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Minor cutting edge–workpiece interactions in drilling of an advanced nickel-based superalloy
ARTICLE IN PRESS International Journal of Machine Tools & Manufacture 49 (2009) 645–658
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Minor cutting edge–workpiece interactions in drilling of an advanced nickel-based superalloy J. Kwong a, D.A. Axinte a,, P.J. Withers b, M.C. Hardy c a b c
School of Mechanical, Materials and Manufacturing Engineering, The University of Nottingham, Nottingham NG7 2RD, UK School of Materials, The University of Manchester, Grosvenor Street, M1 7HS, UK Rolls-Royce plc, P.O. Box 31, Derby DE24 8BJ, UK
a r t i c l e in fo
abstract
Article history: Received 4 October 2008 Received in revised form 14 January 2009 Accepted 16 January 2009 Available online 7 February 2009
Drilling is one of the key machining operations for manufacturing safety critical components that must comply with strict surface quality standards. The influence of major flank wear of drilling tools on workpiece surface quality has been well established; however, similar information concerning minor cutting edge is currently missing from literature. This paper presents a comprehensive analysis and discussions of the influence of the drill’s minor cutting edge to workpiece surface integrity and residual stress distribution for RR1000, a newly developed nickel-based superalloy. These effects are critical to the acceptance of this new material in relation to tool geometry and machining strategies. The thickness of material drag in the hoop direction has been found to be the highest at the top and the least at the bottom of the hole, which is directly related to the contact duration between the minor cutting edge and workpiece surfaces; moreover this difference increased at higher levels of wear of the minor cutting edge. On-line process monitoring techniques have been employed to further understand the material drag phenomena, including feed force, torque and acoustic emission. Compressive axial and tensile hoop stresses at the surface of the holes have been measured as a function of depth and correlated both with metallurgical analysis of drilled surfaces and the process monitoring signals. It was found that the increased material drag associated with a worn tool resulted in compressive hoop surface residual stresses near the entrance hole in correspondence with trends in the processed acoustic emission signal. This work suggests that material drag increases with the duration of the minor cutting edge–workpiece interaction such that plastic deformation is the greatest near the drill entrance holes and that process monitoring of the degree of material drag in hoop direction can be practicable. & 2009 Elsevier Ltd. All rights reserved.
Keywords: Nickel-based superalloy Drilling Material drag Residual stresses Acoustic emission
1. Introduction One of the challenges facing the aerospace manufacturing industry is to introduce increasingly high-strength superalloys, while continuously adapting manufacturing processes to enable the generation of high-integrity components at lower manufacturing costs, higher production rates and lower scrap rates. Safety critical components, such as turbine discs, are machined under strict quality standards set by aerospace manufacturers; in this respect surface integrity issues are of critical importance, especially when new superalloys are adopted. Due to their low thermal conductivity, their tendency for strain hardening and high strength at elevated temperatures, nickel-based superalloys, such as Inconel, Waspaloy and Udimet 720Li, are notoriously difficultto-cut materials [1,2]. Indeed, achieving the high degree of workpiece surface integrity required for the safe utilization of
Corresponding author.
E-mail address: [email protected] (D.A. Axinte). 0890-6955/$ - see front matter & 2009 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijmachtools.2009.01.012
aeroengine parts made of superalloys subjected to high thermal and mechanical loading is a demanding task and all the outcomes of the employed machining processes must be well understood [3]. A variety of surface damages (i.e. anomalies) such as material drag, cracking and plucking [4] can result when nickel-based superalloys are machined with inappropriate parameters or tool conditions. These anomalies can have a detrimental effect on the fatigue performance of safety critical components during their operational life [5]. Safety critical (rotating) parts (e.g. discs) contain a variety of geometrical features that, according to their anticipated loading distributions, have various degrees of importance when it comes to influencing the fatigue performance of the manufactured component. Turning, milling, broaching, grinding and hole making are all involved in the manufacture of such aerospace components. While turning [6], milling [7] and grinding [8] of nickel-based alloys have been well documented in the academic literature, as well as industrial publications, holemaking operations have not received in-depth analysis with regard to workpiece surface quality in terms of metallurgical/ geometrical integrity, residual stresses and their implications.
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This is perhaps a surprising fact given that hole features have the highest degree of criticality for the safe exploitation of rotating aerospace components loaded in fatigue conditions [9]. Wear of the major flank of the drills is the dominant mechanism limiting tool life with worn drills raising the prospect of introducing damage into the workpiece [10,11] as well as additional costs associated with the necessary re-manufacture of the out-of-specification machined components. Consequently, it is important to better understand the relationship between drill wear and quality of machined surfaces. Some information on tool wear and chip formation in drilling of nickel-based superalloys has been reported [12] but surprisingly little reference has been made to the more important aspect of the damage induced in the machined surfaces. A limited published research has been focussed only on the metallurgical inspection in the axial direction for various hole-making techniques, including drilling, reaming and plunge/spiral milling [11,13]; this is a very limitative inspection that gives information on the surface integrity of the machined holes only along feed direction, which might not be (as this report shows in the following) the ‘‘preferential’’ direction on which surface damage occurs in hole-making operations. As the main cutting edges of the drill follow a helical trajectory, analysis in the hoop (cutting) direction is needed to facilitate a more complete interpretation of the phenomena occurring at the tool–workpiece interface during drilling and thus to enable better understanding of their implications for surface integrity, including metallurgical (e.g. phase transformations) and mechanical (e.g. residual stress) aspects. Typically drilling tools (Fig. 1a) have a cylindrical geometry [14] (i.e. not back tapered), which has raised concerns within the aerospace machining community on issues
related to the influence of continuous interactions between the minor cutting edge (S00 ) and the walls of the workpiece on the integrity of final machined surfaces. Previous research has examined the wear of minor cutting edges in drilling [15], but to the authors’ knowledge, its effect on workpiece surface integrity has not been reported. This aspect could be of crucial importance on the hole-making methods when it comes to generating damage-free surfaces like those required in aerospace industry. In view of the importance of tool wear in drilling, on-line wear monitoring systems to help decide when to schedule tool replacement are currently in development as part of the move towards fully automated manufacturing cells. Tool condition monitoring systems allow the prediction of conditions that can lead to catastrophic damage modes. Traditionally, tool wear has been measured optically, but this requires off-line inspection. To avoid this, on-line process monitoring techniques have been employed to determine tool wear levels from readings of the cutting forces, torque, spindle power, vibrations and acoustic emission (AE) signals. As with other conventional machining operations, in drilling, forces increase as tool wear develops [16,17]. Often only the feed force is considered, but recent research has advocated condition monitoring using simultaneous consideration of feed force along with cutting torque [18]. While a direct representation of the cutting forces in drilling can be based on the use of torque readings, the linear increase of torque signal, over the period that the minor cutting edges of the drill come into contact with the workpiece until full completion of the hole, remains unexplained. Spindle current/power signals have been shown to be a reliable way of detecting tool breakage, but
Fig. 1. (a) Basic geometrical specifications of a two flutes twist drill geometry (ISO 2003/1). (b) Illustration of the various stages/cutting times (t1–t5) of the drill minor cutting edges (S0 )–workpiece interaction: t1—corners of drill reaches workpiece; t2—S0 –workpiece interaction begins; t3—material drag increases cumulatively; t4)—corners of drill exits workpiece; t5—material drag continues until tool retraction.
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insufficient to provide a robust means of determining lesser increases in tool wear [19,20]. It is well known that the main contributors of AE in machining come from friction at the clearance faces (for both major and minor cutting edges), plastic deformations in the shear zone and chip breakage [21]. Research has shown that the energy count (Measured Area of the Rectified Signal Envelope—MARSE) of the AE signal increases with tool wear; whereas envelope analysis is good for detecting tool wear only at transient stages of the drilling process [22]. Analytical techniques for prediction of surface quality from machining have been previously employed, including frequency analysis [23] and interpretation of RMS signal [24]. The use of vibration signals has been employed, but since these contain both ‘‘useful’’ and noise signals, the latter can have a significant effect on the robustness of such techniques [25]. In summary, all the current process monitoring techniques focus mainly on detecting excessive flank wear and tool breakages in the drilling operation. By contrast, information relating process signals to the occurrence of damage to workpiece surface integrity is very limited in machining, which includes the defection of plucking [26], white layers [27] and material drag [28] from AE analysis. In particular, information is missing concerning minor cutting edge–workpiece interactions and their implications for nearsurface damage modes, machining force/torque signals and AE signals. In general, machining may result in a high degree of plastic deformation and strain on the workpiece surface when the cutting parameters are not optimised from a surface quality viewpoint [29]. Non-uniform plastic deformation will generally give rise to locked-in residual stresses, comprising balancing compressive and tensile stresses [30,31]. In machining, the nature of the residual stress is determined by two factors, namely mechanical and thermal effects. Mechanically induced residual stress arises from plasticity deformed region near the tool–workpiece interface; the material ahead of the cutting edge is believed to be in compression whilst material behind will be in tension. As a consequence, if plasticity ahead of the cutting edge is dominant then the residual stresses caused by mechanical effects would be tensile; however if plasticity behind the tool is more influential then this favors compressive stresses [32]. Locally significant temperatures can be generated as a combination of plastic deformation of the workpiece and friction at the tool–workpiece interface. The sharp thermal gradients generated in such a way can contribute to tensile stresses in the superficial layer of the workpiece [33]. Tensile residual stresses in the surface and subsurface regions are known to significantly impair fatigue resistance of critical components during operation of critical gas turbine components whereas compressive will prolong the operating life [34]. In literature, reports on the measurements of residual stress associated with drilling are scarce. Furthermore, the reports tend to focus only on stresses in the axial direction and measurements in the hoop direction have not been considered primarily due to the difficulties of making X-ray measurements on a tightly curved surface [35]. Moreover, mapping of the residual stress profiles along the height and beneath the free surfaces of drilled holes have received little attention, preventing the establishment of an understanding of the process by which residual stresses build up during the drilling operation.
