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Residual stresses in face finish turning of high strength nickel-based superalloy
Journal of Materials Processing Technology 209 (2009) 4896–4902
Contents lists available at ScienceDirect
Journal of Materials Processing Technology journal homepage: www.elsevier.com/locate/jmatprotec
Residual stresses in face finish turning of high strength nickel-based superalloy W. Li a , P.J. Withers a,∗ , D. Axinte b , M. Preuss a , P. Andrews c a
School of Materials, University of Manchester, Grosvenor St M1 7HS, UK School of Mechanical, Materials and Manufacturing Engineering, The University of Nottingham, University Park, Nottingham NG7 2RD, UK c Rolls-Royce plc PO Box 31, Derby DE24 8BJ, UK b
a r t i c l e
i n f o
Article history: Received 16 July 2008 Received in revised form 22 December 2008 Accepted 15 January 2009 Keywords: Superalloy X-ray diffraction Plastic work Tooling condition Turning
a b s t r a c t This paper investigates the residual stresses distributions introduced in a new generation nickel-based superalloy RR1000 by surface finish turning. The residual stresses introduced as a function of depth have been analysed for a series of machining trials with round and rhombic inserts, coated and uncoated inserts, new and worn tools, and chipped tool. The residual stress depth profiles obtained by X-ray diffraction, and layer removal show that the tool type, tool coating, tool wear and tool breakage influence the residual stress. The extent of plastic deformation for different cutting conditions has been inferred from the different peak width. Overall, residual stresses tend to have a tensile character at all depths in the hoop direction, but exhibit a significant compressively stressed zone in the radial direction. © 2009 Elsevier B.V. All rights reserved.
1. Introduction Recent metallurgical developments in the gas turbine industry have been directed at a new generation of nickel-based superalloys to meet the demand of higher operating temperatures. These contain even higher volume fractions of strengthening ␥ precipitates and refractory elements than conventional superalloys, thereby achieving higher strength at high temperatures. This combined with low thermal conductivity makes them difficult to machine. In order to achieve acceptable performance of safety critical components, optimized tool material and cutting parameters are needed. In this regard residual stress is an influential factor determining fatigue life and surface integrity. As a result, process optimization needs to consider the dependency of surface and near surface residual stress upon cutting and tooling conditions. To date, numerous studies have been conducted to establish the relationship between the tooling condition and residual stresses. Henriksen (1951) in 1950s made the earliest attempt to examine the effect of varying rake angles on residual stress and found that residual stresses decreased with increasing rake angle. Liu and Barash (1976a,b) carried out a notable study in 1970s on the mechanical state of the machined layer affected by tool wear. The authors attempted to describe the deformation in the machined layer by
∗ Corresponding author. Tel.: +44 161 306 8872; fax: +44 161 306 5066. E-mail addresses: [email protected], [email protected] (P.J. Withers). 0924-0136/$ – see front matter © 2009 Elsevier B.V. All rights reserved. doi:10.1016/j.jmatprotec.2009.01.012
considering stress–strain curves and the associated impact of thermal effects due to the tool wear. Later, El-Wardany et al. (2000) used finite element models to identify the separate effects of the thermal and mechanical loads suggesting that the thermal load promotes tensile stress whereas the mechanical load suppresses it. Ordas et al. (2003) also carried out a research on the effect of tool wear on the subsurface residual stress profiles. They used the diffraction peak full width at half maximum (FWHM) to identify the significant microstructural change in the machined layer due to the high temperature caused by the worn tool. Other tool geometry parameters have also been of interest. Arunachalam et al. (2004) established that negative rake angle tends to generate compressive residual stress whereas positive rake angle tends to induce tensile residual stress. Creating a chamfer on the cutting edge has a similar effect as creating a negative rake angle, by which the maximum compressive residual stress in the subsurface can be increased. Interestingly, several research groups including Jang et al. (1996), Jacobus et al. (2000) and Hua et al. (2005) found increasing edge radius also has a positive effect on the maximum compressive residual stress. Sasahara et al. (2004) further revealed that geometrically, increasing edge radius to a point that the edge roundness is larger than the uncut chip thickness actually decreases the effective rake angle towards the negative value. Hence a large edge radius has a similar effect to a negative rake angle. Ozel and Zeren (2006) studied the effect of a large radius via FE modelling. The FEM results indicate that a large edge radius is associated with a deep temperature and stress field compared with small edge radius. Moreover, Ordas et al. (2003) and Sharman et al. (2006) reported that increasing flank
