.10
3. Results and discussion 3.1. Drilling Tests performed at 30 and 45 m/min exhibited similarly high initial wear rates during the first 40 holes (70 mm average maximum flank wear), after which further progression became gradual and indeed VBmax was only 90 mm after 150 holes. In contrast, the trial with the highest cutting speed resulted in fracture at the outer corner radius of the drill after only 30 holes; see Fig. 2 for wear scar micrographs. In general, continuous, conically shaped helical swarf was produced. [()TD$FIG]
Fig. 2. Wear scar micrographs of drill cutting edges.
Thrust forces (Fz) did not exceed 1600 N for all 3 test conditions when the tool was new. With the onset of tool wear, Fz for Test D1 and D2 increased to approximately 1850 N and 1700 N respectively after 150 holes, while associated torque stabilised to between 5.25 and 5.40 Nm. Both thrust force and torque values when operating at 60 m/min increased exponentially in line with wear evolution and reached 1800 N and 6.75 Nm at the point of tool failure. Analysis of first holes drilled for Test D3 produced an oversize of 40 mm compared to 25 mm for Tests D1 and D2. In general, hole diameter decreased with increasing flank wear although all holes were still above the nominal size at the cessation of each test. Variation in cutting speed appeared to have no discernable effect on hole surface roughness when drills were new, with Ra values between 0.8 and 0.9 mm. These reduced as tests progressed, and in the case of Test D1, was 0.3 mm Ra at the end
S.L. Soo et al. / CIRP Annals - Manufacturing Technology 60 (2011) 89–92
of the trial. All samples exhibited a similar microhardness response, with a hardened layer extending to a depth of up to 100 mm and a maximum of 130 HK0.05 above the bulk value at the machined surface for new drills, see Fig. 3. No significant difference was detected between the feed and radial directions, although the hardness at hole surfaces produced using worn tools were 20 HK0.05 lower. [()TD$FIG]
91
The relatively low subsurface damage observed prompted a further test at 30 m/min to evaluate performance under dry conditions. Results detailed in Fig. 7 showed a white layer covering over 90% of the surface extending to a maximum depth of 15 mm in both radial and feed directions, which agreed with data reported by Kwong et al. [7] when drilling coarse grained RR1000 under equivalent parameters. Further deformation of the microstructure beneath the white layer was also seen, however this only extended 10 mm. Burr size increased considerably at both hole entry and exit, with the latter measuring up to 1 mm in width and 3 mm high. This would suggest that efficient cooling is critical to ensure that temperatures at the cutting zone are kept low in order to produce holes with acceptable surface integrity. [()TD$FIG]
Fig. 7. Surface damage when drilling without cutting fluid. Fig. 3. Microhardness depth profiles with new drills.
SEM analysis showed re-deposited material present on the surfaces of every cross-sectioned hole assessed, with the size of adhered chips varying from a few microns to 100 mm, see Fig. 4. An increase in smeared material was also observed towards the top section of holes however there was no obvious trend between the size/incidence of re-deposited material and flank wear level or operating parameters. Burring occurred at both hole entry and exit although this was relatively moderate as even with worn tools, the maximum burr width/height was <100 mm, see Fig. 5. [()TD$FIG]
Fig. 4. SEM analysis of hole surfaces.