2. Scope of research This paper is the first to focus attention on the influences of the interaction between the drill’s minor cutting edge (S0 ) and workpiece surface that can affect the integrity of machined surfaces. This is discussed in relation to drilling of a new nickel-
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based superalloy, i.e. RR1000, which is regarded as a nextgeneration aerospace material for the manufacture of safety critical parts. RR1000 possesses superior properties, including 25–50 1C higher operating temperature capability over the current disc alloy, 720Li, whilst having an equivalent crack growth resistance to coarse grain Waspaloy [36]. These properties make RR1000 a more difficult-to-cut material than the existing nickelbased alloys (e.g. Inconel, Udimet 720), making it more susceptible to the generation of surface anomalies caused by minor cutting edge during drilling operation. Consequently, this paper addresses the following aspects: (a) investigation of the S0 –workpiece interactions and their effect on the surface integrity; (b) in-depth analysis and comparison of metallographic, 2D residual stress distributions and plastic deformations in the superficial layer of the workpiece to enable quantitative and qualitative evaluations of the outcomes of S0 –workpiece interactions; (c) assessment of potential on-line process monitoring techniques, including feed force, torque and acoustic emission signals as a means of determining the magnitude of S0 –workpiece interaction phenomena to propose process supervision methods that enable machining of damage-free surfaces.
3. Analysis of minor cutting edge–workpiece interactions The mechanism of the minor cutting edge–workpiece interaction can be broken down and discussed in different stages (t1–t5) as presented in Fig. 1b and is discussed in the following: (a) In the first stage of drilling, tip of the drill comes into contact with the workpiece surface and the contact area with the major cutting edges starts to increase. Once the major cutting edges have fully immerged into the workpiece at time, t1, the minor cutting edges of the drill begin to interact with the workpiece surface (at the hole depth), d1. Until, t1, hoop plastic deformation (material drag) caused solely by the minor cutting edges (S0 ) cannot take place. (b) At time t2, the drill progresses deeper into the workpiece and the total contact area between the S0 and workpiece surface increases proportionally with cutting time. (c) During this stage, the S0 –workpiece interaction can cause material from the surface to plastically deform in the hoop (cutting) direction by a dragging mechanism. The extent of the interaction is dependent on the time that the minor cutting edges of the drill are in contact with the workpiece material (walls of the hole) and the intensity of the associated friction phenomena (‘‘embedded’’ within cutting torque). Consequently the material near the entrance of the hole is exposed to interaction with the minor cutting edges of the drill for longer time than that near the exit, allowing for a larger cumulative dragging effect. Thus, a ‘‘triangular plastic deformation’’ (TPD) distribution might be expected to describe the volume of material drag caused by S0 –workpiece interaction at any time during the drilling operation. Of course, the extent of shearing within this zone may be limited if it exceeds a critical value, leading to local material fractures. (d) On increasing the drilling depth to d3 at time t3, the TPD zone extends in both axial and hoop directions at a rate proportional to the total cutting time, provided contact between the S0 and the workpiece is maintained along the depth of the drilled hole. (e) When the major cutting edges of the drill exit from the workpiece at time t4, the complete depth of the hole has
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experienced the minor cutting edge–workpiece interaction and the TPD zone would be expected to extend through the whole depth of the hole. (f) During the push-through period, at time t5 the only contact between the drill and the workpiece is with the minor cutting edges as the major cutting edges have fully emerged from bottom of the hole. At this time, material dragging and heat generation continues and can occur until the tool is retracted, causing the generation of plastic deformation to persist even after the corners of the drill exit the hole at depth d4. This analysis suggests that the longer the S0 –workpiece interaction, the greater the material drag in the hoop (cutting) cutting direction; this is provided that continuous contact between the minor cutting edges and wall of the drilled is maintained behind the tip, a situation that is likely to happen when employing cylindrical drills in generating relatively short holes (h/dE1). Therefore, in the light of these considerations, regions near the start (top) of the hole would be expected to experience a higher material drag than those at the end, in this way leading to the creation of a TPD zone. Additionally, with the increased flank wear of the major cutting edges it is expected that more stock of material is left for removal by minor cutting edges, which is likely to result in more material dragged in the hoop direction; therefore, in such instances, it is expected that TPD zone will expand deeper into drilled surfaces, a situation that will be proven later in the paper. TPD zone that is the result of minor cutting edges–workpiece interaction is experimentally investigated in this paper for new, intermediate and worn drilling tools. In addition, the extent to which both the mechanical effect and the heating effect may give rise to residual stresses that vary along the depth of the hole has also been examined. If a build-up of material drag due to the S0 –workpiece interaction is identified, appropriate measures to reduce this effect could be developed (i.e. on-line monitoring
solutions) to enable high workpiece surface integrity components for sensitive applications (i.e. safety critical component) such as those in aerospace industry.
4. Experimental set-up and methodology In the following, testing programme drills of new, intermediate and worn cutting edges have been employed for hole making on the proprietary nickel-based CG RR1000 superalloy manufactured via the powder metallurgy route [36], which display the following chemical composition: 15% Cr, 18.5% Co, 5% Mo, 3% Al, 3.6%Ti, 2% Ta, 0.5% Hf, 0.06% Zr, 0.027% C, 0.015% B, Ni balance. The high strength at elevated temperatures of CG RR1000 derives from a combination of solid solution strengthening and g0 precipitates [37], which makes it difficult to cut with a high tendency to retain tensile residual stresses after machining. Cylindrical plates of CG RR1000 were ground on both sides to ensure the workpiece to be flat when placed in a special rig for drilling. Prior to any machining, hardness and residual stress measurements have been performed on the workpiece to ensure that the sample has not been work hardened or does not have pre-machining induced residual stresses. A five-axis Makino A55 machine equipped with Corogrip Precision Power Chuck displaying reduced tool run-out (o2 mm) has been employed throughout the experimental trials. A waterbased synthetic cutting fluid (Houghton 3380) was supplied through tool for selected drilling trials. A circular array of holes was drilled; this enabled concentric rings to be cut by electrodischarge machining (EDM) for subsequent workpiece surface inspections. A process monitoring set-up (Fig. 2) consisting of a Kistler four-axis dynamometer (9272) acquiring at 10 kHz torque and x, y and z force signals was mounted behind the fixturing system. An AE sensor (Kistler 815B121) was thread mounted within the fixturing system. The AE signals were acquired at a
Fig. 2. Schematic illustration of drilling set-up (a) and fixturing system on Makino A55 (b).
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frequency of 800 kHz to enable spectral analysis of the signal within the response frequency of the sensor (50–400 kHz). The acquired signals were then transferred via two synchronised data acquisition cards (NI PCI-MIO-16E-dynamometer signals, NI-6115AE sensor) and saved/processed using an in-house LabView programme. To develop an appreciation of the operating window on which the material can be satisfactory drilled, a wide range of drilling parameters has been tested. To address the scope of the paper, only those trials needed to exemplify the significance of the S0 –workpiece interactions on surface integrity are reported here. Factors such as tool wear (VBmax ¼ 0–0.3 mm), cutting parameters (f ¼ 0.08–0.1 mm/rev, n ¼ 1061–1592 RPM) and coolant conditions (dry or through tool coolant supply at 30 bar) were varied. Except where indicated, the discussions on the S0 –workpiece interactions will predominantly refer to three cutting conditions employing new, intermediate (VB ¼ 0.1 mm) and worn (VB ¼ 0.3 mm) tool used at a constant feed rate, f ¼ 0.10 mm/rev and variable spindle speed, n ¼ 1061–1592 RPM whilst keeping coolant on. The machined holes were initially sectioned along the axial direction via cut-off wheels to obtain two halves; one used for residual stress measurements and the other for metallurgical analysis. To perform scanning electron microscope (SEM) inspections, it was necessary to section the hole in the hoop direction at 1 mm intervals along depth of the holes as illustrated in Fig. 3. The specimens were then mounted in conductive bakelite for in-depth micro-scale metallurgical analysis using an environmental field emission gun SEM (Philips XL30 ESEM-FEG). The specimens were ground using silicon carbide paper through to 2400 grit and polished to a 1 mm finish on DUR (6 mm) and NAP (1 mm) cloths. Microstructure was revealed using electrolytic etching with 10% ortho-phosphoric acid in distilled water using a current of 2–3 A with stainless-steel wires as anode and cathode.
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Residual stress measurements were performed on a Proto X-ray diffractometer using Mn Ka radiation (l ¼ 2.1031 A˚) to acquire the (3 11) diffraction peak at a 2y angle of approximately 1521 using a spot size of 1 mm diameter. Prior to the measurements of the samples, a reference (flat and stress free sample of RR1000) was measured. All samples were sectioned to ensure no shadowing effect or stress relaxation would occur during the measurements, which would otherwise affect the accuracy of the results. This was confirmed with the comparison against the defined gain. Sin2 C technique measurements were performed at multiple C tilt angles (b angles with the Proto iXRD) with a 31 oscillation over a C range of 7271. The penetration depth of Mn Ka radiation in Ni is approximately 22 mm considering 99% absorption by the workpiece at C ¼ 01 with negligible changes in the range of tilts performed in the full measurement. The Young’s modulus and Poisson’s ratio used for the {3 11} reflection were 204 GPa and 0.27, respectively.
5. Experimental results and discussion 5.1. Characterisation of material drag in drilling In Fig. 4, preliminary metallographic examinations in the axial direction after drilling operation when either new or worn tool are employed do not reveal any surface anomalies. The clean surface generated from a worn tool might be a surprising observation as typically worn cutting edges have a high tendency to generate surface damages (i.e. material drag, cracks). However, as presented in Fig. 1b, the main drill cutting edges follow helical paths and therefore the generation of surface anomalies may be more dominant in the hoop direction whilst leaving the surface in axial direction much less damaged. To provide a better understanding, metallurgical analysis in hoop direction must be performed and
Fig. 3. Preparation of specimen for SEM analysis in axial and hoop directions.
Fig. 4. Secondary electron images ( 5000) of the workpiece surface viewed in the axial direction (VB ¼ 0.0 mm, f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on): (a) new tool and (b) worn tool.