W. Li et al. / Journal of Materials Processing Technology 209 (2009) 4896–4902
wear also alters the cutting edge geometry and increases the tensile residual stress at the surface as well as the compressive residual stress in the depth. Recently, several studies on machining residual stress of nickelbased superalloy have appeared in the literature. Axinte et al. (2006) reported that the tensile residual stress at the machined surface of a nickel disc alloy RR1000 tended to increase with the spiral length that the tool passes. When the alloy was machined with a chipped tool a drastic metallurgical alteration (white etching layer) occurred near the surface Li et al. (2006) presented a method to determine the distribution of plastic work using calibrated diffraction peak width and also reported that that the characteristic residual stress depth profiles are different in the two principal in-plane directions for the machined RR1000. Sharman et al. (2006) examined the effect of using coated and uncoated tungsten carbide (WC) tools at various cutting speeds, feeds and levels of tool wear for turning of Inconel 718. It was found that a worn tool could cause a significant increase in both the levels and the depth of the stressed layer, while the use of uncoated WC tools can lead to greater tensile residual stress at the surface than for the coated tools. The present study aims to evaluate the effect of varying tooling conditions on the residual stress distribution in the machined layer for the new generation nickel-base turbine disc alloy RR1000. The targeted tooling conditions include insert type, tool wear, tool coating and tool breakage. Here we focus on the face finish turning as the final surface quality and surface state are largely determined by the finishing operation. 2. Experimental method In this work the X-ray diffraction technique was chosen for the residual stress measurement because of its accuracy and its ability to concurrently acquire the information about the changes in plasticity. In order to evaluate the residual stress distribution with the depth and the variation with the surface locations, the X-ray diffraction stress measurements were carried out in the substrates as well as the surfaces of the samples. 2.1. Materials and specimens RR1000 is a nickel-base superalloy manufactured via powder metallurgy route having a chemical composition of 15% Cr, 14–19% Co, 5% Mo, 3% Al, 4% Ti, 2% Ta, 0.06% Zr, 0.0012–0.033% C, 0.01–0.025% B and Ni balance. The billet used for this study was solution treated at 1120 ◦ C for 4 h, removed from the furnace and then oil quenched. The material was then aged at 760 ◦ C for 8 h. Finish face turning trials were carried out at the Rolls-Royce UTC in University of Nottingham. Real-geometry turbine disc forgings were also provide by Rolls-Royce Plc for the turning trials carried out at the London South Bank University and University of Nottingham in order to assess the effect of tool wear and tool coating, respectively. The forgings had undergone a heat-treatment route slightly different from the billet material: the cooling after the solution treatment was fan air quench instead of oil quench. After the surface integrity assessment all the samples were passed to the University of Manchester for the residual stress study. Three types of samples were produced. 2.1.1. Disc specimens The disc specimens (thickness = 5 mm, diameter = 120 mm ) were produced from the RR1000 billet by the Rolls-Royce UTC, Nottingham University. The principal motion and cutting directions (axial, hoop and radial) are defined in Fig. 1. In each pass the disc surface was cut along the radius from the outer diameter proceeding towards the inner diameter. Two types of coated carbide inserts (round and rhombic) were studied in conjunction with two differ-
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Fig. 1. The geometrical setup of surface finish turning on disc specimen.