Corresponding micrographs of etched samples highlighted instances of apparent white layers on most surfaces however these were generally less than 2 mm thick, see Fig. 6(a). In addition, distorted layers/surface drag to a depth of up to 4 mm were visible mainly on radial sections and are illustrated in Fig. 6(b). [()TD$FIG]
3.2. Radius end milling All tests resulted in heavy initial flank wear of up to 60 mm within the first two minutes of cutting except for Test M1, which achieved the longest tool life of 120 min. Doubling of cutting speed to 50 m/min (while maintaining feed at 0.05 mm/tooth) reduced machining time to 53 min. Conversely, increasing the feed rate from 0.05 to 0.1 mm/tooth (while keeping cutting speed constant at 25 m/min), only reduced tool life to 71 min. When milling using the highest parameters (Test M4), the tool was worn after only 27 min. However, none of the inserts failed catastrophically, all experienced progressive uniform flank wear over the duration of the experiments. It is estimated that an overall tool life of 8 h is possible when operating at the lowest parameters, assuming 4 usable cutting edges per round insert (rotating 908 after reaching tool life criterion). Surface roughness (Ra) of the machined slots was generally low over the cutting parameters tested, which ranged between 0.2 mm and 0.4 mm when measured perpendicular to the feed and from 0.27 mm to 0.55 mm along the feed direction. Tool condition did not appear to have a significant influence on roughness, particularly at the lowest operating regime. Fig. 8 shows microhardness plots measured parallel to the feed direction of the surfaces machined with new and worn cutting tools. With the former, a strain hardened region up to 50 HK0.05 above the bulk hardness of 460 HK0.05 was recorded. In contrast, profiles for worn tooling show an increase of up to 150 HK0.05, due to the greater deformation caused by the blunter tools. Similar results were seen for samples in the perpendicular direction, although the
[()TD$FIG]
[()TD$FIG]
Fig. 5. Burr formation at hole entry and exit in Test D3 using worn drills.
Fig. 6. Microstructural micrographs of hole radial cross-sections.
Fig. 8. Microhardness depth profiles parallel to feed direction.
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S.L. Soo et al. / CIRP Annals - Manufacturing Technology 60 (2011) 89–92
hardness produced by new tools was 100 HK0.5 higher at the machined surface. All hardened layers extended to a depth of 50 mm from the workpiece surface. No discernable trend could be seen relative to an increase in cutting speed or feed rate, with all profiles showing almost identical maximum hardness levels. Assessment of milled specimens showed burrs visible at both entry and exit locations as well as along the sides of the slot when using worn tools. These varied in height from 25 to 150 mm and width from 16 to 50 mm. In general, higher operating parameters caused greater burring of the workpiece, while use of new inserts prevented such defects from forming on the side walls. SEM micrographs of the machined surfaces revealed limited damage, although small (50 mm) deposits of adhered material and laps were present over the bottom of slots irrespective of the operating parameters used or tool condition, see Fig. 9. [()TD$FIG]
Fig. 9. Micrographs of end milled slot surfaces.
In terms of subsurface integrity, minimal deformation was seen when employing new inserts. Contrary to this, white layers of up to 15 mm deep were observed on sections perpendicular to the feed direction produced in Test M1 using worn tooling, see Fig. 10(a). Traces of re-deposited material up to 5 mm deep were also found in this region. Specimens of sections in the parallel direction and Test M2 showed only limited damage with no visible white layers, see Fig. 10(b). Tests M3 and M4 at higher feed rates produced no
[()TD$FIG]
evidence of white layer or re-deposited material even with worn tools, see Fig. 10(c) and (d). An explanation for the poor surface condition seen in M1 is the increased contact time between the tool and workpiece material due to the mix of low cutting speed and feed rate. 4. Conclusions Tool performance was comparable when drilling RR1000 with low (30 m/min) or intermediate (45 m/min) cutting speeds as both achieved 150 holes without exceeding 100 mm flank wear. All holes inspected showed hardened layers to a depth of 100 mm with limited surface anomalies (adhered chips, material drag, etc.) and moderate burring at entry and exit, irrespective of machining conditions. Results from the dry drilling trial revealed relatively thick white layers present over the hole subsurface and indicated that use of cutting fluid (preferably through spindle) is vital to ensure acceptable hole integrity when drilling this material. Despite excellent tool life when milling at low cutting speed and feed rate, white layers were detected in some of the surfaces assessed. These however were not present when utilising a higher feed rate level (0.1 mm/tooth) in addition to minimal subsurface damage even with worn tools. Slot Ra did not exceed 0.55 mm while only minor flaws in the form of re-deposited chips and laps were observed in the majority of specimens analysed. Compared with Inconel 718, results for RR1000 were similar however the pace of cutting tool developments may have countered machinability performance. Acknowledgments We would like to thank Hardinge-Bridgeport for use of the FGC1000 machine at Birmingham, Sandvik Coromant for the supply of tooling and Kistler Instruments for dynamometry.
References
Fig. 10. Sample cross-sectional SEM micrographs of milled surfaces.
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