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more importantly to evaluate the influence of S0 (minor cutting edge) to the surface finish from drilling operation. To explore the severity of material drag caused by S0 (minor cutting edge)–workpiece interaction, in Fig. 5 an axially sectioned hole with an embedded broken tool (VB ¼ 0.3 mm, f ¼ 0.08 mm/rev,
Fig. 5. Axial sectioned optical image of blind hole illustrating the S0 –workpiece interaction.
n ¼ 1326 RPM, coolant on) is presented; this gives an indication of the possible local ‘‘welds’’ between S0 and the wall of the drilled hole. The circled regions (Fig. 5) illustrate that S0 retains intimate contact with the workpiece well behind the tip. Indeed, in extreme cases such as these, local S0 –workpiece welding can contribute (along with high cutting loads at the main cutting edges) to drill breakage. Moreover, for the S0 and workpiece to locally weld, a high local (minor cutting edge) temperature, likely caused by the intensive friction phenomena, must be present, which also facilitates plastic deformation on the surfaces in the hoop direction. Metallographic examination of a hole produced using a new drill (VB ¼ 0 mm) is illustrated in Fig. 6. Preliminary investigations of metallographic analysis in the axial (feed) direction show negligible amount of material drag beneath drilled surfaces; hence the follow-up examinations have been focussed on the hoop direction. The thickness (measured in the radial direction of the hole) of the plastic deformation zone was quantified as the distance from the free surface of the workpiece to the depth beneath the surface where no more plastic deformation is visible on the micrographs; lower magnifications ( 5000) enable the observation of the depth of material drag while higher magnification pictures ( 10,000) put in evidence the angle of the dragged grains. From the SEM images, it is clear that the depth of material drag tends to decrease from the top (Figs. 6a and b) to the bottom (Figs. 6e and f) of the holes. Material drag observed on the surfaces is caused by the ‘‘ploughing’’ interaction between the S0 and walls of
Fig. 6. Secondary electron images ( 5000 on left column and 10,000 on right column) of the workpiece surface viewed in the hoop direction at various depths: (a, b) 1 mm, (c, d) 3 mm and (e, f) 5 mm when using a new drill (VB ¼ 0.0 mm, f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on).
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the hole, which follows the direction of the cutting action. As the contact time between the S0 and workpiece is lower near the exit of the hole, a lesser degree of material drag has been observed (Fig. 6e). As previously explained, from the point of view of the material drag caused by S0 the plastic zone might be expected to have ‘‘zero’’ depth at the bottom of the hole in the hoop direction. Although the major cutting edges do not directly come into contract with the walls of the workpiece, a somewhat pushedthrough effect induces a small degree of material drag in the hoop direction (Fig. 6e). Unsurprisingly, the extent of material drag was found to be more significant when employing worn tools as illustrated in Fig. 7. Material drag to a depth of 32 mm (cf. 20 mm for the new tool) was measured 1 mm from the top of the hole, decreasing to 2 mm at 5 mm depth. This is almost certainly related to the increased tool wear on the minor cutting edges of the drill (Figs. 8a and b) since this is the main contact between the drill and walls of the hole. However, it should also noted that the corners of the drill are most susceptible to wear during cutting (Figs. 8c and d); this is likely to lead to an initially smaller hole diameter and thus, leaving more stock of material for removal by the S0 . Since the material stock removed by the S0 is much lesser than those from the major cutting edges, it is likely that S0 will have a tendency to drag the surface material rather than cut it. Hence, more intense material drag is expected when worn tools are employed. Therefore, the wear of the S0 is also likely to add to this material dragging phenomenon. However, to separate the effects of wear from the major and minor cutting edges on
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material drag is difficult to be evaluated unequivocally and this is an issue under current investigation. The intensity of material drag was found to be more significant at the top of the hole. As previously commented in Section 3 (Fig. 1b), the top of the hole experiences the longest duration of S0 –workpiece interaction, hence causing the maximum material drag in the TPD zone. This effect is evidenced in Figs. 7a and b with the surface layer smeared to such an extreme extent that cracks are visible to a radial depth of around 4 mm. In Fig. 9, the depth of the plastically dragged layer is plotted as a function of depth (1–5 mm) of the hole when employing new and worn drills; values of these measurements are presented in Table 1. From this graph it is clear that the plastically deformed zone is approximately triangular in form and is approximately 50% deeper for the worn tool. Therefore, the initial assumption on the generation of a TPD zone is caused by S0 –workpiece interaction, which has a good correlation with the results of the metallurgical/surface integrity investigations. 5.2. Evolution of feed force and torque To support the understanding of material drag, Fig. 10 shows the torque and feed force signals when drilling CG RR1000 for both new and worn drills with the same cutting conditions. The distinct drilling stages seen in relation to the pattern of sensory signals (Fig. 10) are discussed below. A. Tip entrance: From the moment when the drill tip comes into contact with the workpiece until full immersion of the major
Fig. 7. Secondary electron images ( 5000 on left column and 10,000 on right column) of the workpiece surface viewed in the hoop direction at various depths: (a, b) 1 mm, (c, d) 3 mm and (e, f) 5 mm when using a worn drill (VB ¼ 0.3 mm, f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on).
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Fig. 8. Optical images of drills illustrating different degrees of tool wear (a) new tool–minor cutting edge; (b) worn tool–minor cutting edge; (c) new tool–major cutting edge; (d) worn tool–major cutting edge.
Fig. 9. Scatter plot from averaged values (five measurements) of the thickness of material drag against depth of hole.
Table 1 Averaged thickness of material drag obtained using new and worn drilling tools against the depth of hole (results based on five measurements). Depth of hole (mm)
1 2 3 4 5
Material drag thickness (mm) New tool
Worn tool
0.020 0.018 0.012 0.004 0.002
0.032 0.033 0.018 0.012 0.002
cutting edges, a sharp increase of feed force and torque signals is observed. When a worn drill is employed, the increase in feed force substantially exceeds that for a fresh drill. This can be explained by the decrease in cutting ability when the major cutting edges are worn. B. Development of minor cutting edges–workpiece interaction: Once the major cutting edges are fully immersed into the workpiece, feed force (Fz) and torque (T) reach a steady state, at least for the new tool. At this stage, the contact area of S0 with the walls of the hole increases linearly with cutting time, which contributes to the increase of T as the drill progresses through the depth of the hole. This is where the generation of the TPD distribution begins and the development of material drag initiates. It is clear that feed force varies little with increasing drill depth, being larger for the worn tool. On the other hand, torque appears to increase linearly with depth, the gradient being slight for the new tool and much steeper for the worn tool. As the drill progresses deeper into the hole, the extent of S0 –workpiece contact increases proportionally with depth. The observed linear increase of torque thus corroborates the initial consideration that material drag originates from the workpiece interaction with the minor cutting edges and giving rise to the TPD zone. Further, it is not surprising that this linear increase is evident to a much lesser degree for the new tool as the minor cutting edges efficiently shear off material rather than dragging them. In Fig. 10, the dashed line (TIdeal line) represents the ‘‘ideal’’ level of torque, being equal to that generated solely by the major cutting edges of the drill. It should also be noted the feed force starts to decrease just before the tip exits the workpiece for both new and worn drills. At this depth, the remaining material ahead of the drill has been plastically deformed under the feed force rather than being appropriately cut. This effect is observed only during the latter stages in the full immersion zone since the remainder of workpiece material is significantly thinner leading to reduced strength. C. Tip exit: As the tip of the new drill exits the workpiece an instantaneous decrease in both Fz and T is observed. This is in contrast with the response when a worn drill is used. Instead, a sudden drop in both feed force and torque was not measured until a point 0.8 mm beyond the expected exit of the workpiece
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Fig. 10. Torque and feed force signals vs. time (lower axis) and distance travelled (upper axis) for both new (VB ¼ 0.0 mm) and worn (VB ¼ 0.3 mm) drills in the same cutting conditions (f ¼ 0.10 mm/rev, n ¼ 1592 RPM, coolant on) showing different drilling stages: A—tip entrance, B—full immersion, C—tip exit, D—drill push through and E—drill retraction.
Fig. 11. 3D mapping of plastic deformations/burrs formed at exit of drilled holes (f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on): (a) new tool (VB ¼ 0 mm) and (b) worn tool (VB ¼ 0.3 mm).
(at 6 mm). In this case, as the main cutting edges of the drill are worn they cause more material to be plastically deformed beyond the ‘‘theoretical’’ edge of the workpiece. In this context, Fig. 11 shows an increase from around 108 to around 187 mm in the height of the exit burrs when a worn drill is used. The larger burr formed by a worn tool resulted in increased duration of the Fz and T signals, which is due to the increased contact time between the tool and workpiece surface. As the corners of drill wear out, this leads to lesser material being removed by the major cutting edges;
the ‘‘formed’’ hole has a slightly smaller diameter than the nominal one. The remainder of stock material will be removed by the minor cutting edges and hence this results in extended contract duration measured as feed force and torque. For 0.8 mm of travel beyond the nominal exit hole position the torque remains essentially undiminished since S0 edge–workpiece interaction is still present until the geometry of the exit burr has stabilized. Since the torque reaches a maximum level once the major cutting edges have exited the nominal hole, this corroborates the theory
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that S0 –workpiece interaction is present in drilling. Moreover, S0 –workpiece interface is the main contact between the drill and the walls of the workpiece; therefore such interactions are the main contributor to material dragging in the hoop direction and the formation of TPD zone. D. Drill push through: During this stage no more material can be removed by the major cutting edges of the drill and hence no feed force is recorded for either the new or worn drill. However the torque, which is negligible for the new tool, declines significantly (from 5.3 to 2.6 Nm) over this region. During this period the only interaction between the drill and the workpiece occurs via the minor cutting edges and thus one can deduce, in confirmation of the results for stage B, that the torque arising from the S0 –workpiece interaction is significant for the worn drill and considerably smaller for the new drill. At this particular stage of drilling, the burrs at the exit of the hole are pushed out with increasing hole diameter by S0 , and hence increasing the contact duration and hence reducing the area of contact and torque. E. Drill retraction: Once the drill travels the preset distance, it is retracted; this is revealed through a negative feed force, which demonstrates that S0 is still in contact with the workpiece and that the machined hole is still being deformed with material pushed through during the latter stages of full immersion. Process monitoring and signal analysis of the drilling operation from the perspective of both feed force and torque have provided a proof for minor cutting edges–workpiece interaction of during drilling operation while giving information of the patterns of these signals for different levels of cutting edge conditions. Without such process monitoring techniques it would be difficult to identify which part of the drill is the dominant contributor to material drag in the hoop direction but it is has now been clearly shown that the minor cutting edges are responsible for such plastic deformation. In Fig. 12, the mean torque value is plotted against drilltravelled distance for different tool wear levels and machining parameters (VB ¼ 0–0.3 mm, f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on), highlighting the increase in torque once S0 comes into contact (zone B) with the workpiece. As previously discussed in Fig. 10, the intensity of material drag substantially rises with tool wear, suggesting that an analysis of the torque signals could lead to an estimation of material drag developed on the workpiece surface. However, it should be concluded from Fig. 12 that the effect on torque is strong only for severe tool wear and its pattern can be hardly be linked with the evolution of the amount of drag material along the depth of the hole, i.e. no relationship between the amplitude TPD zone and torque level can be established for drills with acceptable (VB ¼ 0–0.2 mm) edge condition.
Fig. 12. Mean value of torque signals against the depth of the hole when employing drills with different wear levels.
5.3. Correlation of acoustic emission and material drag While the feed force and torque signals give some indication of severe wear, other signals may provide a clearer ‘‘picture’’ of the intensity and distribution of material drag along the depth of the drilled holes. With this in mind, investigations on AE signals have been performed to examine whether there is a correlation between their level/patterns and the material drag arising from the S0 –workpiece interaction. Short-time Fourier transform (STFT) was used to interpret the acoustic activity against time associated with drilling with new and worn tools as illustrated in Figs. 13a and b. When a new tool was employed to generate one hole, a marked increase in magnitude of AE signal was observed from 1.6 to 2.0 s (zones B and C—Fig. 10) across frequencies ranging from 75 to 300 kHz. Over this particular time interval, a linear drop in feed force was observed at the point where exit burr is formed. During the formation of the exit blur, a significant amount of workpiece material is plastically deformed, thereby releasing a high level of acoustic activity. However, at all other drilling time intervals, STFT at much lower amplitudes was recorded, which is presumably related to material shearing performed by the sharp major and minor cutting edges of the drill. Unsurprisingly, the STFT for the worn tool exhibits markedly higher magnitudes of the AE signal across the entire duration of drilling operation; this is believed to be related to higher degrees of friction and plastic deformation occurring during cutting with worn edges [38]. Two frequency bands (80 and 110 kHz) are clearly evident from the STFT, having increasing magnitude for the full duration (Zone B—Fig. 10) of the drilling process, during which the S0 –workpiece contact level constantly increases. Although a response across a similar frequency band might be expected to be observed when a new tool is employed, this was not to be found presumably due to much lower levels of plastic deformation, in accordance with the much lower torque levels recorded in Figs. 9 and 10. Consequently, while STFT shows good agreement with torque levels, the correlation to material drag should be sought within its details such as evolution of AE signal magnitude within particular frequency bands. As the integral of RMS AE quantifies the energy released in the workpiece during particular selected time intervals [39], it may be possible to estimate the amount of acoustic energy released during drilling in order to relate this information with the intensity of material drag at specific time intervals. Fig. 14 shows the variation of the integral of RMS AE (window size of three revolutions at 100–120 kHz bandpass filter) as a function of distance drilled. The use of RMS over the selected window interval enables a better understanding of the acoustic energies released solely from the drilling operation for a given frequency band. Since this range of frequency (100–120 kHz) appears from the AE signal in Fig. 13 to have a strong relationship with the S0 –workpiece interaction, it may be possible to determine the intensity of material drag by ‘‘supervising’’ the RMS AE signal only within this frequency band. From Fig. 14 the intensity of RMS AE clearly rises with increasing levels of tool wear. Consequently, the RMS within the 100–120 kHz frequency band does seem to correlate with higher degrees of plastic deformation (material drag) occurring due to S0 –workpiece interaction. As previously presented in Fig. 9, the thickness of material drag induced by S0 –workpiece interaction was found to decrease towards the exit hole. At the same time, the torque in Fig. 12 and RMS EAE value in Fig. 14 tend to increase linearly with drill depth. Since the thickness of material drag is a cumulative buildup of S0 –workpiece interaction (time dependent), this leads to the highest degree of deformation at the top of the hole. As RMS EAE expresses the friction energy released in the component within a
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Fig. 13. Short-time Fourier transform (STFT) of AE signals from drilling operation (f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on) against drilling time of a hole: (a) new tool (VB ¼ 0.0 mm) and (b) worn tool (VB ¼ 0.3 mm); highlighted region on which later a bandpass filter was applied to AE signal.
plot the graph. From this method it is possible to directly estimate the degree of material drag retained on the surface of the drilled hole induced by S0 –workpiece interaction. Moreover, the magnitudes of material drag and RMS EAE increase with the level of tool wear, which is proven consistently in both Figs. 9 and 14. From the above it can be concluded the RMS EAE distribution accurately reflects the magnitude of material drag caused by S0 –workpiece interaction at different cross-section levels of the hole.
5.4. Residual stress and plastic deformation distribution analysis Fig. 14. Integrals of RMS from acoustic emission against the drill-travelled distance (f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on).
specific time window (S0 –workpiece contact area dependent) hence the highest values would be observed at the exit of the hole. To understand the correlation between material drag and RMS EAE better, it is possible to inverse the x-axis (depth of hole) and re-
A combination of metallurgical analysis and signal processing has identified possible means of detecting the intensity of material drag in the hoop direction during drilling. Another important aspect for prolonging the safe operating life of safety critical components is to achieve compressive residual stresses near surface of machined components [34]. Indeed, the level of material drag might be expected to create plastic misfits around the hole; misfits that along with steep thermal gradients during
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drilling might be expected to give rise to tensile residual stresses. X-ray diffraction has been used to measure the residual stresses along the depth of the drilled holes in two different directions, i.e. axial and hoop. In this respect, the variation in diffraction peak full-width at half-maximum (FWHM) is a good indicator of the level of induced plastic work [40]; in this way the material drag caused by S0 –workpiece interaction can be directly quantified. This is expected to correlate with the above microstructural observations of material drag caused by S0 –workpiece interaction and with the previously presented relationships with sensory signals (i.e. RMS EAE).
5.4.1. Residual stress analysis in the axial direction In Fig. 15, an example of surface residual stress profiles along the depth of the holes drilled with new (VB ¼ 0.0 mm) and worn (VB ¼ 0.3 mm) tools are presented. Highly compressive axial (feed direction) residual stresses were found along the depth of the hole peaking out at 1161 MPa (749.52 MPa) for new and 1209 MPa (756.27 MPa) for worn tools. This is in good agreement with published work [41], which also found compressive stresses of approximately 1000 MPa, which extended some 75 mm below the surface. As mentioned earlier, such compressive stresses maybe the result of tensile misfits introduced by the tendency of tensile straining in the axial direction under large feed forces. Since the work focuses on the minor cutting edge–workpiece interaction that causes material drag in the hoop direction, more effort has been dedicated to analyse the residual stresses in this direction.
5.4.2. Residual stress analysis in the hoop direction By contrast, residual stresses in the hoop direction were found to be essentially tensile for both new and worn tools, except near the hole entrance for the worn tool (Fig. 15). The observation of tensile hoop stresses is perhaps surprising, but is not out of line with observations for other machining operations, such as surface finish turning [31,42,43], which shows significant tensile stresses near surface in the hoop direction, before becoming compressive at greater depths. Since the frictional forces (see comments related to Fig. 10) accounting for S0 are considerably lesser when compared to those from the major cutting edges, it is possible that steep local thermal gradients (associated with the friction phenomena) determine the surface hoop stresses. However, at the entrance of the hole where the dragging phenomena is the most significant due to the longer S0 –workpiece interaction (as described in Section 5.1), mechanical effects leading to compressive residual stresses are apparent. This suggests that the plastic deformation incurred by material dragging has reduced the level of tensile residual stress. This explanation is supported by the FWHM measurements. For the new tool very little variation in FWHM is shown with distance along the depth of the drilled hole surface, whereas for the worn tool, the FWHM whilst similar to that at the exit hole for the new tool becomes significantly greater with distance towards the entrance hole. This trend supports the results of Fig. 9, suggesting that plastic deformation increases significantly with distance from the hole exit. Hence, it can be concluded that increased S0 –workpiece interaction for a worn tool (within the testing condition and workpiece material) is sufficient to reduce the level of tensile residual stress generated from the
Fig. 15. Residual stress and FWHM profile along the depth of hole (f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on): (a) new tool (VB ¼ 0 mm) and (b) worn tool (VB ¼ 0.3 mm).
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steep thermal gradients through severe plastic deformation associated with material drag.
6. Conclusions This paper presents for the first time the influence of the minor cutting edges (S0 )–workpiece interaction exemplified in drilling of a new nickel-based superalloy, coarse grain RR1000. Investigations including surface integrity, process monitoring and residual stresses have been considered to prove the existence and outcomes of the S0 –workpiece interaction with main results and conclusions summarised as follows: 1. Material drag in the hoop direction is mainly caused by interactions between S0 and the workpiece material. The degree of material drag is proportional to the contact duration between S0 and workpiece surface such that the extent of material drag is lesser near the hole exit than near the hole entrance, creating an approximately triangular plastic deformation zone. Metallurgical evidence has shown that increasing tool wear of S0 increases both the intensity and thickness of material drag in the hoop direction. The worn tool has demonstrated a depth of material drag approximately 50% greater than that for the new tool. 2. Feed force and torque signals have demonstrated a basic relationship of the interactions between the workpiece and the major and minor (S0 ) cutting edges, especially when worn tools were employed. In spite of this, torque signals were not found to be useful in relation to the evolution of material drag along the depth of the hole (i.e. TPD zone). Conversely, analysis of the acoustic emission signals in the form of STFT highlighted the frequency bands (100–120 kHz) in which the S0 –workpiece interaction (generation of material drag in the hoop direction) is most evident. Understanding the signal using RMS EAE techniques from appropriate frequency bands delivered a more sensitive means for quantifying the level of material drag (i.e. amplitude of TPD zone) down the depth of the hole, which can be used as a base for developing an on-line surface damage system. 3. Compressive axial and tensile hoop residual stresses were measured at the surface of the hole. The tensile residual stress values along the depth of the hole for a new tool were caused by steep thermal gradients during drilling operation. However, for worn tool near the hole entrance, extensive plastic work caused by material drag appears to reduce the tensile hoop stresses to compression. The FWHM curves for the worn tool corroborate the linear increases in the torque curve, the RMS EAE signal and hence confirm the existence of the TPD zone. Acknowledgements The authors would like to thank Dr. Jamie McGourlay, Dr. Robert Mitchell and Mr. James Fortune from Rolls-Royce plc for their continuous financial and technical support towards performing the research. Special thanks also go to Dr Judith Shackleton from the Stress and Damage Characterisation Unit at the University of Manchester for her technical support during the residual stress measurements. References [1] I.A. Choudhury, M.A. El-Baradie, Machinability of nickel-base super alloys: a general review, Journal of Materials Processing Technology 77 (1998) 278–284.
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Contents lists available at ScienceDirect
International Journal of Machine Tools & Manufacture journal homepage: www.elsevier.com/locate/ijmactool
Minor cutting edge–workpiece interactions in drilling of an advanced nickel-based superalloy J. Kwong a, D.A. Axinte a,, P.J. Withers b, M.C. Hardy c a b c
School of Mechanical, Materials and Manufacturing Engineering, The University of Nottingham, Nottingham NG7 2RD, UK School of Materials, The University of Manchester, Grosvenor Street, M1 7HS, UK Rolls-Royce plc, P.O. Box 31, Derby DE24 8BJ, UK
a r t i c l e in fo
abstract
Article history: Received 4 October 2008 Received in revised form 14 January 2009 Accepted 16 January 2009 Available online 7 February 2009
Drilling is one of the key machining operations for manufacturing safety critical components that must comply with strict surface quality standards. The influence of major flank wear of drilling tools on workpiece surface quality has been well established; however, similar information concerning minor cutting edge is currently missing from literature. This paper presents a comprehensive analysis and discussions of the influence of the drill’s minor cutting edge to workpiece surface integrity and residual stress distribution for RR1000, a newly developed nickel-based superalloy. These effects are critical to the acceptance of this new material in relation to tool geometry and machining strategies. The thickness of material drag in the hoop direction has been found to be the highest at the top and the least at the bottom of the hole, which is directly related to the contact duration between the minor cutting edge and workpiece surfaces; moreover this difference increased at higher levels of wear of the minor cutting edge. On-line process monitoring techniques have been employed to further understand the material drag phenomena, including feed force, torque and acoustic emission. Compressive axial and tensile hoop stresses at the surface of the holes have been measured as a function of depth and correlated both with metallurgical analysis of drilled surfaces and the process monitoring signals. It was found that the increased material drag associated with a worn tool resulted in compressive hoop surface residual stresses near the entrance hole in correspondence with trends in the processed acoustic emission signal. This work suggests that material drag increases with the duration of the minor cutting edge–workpiece interaction such that plastic deformation is the greatest near the drill entrance holes and that process monitoring of the degree of material drag in hoop direction can be practicable. & 2009 Elsevier Ltd. All rights reserved.
Keywords: Nickel-based superalloy Drilling Material drag Residual stresses Acoustic emission
1. Introduction One of the challenges facing the aerospace manufacturing industry is to introduce increasingly high-strength superalloys, while continuously adapting manufacturing processes to enable the generation of high-integrity components at lower manufacturing costs, higher production rates and lower scrap rates. Safety critical components, such as turbine discs, are machined under strict quality standards set by aerospace manufacturers; in this respect surface integrity issues are of critical importance, especially when new superalloys are adopted. Due to their low thermal conductivity, their tendency for strain hardening and high strength at elevated temperatures, nickel-based superalloys, such as Inconel, Waspaloy and Udimet 720Li, are notoriously difficultto-cut materials [1,2]. Indeed, achieving the high degree of workpiece surface integrity required for the safe utilization of
Corresponding author.
E-mail address: [email protected] (D.A. Axinte). 0890-6955/$ - see front matter & 2009 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijmachtools.2009.01.012
aeroengine parts made of superalloys subjected to high thermal and mechanical loading is a demanding task and all the outcomes of the employed machining processes must be well understood [3]. A variety of surface damages (i.e. anomalies) such as material drag, cracking and plucking [4] can result when nickel-based superalloys are machined with inappropriate parameters or tool conditions. These anomalies can have a detrimental effect on the fatigue performance of safety critical components during their operational life [5]. Safety critical (rotating) parts (e.g. discs) contain a variety of geometrical features that, according to their anticipated loading distributions, have various degrees of importance when it comes to influencing the fatigue performance of the manufactured component. Turning, milling, broaching, grinding and hole making are all involved in the manufacture of such aerospace components. While turning [6], milling [7] and grinding [8] of nickel-based alloys have been well documented in the academic literature, as well as industrial publications, holemaking operations have not received in-depth analysis with regard to workpiece surface quality in terms of metallurgical/ geometrical integrity, residual stresses and their implications.
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This is perhaps a surprising fact given that hole features have the highest degree of criticality for the safe exploitation of rotating aerospace components loaded in fatigue conditions [9]. Wear of the major flank of the drills is the dominant mechanism limiting tool life with worn drills raising the prospect of introducing damage into the workpiece [10,11] as well as additional costs associated with the necessary re-manufacture of the out-of-specification machined components. Consequently, it is important to better understand the relationship between drill wear and quality of machined surfaces. Some information on tool wear and chip formation in drilling of nickel-based superalloys has been reported [12] but surprisingly little reference has been made to the more important aspect of the damage induced in the machined surfaces. A limited published research has been focussed only on the metallurgical inspection in the axial direction for various hole-making techniques, including drilling, reaming and plunge/spiral milling [11,13]; this is a very limitative inspection that gives information on the surface integrity of the machined holes only along feed direction, which might not be (as this report shows in the following) the ‘‘preferential’’ direction on which surface damage occurs in hole-making operations. As the main cutting edges of the drill follow a helical trajectory, analysis in the hoop (cutting) direction is needed to facilitate a more complete interpretation of the phenomena occurring at the tool–workpiece interface during drilling and thus to enable better understanding of their implications for surface integrity, including metallurgical (e.g. phase transformations) and mechanical (e.g. residual stress) aspects. Typically drilling tools (Fig. 1a) have a cylindrical geometry [14] (i.e. not back tapered), which has raised concerns within the aerospace machining community on issues
related to the influence of continuous interactions between the minor cutting edge (S00 ) and the walls of the workpiece on the integrity of final machined surfaces. Previous research has examined the wear of minor cutting edges in drilling [15], but to the authors’ knowledge, its effect on workpiece surface integrity has not been reported. This aspect could be of crucial importance on the hole-making methods when it comes to generating damage-free surfaces like those required in aerospace industry. In view of the importance of tool wear in drilling, on-line wear monitoring systems to help decide when to schedule tool replacement are currently in development as part of the move towards fully automated manufacturing cells. Tool condition monitoring systems allow the prediction of conditions that can lead to catastrophic damage modes. Traditionally, tool wear has been measured optically, but this requires off-line inspection. To avoid this, on-line process monitoring techniques have been employed to determine tool wear levels from readings of the cutting forces, torque, spindle power, vibrations and acoustic emission (AE) signals. As with other conventional machining operations, in drilling, forces increase as tool wear develops [16,17]. Often only the feed force is considered, but recent research has advocated condition monitoring using simultaneous consideration of feed force along with cutting torque [18]. While a direct representation of the cutting forces in drilling can be based on the use of torque readings, the linear increase of torque signal, over the period that the minor cutting edges of the drill come into contact with the workpiece until full completion of the hole, remains unexplained. Spindle current/power signals have been shown to be a reliable way of detecting tool breakage, but
Fig. 1. (a) Basic geometrical specifications of a two flutes twist drill geometry (ISO 2003/1). (b) Illustration of the various stages/cutting times (t1–t5) of the drill minor cutting edges (S0 )–workpiece interaction: t1—corners of drill reaches workpiece; t2—S0 –workpiece interaction begins; t3—material drag increases cumulatively; t4)—corners of drill exits workpiece; t5—material drag continues until tool retraction.
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insufficient to provide a robust means of determining lesser increases in tool wear [19,20]. It is well known that the main contributors of AE in machining come from friction at the clearance faces (for both major and minor cutting edges), plastic deformations in the shear zone and chip breakage [21]. Research has shown that the energy count (Measured Area of the Rectified Signal Envelope—MARSE) of the AE signal increases with tool wear; whereas envelope analysis is good for detecting tool wear only at transient stages of the drilling process [22]. Analytical techniques for prediction of surface quality from machining have been previously employed, including frequency analysis [23] and interpretation of RMS signal [24]. The use of vibration signals has been employed, but since these contain both ‘‘useful’’ and noise signals, the latter can have a significant effect on the robustness of such techniques [25]. In summary, all the current process monitoring techniques focus mainly on detecting excessive flank wear and tool breakages in the drilling operation. By contrast, information relating process signals to the occurrence of damage to workpiece surface integrity is very limited in machining, which includes the defection of plucking [26], white layers [27] and material drag [28] from AE analysis. In particular, information is missing concerning minor cutting edge–workpiece interactions and their implications for nearsurface damage modes, machining force/torque signals and AE signals. In general, machining may result in a high degree of plastic deformation and strain on the workpiece surface when the cutting parameters are not optimised from a surface quality viewpoint [29]. Non-uniform plastic deformation will generally give rise to locked-in residual stresses, comprising balancing compressive and tensile stresses [30,31]. In machining, the nature of the residual stress is determined by two factors, namely mechanical and thermal effects. Mechanically induced residual stress arises from plasticity deformed region near the tool–workpiece interface; the material ahead of the cutting edge is believed to be in compression whilst material behind will be in tension. As a consequence, if plasticity ahead of the cutting edge is dominant then the residual stresses caused by mechanical effects would be tensile; however if plasticity behind the tool is more influential then this favors compressive stresses [32]. Locally significant temperatures can be generated as a combination of plastic deformation of the workpiece and friction at the tool–workpiece interface. The sharp thermal gradients generated in such a way can contribute to tensile stresses in the superficial layer of the workpiece [33]. Tensile residual stresses in the surface and subsurface regions are known to significantly impair fatigue resistance of critical components during operation of critical gas turbine components whereas compressive will prolong the operating life [34]. In literature, reports on the measurements of residual stress associated with drilling are scarce. Furthermore, the reports tend to focus only on stresses in the axial direction and measurements in the hoop direction have not been considered primarily due to the difficulties of making X-ray measurements on a tightly curved surface [35]. Moreover, mapping of the residual stress profiles along the height and beneath the free surfaces of drilled holes have received little attention, preventing the establishment of an understanding of the process by which residual stresses build up during the drilling operation.
2. Scope of research This paper is the first to focus attention on the influences of the interaction between the drill’s minor cutting edge (S0 ) and workpiece surface that can affect the integrity of machined surfaces. This is discussed in relation to drilling of a new nickel-
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based superalloy, i.e. RR1000, which is regarded as a nextgeneration aerospace material for the manufacture of safety critical parts. RR1000 possesses superior properties, including 25–50 1C higher operating temperature capability over the current disc alloy, 720Li, whilst having an equivalent crack growth resistance to coarse grain Waspaloy [36]. These properties make RR1000 a more difficult-to-cut material than the existing nickelbased alloys (e.g. Inconel, Udimet 720), making it more susceptible to the generation of surface anomalies caused by minor cutting edge during drilling operation. Consequently, this paper addresses the following aspects: (a) investigation of the S0 –workpiece interactions and their effect on the surface integrity; (b) in-depth analysis and comparison of metallographic, 2D residual stress distributions and plastic deformations in the superficial layer of the workpiece to enable quantitative and qualitative evaluations of the outcomes of S0 –workpiece interactions; (c) assessment of potential on-line process monitoring techniques, including feed force, torque and acoustic emission signals as a means of determining the magnitude of S0 –workpiece interaction phenomena to propose process supervision methods that enable machining of damage-free surfaces.
3. Analysis of minor cutting edge–workpiece interactions The mechanism of the minor cutting edge–workpiece interaction can be broken down and discussed in different stages (t1–t5) as presented in Fig. 1b and is discussed in the following: (a) In the first stage of drilling, tip of the drill comes into contact with the workpiece surface and the contact area with the major cutting edges starts to increase. Once the major cutting edges have fully immerged into the workpiece at time, t1, the minor cutting edges of the drill begin to interact with the workpiece surface (at the hole depth), d1. Until, t1, hoop plastic deformation (material drag) caused solely by the minor cutting edges (S0 ) cannot take place. (b) At time t2, the drill progresses deeper into the workpiece and the total contact area between the S0 and workpiece surface increases proportionally with cutting time. (c) During this stage, the S0 –workpiece interaction can cause material from the surface to plastically deform in the hoop (cutting) direction by a dragging mechanism. The extent of the interaction is dependent on the time that the minor cutting edges of the drill are in contact with the workpiece material (walls of the hole) and the intensity of the associated friction phenomena (‘‘embedded’’ within cutting torque). Consequently the material near the entrance of the hole is exposed to interaction with the minor cutting edges of the drill for longer time than that near the exit, allowing for a larger cumulative dragging effect. Thus, a ‘‘triangular plastic deformation’’ (TPD) distribution might be expected to describe the volume of material drag caused by S0 –workpiece interaction at any time during the drilling operation. Of course, the extent of shearing within this zone may be limited if it exceeds a critical value, leading to local material fractures. (d) On increasing the drilling depth to d3 at time t3, the TPD zone extends in both axial and hoop directions at a rate proportional to the total cutting time, provided contact between the S0 and the workpiece is maintained along the depth of the drilled hole. (e) When the major cutting edges of the drill exit from the workpiece at time t4, the complete depth of the hole has
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experienced the minor cutting edge–workpiece interaction and the TPD zone would be expected to extend through the whole depth of the hole. (f) During the push-through period, at time t5 the only contact between the drill and the workpiece is with the minor cutting edges as the major cutting edges have fully emerged from bottom of the hole. At this time, material dragging and heat generation continues and can occur until the tool is retracted, causing the generation of plastic deformation to persist even after the corners of the drill exit the hole at depth d4. This analysis suggests that the longer the S0 –workpiece interaction, the greater the material drag in the hoop (cutting) cutting direction; this is provided that continuous contact between the minor cutting edges and wall of the drilled is maintained behind the tip, a situation that is likely to happen when employing cylindrical drills in generating relatively short holes (h/dE1). Therefore, in the light of these considerations, regions near the start (top) of the hole would be expected to experience a higher material drag than those at the end, in this way leading to the creation of a TPD zone. Additionally, with the increased flank wear of the major cutting edges it is expected that more stock of material is left for removal by minor cutting edges, which is likely to result in more material dragged in the hoop direction; therefore, in such instances, it is expected that TPD zone will expand deeper into drilled surfaces, a situation that will be proven later in the paper. TPD zone that is the result of minor cutting edges–workpiece interaction is experimentally investigated in this paper for new, intermediate and worn drilling tools. In addition, the extent to which both the mechanical effect and the heating effect may give rise to residual stresses that vary along the depth of the hole has also been examined. If a build-up of material drag due to the S0 –workpiece interaction is identified, appropriate measures to reduce this effect could be developed (i.e. on-line monitoring
solutions) to enable high workpiece surface integrity components for sensitive applications (i.e. safety critical component) such as those in aerospace industry.
4. Experimental set-up and methodology In the following, testing programme drills of new, intermediate and worn cutting edges have been employed for hole making on the proprietary nickel-based CG RR1000 superalloy manufactured via the powder metallurgy route [36], which display the following chemical composition: 15% Cr, 18.5% Co, 5% Mo, 3% Al, 3.6%Ti, 2% Ta, 0.5% Hf, 0.06% Zr, 0.027% C, 0.015% B, Ni balance. The high strength at elevated temperatures of CG RR1000 derives from a combination of solid solution strengthening and g0 precipitates [37], which makes it difficult to cut with a high tendency to retain tensile residual stresses after machining. Cylindrical plates of CG RR1000 were ground on both sides to ensure the workpiece to be flat when placed in a special rig for drilling. Prior to any machining, hardness and residual stress measurements have been performed on the workpiece to ensure that the sample has not been work hardened or does not have pre-machining induced residual stresses. A five-axis Makino A55 machine equipped with Corogrip Precision Power Chuck displaying reduced tool run-out (o2 mm) has been employed throughout the experimental trials. A waterbased synthetic cutting fluid (Houghton 3380) was supplied through tool for selected drilling trials. A circular array of holes was drilled; this enabled concentric rings to be cut by electrodischarge machining (EDM) for subsequent workpiece surface inspections. A process monitoring set-up (Fig. 2) consisting of a Kistler four-axis dynamometer (9272) acquiring at 10 kHz torque and x, y and z force signals was mounted behind the fixturing system. An AE sensor (Kistler 815B121) was thread mounted within the fixturing system. The AE signals were acquired at a
Fig. 2. Schematic illustration of drilling set-up (a) and fixturing system on Makino A55 (b).
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frequency of 800 kHz to enable spectral analysis of the signal within the response frequency of the sensor (50–400 kHz). The acquired signals were then transferred via two synchronised data acquisition cards (NI PCI-MIO-16E-dynamometer signals, NI-6115AE sensor) and saved/processed using an in-house LabView programme. To develop an appreciation of the operating window on which the material can be satisfactory drilled, a wide range of drilling parameters has been tested. To address the scope of the paper, only those trials needed to exemplify the significance of the S0 –workpiece interactions on surface integrity are reported here. Factors such as tool wear (VBmax ¼ 0–0.3 mm), cutting parameters (f ¼ 0.08–0.1 mm/rev, n ¼ 1061–1592 RPM) and coolant conditions (dry or through tool coolant supply at 30 bar) were varied. Except where indicated, the discussions on the S0 –workpiece interactions will predominantly refer to three cutting conditions employing new, intermediate (VB ¼ 0.1 mm) and worn (VB ¼ 0.3 mm) tool used at a constant feed rate, f ¼ 0.10 mm/rev and variable spindle speed, n ¼ 1061–1592 RPM whilst keeping coolant on. The machined holes were initially sectioned along the axial direction via cut-off wheels to obtain two halves; one used for residual stress measurements and the other for metallurgical analysis. To perform scanning electron microscope (SEM) inspections, it was necessary to section the hole in the hoop direction at 1 mm intervals along depth of the holes as illustrated in Fig. 3. The specimens were then mounted in conductive bakelite for in-depth micro-scale metallurgical analysis using an environmental field emission gun SEM (Philips XL30 ESEM-FEG). The specimens were ground using silicon carbide paper through to 2400 grit and polished to a 1 mm finish on DUR (6 mm) and NAP (1 mm) cloths. Microstructure was revealed using electrolytic etching with 10% ortho-phosphoric acid in distilled water using a current of 2–3 A with stainless-steel wires as anode and cathode.
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Residual stress measurements were performed on a Proto X-ray diffractometer using Mn Ka radiation (l ¼ 2.1031 A˚) to acquire the (3 11) diffraction peak at a 2y angle of approximately 1521 using a spot size of 1 mm diameter. Prior to the measurements of the samples, a reference (flat and stress free sample of RR1000) was measured. All samples were sectioned to ensure no shadowing effect or stress relaxation would occur during the measurements, which would otherwise affect the accuracy of the results. This was confirmed with the comparison against the defined gain. Sin2 C technique measurements were performed at multiple C tilt angles (b angles with the Proto iXRD) with a 31 oscillation over a C range of 7271. The penetration depth of Mn Ka radiation in Ni is approximately 22 mm considering 99% absorption by the workpiece at C ¼ 01 with negligible changes in the range of tilts performed in the full measurement. The Young’s modulus and Poisson’s ratio used for the {3 11} reflection were 204 GPa and 0.27, respectively.
5. Experimental results and discussion 5.1. Characterisation of material drag in drilling In Fig. 4, preliminary metallographic examinations in the axial direction after drilling operation when either new or worn tool are employed do not reveal any surface anomalies. The clean surface generated from a worn tool might be a surprising observation as typically worn cutting edges have a high tendency to generate surface damages (i.e. material drag, cracks). However, as presented in Fig. 1b, the main drill cutting edges follow helical paths and therefore the generation of surface anomalies may be more dominant in the hoop direction whilst leaving the surface in axial direction much less damaged. To provide a better understanding, metallurgical analysis in hoop direction must be performed and
Fig. 3. Preparation of specimen for SEM analysis in axial and hoop directions.
Fig. 4. Secondary electron images ( 5000) of the workpiece surface viewed in the axial direction (VB ¼ 0.0 mm, f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on): (a) new tool and (b) worn tool.
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more importantly to evaluate the influence of S0 (minor cutting edge) to the surface finish from drilling operation. To explore the severity of material drag caused by S0 (minor cutting edge)–workpiece interaction, in Fig. 5 an axially sectioned hole with an embedded broken tool (VB ¼ 0.3 mm, f ¼ 0.08 mm/rev,
Fig. 5. Axial sectioned optical image of blind hole illustrating the S0 –workpiece interaction.
n ¼ 1326 RPM, coolant on) is presented; this gives an indication of the possible local ‘‘welds’’ between S0 and the wall of the drilled hole. The circled regions (Fig. 5) illustrate that S0 retains intimate contact with the workpiece well behind the tip. Indeed, in extreme cases such as these, local S0 –workpiece welding can contribute (along with high cutting loads at the main cutting edges) to drill breakage. Moreover, for the S0 and workpiece to locally weld, a high local (minor cutting edge) temperature, likely caused by the intensive friction phenomena, must be present, which also facilitates plastic deformation on the surfaces in the hoop direction. Metallographic examination of a hole produced using a new drill (VB ¼ 0 mm) is illustrated in Fig. 6. Preliminary investigations of metallographic analysis in the axial (feed) direction show negligible amount of material drag beneath drilled surfaces; hence the follow-up examinations have been focussed on the hoop direction. The thickness (measured in the radial direction of the hole) of the plastic deformation zone was quantified as the distance from the free surface of the workpiece to the depth beneath the surface where no more plastic deformation is visible on the micrographs; lower magnifications ( 5000) enable the observation of the depth of material drag while higher magnification pictures ( 10,000) put in evidence the angle of the dragged grains. From the SEM images, it is clear that the depth of material drag tends to decrease from the top (Figs. 6a and b) to the bottom (Figs. 6e and f) of the holes. Material drag observed on the surfaces is caused by the ‘‘ploughing’’ interaction between the S0 and walls of
Fig. 6. Secondary electron images ( 5000 on left column and 10,000 on right column) of the workpiece surface viewed in the hoop direction at various depths: (a, b) 1 mm, (c, d) 3 mm and (e, f) 5 mm when using a new drill (VB ¼ 0.0 mm, f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on).
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the hole, which follows the direction of the cutting action. As the contact time between the S0 and workpiece is lower near the exit of the hole, a lesser degree of material drag has been observed (Fig. 6e). As previously explained, from the point of view of the material drag caused by S0 the plastic zone might be expected to have ‘‘zero’’ depth at the bottom of the hole in the hoop direction. Although the major cutting edges do not directly come into contract with the walls of the workpiece, a somewhat pushedthrough effect induces a small degree of material drag in the hoop direction (Fig. 6e). Unsurprisingly, the extent of material drag was found to be more significant when employing worn tools as illustrated in Fig. 7. Material drag to a depth of 32 mm (cf. 20 mm for the new tool) was measured 1 mm from the top of the hole, decreasing to 2 mm at 5 mm depth. This is almost certainly related to the increased tool wear on the minor cutting edges of the drill (Figs. 8a and b) since this is the main contact between the drill and walls of the hole. However, it should also noted that the corners of the drill are most susceptible to wear during cutting (Figs. 8c and d); this is likely to lead to an initially smaller hole diameter and thus, leaving more stock of material for removal by the S0 . Since the material stock removed by the S0 is much lesser than those from the major cutting edges, it is likely that S0 will have a tendency to drag the surface material rather than cut it. Hence, more intense material drag is expected when worn tools are employed. Therefore, the wear of the S0 is also likely to add to this material dragging phenomenon. However, to separate the effects of wear from the major and minor cutting edges on
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material drag is difficult to be evaluated unequivocally and this is an issue under current investigation. The intensity of material drag was found to be more significant at the top of the hole. As previously commented in Section 3 (Fig. 1b), the top of the hole experiences the longest duration of S0 –workpiece interaction, hence causing the maximum material drag in the TPD zone. This effect is evidenced in Figs. 7a and b with the surface layer smeared to such an extreme extent that cracks are visible to a radial depth of around 4 mm. In Fig. 9, the depth of the plastically dragged layer is plotted as a function of depth (1–5 mm) of the hole when employing new and worn drills; values of these measurements are presented in Table 1. From this graph it is clear that the plastically deformed zone is approximately triangular in form and is approximately 50% deeper for the worn tool. Therefore, the initial assumption on the generation of a TPD zone is caused by S0 –workpiece interaction, which has a good correlation with the results of the metallurgical/surface integrity investigations. 5.2. Evolution of feed force and torque To support the understanding of material drag, Fig. 10 shows the torque and feed force signals when drilling CG RR1000 for both new and worn drills with the same cutting conditions. The distinct drilling stages seen in relation to the pattern of sensory signals (Fig. 10) are discussed below. A. Tip entrance: From the moment when the drill tip comes into contact with the workpiece until full immersion of the major
Fig. 7. Secondary electron images ( 5000 on left column and 10,000 on right column) of the workpiece surface viewed in the hoop direction at various depths: (a, b) 1 mm, (c, d) 3 mm and (e, f) 5 mm when using a worn drill (VB ¼ 0.3 mm, f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on).
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Fig. 8. Optical images of drills illustrating different degrees of tool wear (a) new tool–minor cutting edge; (b) worn tool–minor cutting edge; (c) new tool–major cutting edge; (d) worn tool–major cutting edge.
Fig. 9. Scatter plot from averaged values (five measurements) of the thickness of material drag against depth of hole.
Table 1 Averaged thickness of material drag obtained using new and worn drilling tools against the depth of hole (results based on five measurements). Depth of hole (mm)
1 2 3 4 5
Material drag thickness (mm) New tool
Worn tool
0.020 0.018 0.012 0.004 0.002
0.032 0.033 0.018 0.012 0.002
cutting edges, a sharp increase of feed force and torque signals is observed. When a worn drill is employed, the increase in feed force substantially exceeds that for a fresh drill. This can be explained by the decrease in cutting ability when the major cutting edges are worn. B. Development of minor cutting edges–workpiece interaction: Once the major cutting edges are fully immersed into the workpiece, feed force (Fz) and torque (T) reach a steady state, at least for the new tool. At this stage, the contact area of S0 with the walls of the hole increases linearly with cutting time, which contributes to the increase of T as the drill progresses through the depth of the hole. This is where the generation of the TPD distribution begins and the development of material drag initiates. It is clear that feed force varies little with increasing drill depth, being larger for the worn tool. On the other hand, torque appears to increase linearly with depth, the gradient being slight for the new tool and much steeper for the worn tool. As the drill progresses deeper into the hole, the extent of S0 –workpiece contact increases proportionally with depth. The observed linear increase of torque thus corroborates the initial consideration that material drag originates from the workpiece interaction with the minor cutting edges and giving rise to the TPD zone. Further, it is not surprising that this linear increase is evident to a much lesser degree for the new tool as the minor cutting edges efficiently shear off material rather than dragging them. In Fig. 10, the dashed line (TIdeal line) represents the ‘‘ideal’’ level of torque, being equal to that generated solely by the major cutting edges of the drill. It should also be noted the feed force starts to decrease just before the tip exits the workpiece for both new and worn drills. At this depth, the remaining material ahead of the drill has been plastically deformed under the feed force rather than being appropriately cut. This effect is observed only during the latter stages in the full immersion zone since the remainder of workpiece material is significantly thinner leading to reduced strength. C. Tip exit: As the tip of the new drill exits the workpiece an instantaneous decrease in both Fz and T is observed. This is in contrast with the response when a worn drill is used. Instead, a sudden drop in both feed force and torque was not measured until a point 0.8 mm beyond the expected exit of the workpiece
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Fig. 10. Torque and feed force signals vs. time (lower axis) and distance travelled (upper axis) for both new (VB ¼ 0.0 mm) and worn (VB ¼ 0.3 mm) drills in the same cutting conditions (f ¼ 0.10 mm/rev, n ¼ 1592 RPM, coolant on) showing different drilling stages: A—tip entrance, B—full immersion, C—tip exit, D—drill push through and E—drill retraction.
Fig. 11. 3D mapping of plastic deformations/burrs formed at exit of drilled holes (f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on): (a) new tool (VB ¼ 0 mm) and (b) worn tool (VB ¼ 0.3 mm).
(at 6 mm). In this case, as the main cutting edges of the drill are worn they cause more material to be plastically deformed beyond the ‘‘theoretical’’ edge of the workpiece. In this context, Fig. 11 shows an increase from around 108 to around 187 mm in the height of the exit burrs when a worn drill is used. The larger burr formed by a worn tool resulted in increased duration of the Fz and T signals, which is due to the increased contact time between the tool and workpiece surface. As the corners of drill wear out, this leads to lesser material being removed by the major cutting edges;
the ‘‘formed’’ hole has a slightly smaller diameter than the nominal one. The remainder of stock material will be removed by the minor cutting edges and hence this results in extended contract duration measured as feed force and torque. For 0.8 mm of travel beyond the nominal exit hole position the torque remains essentially undiminished since S0 edge–workpiece interaction is still present until the geometry of the exit burr has stabilized. Since the torque reaches a maximum level once the major cutting edges have exited the nominal hole, this corroborates the theory
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that S0 –workpiece interaction is present in drilling. Moreover, S0 –workpiece interface is the main contact between the drill and the walls of the workpiece; therefore such interactions are the main contributor to material dragging in the hoop direction and the formation of TPD zone. D. Drill push through: During this stage no more material can be removed by the major cutting edges of the drill and hence no feed force is recorded for either the new or worn drill. However the torque, which is negligible for the new tool, declines significantly (from 5.3 to 2.6 Nm) over this region. During this period the only interaction between the drill and the workpiece occurs via the minor cutting edges and thus one can deduce, in confirmation of the results for stage B, that the torque arising from the S0 –workpiece interaction is significant for the worn drill and considerably smaller for the new drill. At this particular stage of drilling, the burrs at the exit of the hole are pushed out with increasing hole diameter by S0 , and hence increasing the contact duration and hence reducing the area of contact and torque. E. Drill retraction: Once the drill travels the preset distance, it is retracted; this is revealed through a negative feed force, which demonstrates that S0 is still in contact with the workpiece and that the machined hole is still being deformed with material pushed through during the latter stages of full immersion. Process monitoring and signal analysis of the drilling operation from the perspective of both feed force and torque have provided a proof for minor cutting edges–workpiece interaction of during drilling operation while giving information of the patterns of these signals for different levels of cutting edge conditions. Without such process monitoring techniques it would be difficult to identify which part of the drill is the dominant contributor to material drag in the hoop direction but it is has now been clearly shown that the minor cutting edges are responsible for such plastic deformation. In Fig. 12, the mean torque value is plotted against drilltravelled distance for different tool wear levels and machining parameters (VB ¼ 0–0.3 mm, f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on), highlighting the increase in torque once S0 comes into contact (zone B) with the workpiece. As previously discussed in Fig. 10, the intensity of material drag substantially rises with tool wear, suggesting that an analysis of the torque signals could lead to an estimation of material drag developed on the workpiece surface. However, it should be concluded from Fig. 12 that the effect on torque is strong only for severe tool wear and its pattern can be hardly be linked with the evolution of the amount of drag material along the depth of the hole, i.e. no relationship between the amplitude TPD zone and torque level can be established for drills with acceptable (VB ¼ 0–0.2 mm) edge condition.
Fig. 12. Mean value of torque signals against the depth of the hole when employing drills with different wear levels.
5.3. Correlation of acoustic emission and material drag While the feed force and torque signals give some indication of severe wear, other signals may provide a clearer ‘‘picture’’ of the intensity and distribution of material drag along the depth of the drilled holes. With this in mind, investigations on AE signals have been performed to examine whether there is a correlation between their level/patterns and the material drag arising from the S0 –workpiece interaction. Short-time Fourier transform (STFT) was used to interpret the acoustic activity against time associated with drilling with new and worn tools as illustrated in Figs. 13a and b. When a new tool was employed to generate one hole, a marked increase in magnitude of AE signal was observed from 1.6 to 2.0 s (zones B and C—Fig. 10) across frequencies ranging from 75 to 300 kHz. Over this particular time interval, a linear drop in feed force was observed at the point where exit burr is formed. During the formation of the exit blur, a significant amount of workpiece material is plastically deformed, thereby releasing a high level of acoustic activity. However, at all other drilling time intervals, STFT at much lower amplitudes was recorded, which is presumably related to material shearing performed by the sharp major and minor cutting edges of the drill. Unsurprisingly, the STFT for the worn tool exhibits markedly higher magnitudes of the AE signal across the entire duration of drilling operation; this is believed to be related to higher degrees of friction and plastic deformation occurring during cutting with worn edges [38]. Two frequency bands (80 and 110 kHz) are clearly evident from the STFT, having increasing magnitude for the full duration (Zone B—Fig. 10) of the drilling process, during which the S0 –workpiece contact level constantly increases. Although a response across a similar frequency band might be expected to be observed when a new tool is employed, this was not to be found presumably due to much lower levels of plastic deformation, in accordance with the much lower torque levels recorded in Figs. 9 and 10. Consequently, while STFT shows good agreement with torque levels, the correlation to material drag should be sought within its details such as evolution of AE signal magnitude within particular frequency bands. As the integral of RMS AE quantifies the energy released in the workpiece during particular selected time intervals [39], it may be possible to estimate the amount of acoustic energy released during drilling in order to relate this information with the intensity of material drag at specific time intervals. Fig. 14 shows the variation of the integral of RMS AE (window size of three revolutions at 100–120 kHz bandpass filter) as a function of distance drilled. The use of RMS over the selected window interval enables a better understanding of the acoustic energies released solely from the drilling operation for a given frequency band. Since this range of frequency (100–120 kHz) appears from the AE signal in Fig. 13 to have a strong relationship with the S0 –workpiece interaction, it may be possible to determine the intensity of material drag by ‘‘supervising’’ the RMS AE signal only within this frequency band. From Fig. 14 the intensity of RMS AE clearly rises with increasing levels of tool wear. Consequently, the RMS within the 100–120 kHz frequency band does seem to correlate with higher degrees of plastic deformation (material drag) occurring due to S0 –workpiece interaction. As previously presented in Fig. 9, the thickness of material drag induced by S0 –workpiece interaction was found to decrease towards the exit hole. At the same time, the torque in Fig. 12 and RMS EAE value in Fig. 14 tend to increase linearly with drill depth. Since the thickness of material drag is a cumulative buildup of S0 –workpiece interaction (time dependent), this leads to the highest degree of deformation at the top of the hole. As RMS EAE expresses the friction energy released in the component within a
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Fig. 13. Short-time Fourier transform (STFT) of AE signals from drilling operation (f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on) against drilling time of a hole: (a) new tool (VB ¼ 0.0 mm) and (b) worn tool (VB ¼ 0.3 mm); highlighted region on which later a bandpass filter was applied to AE signal.
plot the graph. From this method it is possible to directly estimate the degree of material drag retained on the surface of the drilled hole induced by S0 –workpiece interaction. Moreover, the magnitudes of material drag and RMS EAE increase with the level of tool wear, which is proven consistently in both Figs. 9 and 14. From the above it can be concluded the RMS EAE distribution accurately reflects the magnitude of material drag caused by S0 –workpiece interaction at different cross-section levels of the hole.
5.4. Residual stress and plastic deformation distribution analysis Fig. 14. Integrals of RMS from acoustic emission against the drill-travelled distance (f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on).
specific time window (S0 –workpiece contact area dependent) hence the highest values would be observed at the exit of the hole. To understand the correlation between material drag and RMS EAE better, it is possible to inverse the x-axis (depth of hole) and re-
A combination of metallurgical analysis and signal processing has identified possible means of detecting the intensity of material drag in the hoop direction during drilling. Another important aspect for prolonging the safe operating life of safety critical components is to achieve compressive residual stresses near surface of machined components [34]. Indeed, the level of material drag might be expected to create plastic misfits around the hole; misfits that along with steep thermal gradients during
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drilling might be expected to give rise to tensile residual stresses. X-ray diffraction has been used to measure the residual stresses along the depth of the drilled holes in two different directions, i.e. axial and hoop. In this respect, the variation in diffraction peak full-width at half-maximum (FWHM) is a good indicator of the level of induced plastic work [40]; in this way the material drag caused by S0 –workpiece interaction can be directly quantified. This is expected to correlate with the above microstructural observations of material drag caused by S0 –workpiece interaction and with the previously presented relationships with sensory signals (i.e. RMS EAE).
5.4.1. Residual stress analysis in the axial direction In Fig. 15, an example of surface residual stress profiles along the depth of the holes drilled with new (VB ¼ 0.0 mm) and worn (VB ¼ 0.3 mm) tools are presented. Highly compressive axial (feed direction) residual stresses were found along the depth of the hole peaking out at 1161 MPa (749.52 MPa) for new and 1209 MPa (756.27 MPa) for worn tools. This is in good agreement with published work [41], which also found compressive stresses of approximately 1000 MPa, which extended some 75 mm below the surface. As mentioned earlier, such compressive stresses maybe the result of tensile misfits introduced by the tendency of tensile straining in the axial direction under large feed forces. Since the work focuses on the minor cutting edge–workpiece interaction that causes material drag in the hoop direction, more effort has been dedicated to analyse the residual stresses in this direction.
5.4.2. Residual stress analysis in the hoop direction By contrast, residual stresses in the hoop direction were found to be essentially tensile for both new and worn tools, except near the hole entrance for the worn tool (Fig. 15). The observation of tensile hoop stresses is perhaps surprising, but is not out of line with observations for other machining operations, such as surface finish turning [31,42,43], which shows significant tensile stresses near surface in the hoop direction, before becoming compressive at greater depths. Since the frictional forces (see comments related to Fig. 10) accounting for S0 are considerably lesser when compared to those from the major cutting edges, it is possible that steep local thermal gradients (associated with the friction phenomena) determine the surface hoop stresses. However, at the entrance of the hole where the dragging phenomena is the most significant due to the longer S0 –workpiece interaction (as described in Section 5.1), mechanical effects leading to compressive residual stresses are apparent. This suggests that the plastic deformation incurred by material dragging has reduced the level of tensile residual stress. This explanation is supported by the FWHM measurements. For the new tool very little variation in FWHM is shown with distance along the depth of the drilled hole surface, whereas for the worn tool, the FWHM whilst similar to that at the exit hole for the new tool becomes significantly greater with distance towards the entrance hole. This trend supports the results of Fig. 9, suggesting that plastic deformation increases significantly with distance from the hole exit. Hence, it can be concluded that increased S0 –workpiece interaction for a worn tool (within the testing condition and workpiece material) is sufficient to reduce the level of tensile residual stress generated from the
Fig. 15. Residual stress and FWHM profile along the depth of hole (f ¼ 0.10 mm/rev, n ¼ 1061–1592 RPM, coolant on): (a) new tool (VB ¼ 0 mm) and (b) worn tool (VB ¼ 0.3 mm).
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steep thermal gradients through severe plastic deformation associated with material drag.
6. Conclusions This paper presents for the first time the influence of the minor cutting edges (S0 )–workpiece interaction exemplified in drilling of a new nickel-based superalloy, coarse grain RR1000. Investigations including surface integrity, process monitoring and residual stresses have been considered to prove the existence and outcomes of the S0 –workpiece interaction with main results and conclusions summarised as follows: 1. Material drag in the hoop direction is mainly caused by interactions between S0 and the workpiece material. The degree of material drag is proportional to the contact duration between S0 and workpiece surface such that the extent of material drag is lesser near the hole exit than near the hole entrance, creating an approximately triangular plastic deformation zone. Metallurgical evidence has shown that increasing tool wear of S0 increases both the intensity and thickness of material drag in the hoop direction. The worn tool has demonstrated a depth of material drag approximately 50% greater than that for the new tool. 2. Feed force and torque signals have demonstrated a basic relationship of the interactions between the workpiece and the major and minor (S0 ) cutting edges, especially when worn tools were employed. In spite of this, torque signals were not found to be useful in relation to the evolution of material drag along the depth of the hole (i.e. TPD zone). Conversely, analysis of the acoustic emission signals in the form of STFT highlighted the frequency bands (100–120 kHz) in which the S0 –workpiece interaction (generation of material drag in the hoop direction) is most evident. Understanding the signal using RMS EAE techniques from appropriate frequency bands delivered a more sensitive means for quantifying the level of material drag (i.e. amplitude of TPD zone) down the depth of the hole, which can be used as a base for developing an on-line surface damage system. 3. Compressive axial and tensile hoop residual stresses were measured at the surface of the hole. The tensile residual stress values along the depth of the hole for a new tool were caused by steep thermal gradients during drilling operation. However, for worn tool near the hole entrance, extensive plastic work caused by material drag appears to reduce the tensile hoop stresses to compression. The FWHM curves for the worn tool corroborate the linear increases in the torque curve, the RMS EAE signal and hence confirm the existence of the TPD zone. Acknowledgements The authors would like to thank Dr. Jamie McGourlay, Dr. Robert Mitchell and Mr. James Fortune from Rolls-Royce plc for their continuous financial and technical support towards performing the research. Special thanks also go to Dr Judith Shackleton from the Stress and Damage Characterisation Unit at the University of Manchester for her technical support during the residual stress measurements. References [1] I.A. Choudhury, M.A. El-Baradie, Machinability of nickel-base super alloys: a general review, Journal of Materials Processing Technology 77 (1998) 278–284.
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