ent tool holders to assess the effects of the insert type (Fig. 2). The conditions for the turning trials are summarized in Table 1. 2.1.2. Ring specimen Two 300 mm diameter ring samples made of the RR1000 disc forgings were face finish turned using four different machining conditions at London South Bank University before radial sections were cut from them for examination (Fig. 3). Four surface areas produced on two rings were designated A, B, C and D to test various cutting parameters. As shown in Fig. 3 surfaces A and D were turned using an uncoated rhombic insert, whereas C and B were face turned using a similar rhombic insert having multilayer coating (see Table 1 for cutting conditions). 2.1.3. Washer specimen The washer was produced from the RR1000 disc forgings by the Rolls-Royce UTC at Nottingham University (Fig. 4). The gauge surface on one side of the washer was machined by a relatively new tool which experienced a total cutting time of 480 s, while on the other side it was machined by a relatively worn tool which experienced a total cutting time of 840 s. Residual stresses were measured at the middle of remaining gauge. The cutting conditions are given in Table 1. 2.2. X-ray diffraction The principles of X-ray diffraction residual stress measurement can be found in Cullity and Stock (2001). In this study residual stresses were measured on a portable ProtoXRD unit using manganese K␣ radiation ( = 2.1031 Å) at 20 kV 4 mA to acquire the (3 1 1) diffraction peak at a 2 angle of about 152◦ , using a spot size of 1 mm in diameter. Sin2 measurements were performed at eleven ± angles (ˇ angles with ProtoXRD) using a 3◦ oscillation at each ˇ angle. The Young’s modulus and Poisson’s ratio used for the (3 1 1) reflection were 204 GPa and 0.27, respectively. Stresses in the hoop and radial direction were evaluated using the usual assumption that because the penetration depth is limited the residual stress in the direction normal to the free machined surface to be zero. Considering the accumulation of heat which might affect residual stress during machining as a function of radial positions, all stress measurements were carried out at the outer radius of the disc surface, about 3 mm from the edge. The residual stress
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Fig. 2. Inserts and tool holders used for finishing trials.
Table 1 Cutting and tooling parameters for the turning trials carried out under the same feed rate and depth of cut. Trial type
Sample type
Insert type
Insert radius (mm)
Cutting speed (rpm)
Cutting time (s)
Tool Type Ro1 Tool Type Ro2 Tool Type Rh1 Tool Type Rh2 Coating 1 Coating 2 Coating 3 Coating 4 Wear 1 Wear 2 Breakage
Disc Disc Disc Disc Ring Ring Ring Ring Washer Washer Disc
Coated round Coated round Coated rhombic Coated rhombic Coated rhombic Coated rhombic Uncoated rhombic Uncoated rhombic Coated round Coated round Coated round
6 6 0.8 0.8 1.2 1.2 1.2 1.2 6 6 6
7955 7955 7955 7955 3982 2292 3982 2292 6818 6818 9549
840 1932 1008 1344 NA NA NA NA 420 840 1260
vs. depth profiles were measured by removing successive layers of material to a depth in excess of 100 m by electro-polishing. Considering the relatively large thickness of the disc sample (∼5 mm) in comparison with the stressed layer (<100 m), the difference between the redistributed stress due to layer removal and the true stress state was indeed negligible. Therefore no correction was made for layer removal.
The levels of plastic deformation were estimated from the variation in diffraction peak width (Prevey, 2000). Peak width is usually represented by the (in degrees) FWHM. Peak broadening can be affected by low-angle grain boundaries, very small particle or grain size, plastic deformation, compositional inhomogeneity, as well as variations in macro- or micro-lattice strain, etc. In this case the grain size and other broadening factors besides plastic work are taken
Fig. 3. The specimen for evaluation of the tool coating effect on residual stress.
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Fig. 4. Turned washer with gauge surfaces.
to be insignificant across the sample and FWHM is considered to be a good indicator of plastic work. Li et al. (2006) correlated the measured plastic work with FWHM and thus obtained a calibration curve (Fig. 5). The calibration curve can also be used in this study to infer the plastic work distribution in the subsurface layer. 3. Results and discussion 3.1. Comparison of round and rhombic inserts The residual stress depth profiles measured on the disc sample for cutting with round (Ro) and rhombic (Rh) inserts are shown in Fig. 6. It is evident that the subsurface residual stresses differ dramatically between the hoop and radial directions. The hoop stresses
Fig. 5. Diffraction peak width broadening with plastic work: (a) broadening of (3 1 1) peak width with increasing plastic work (equivalent uniaxial tensile true strain); (b) percent plastic work plotted against FWHM.
Fig. 6. Residual stress depth profiles measured on the disc sample for cutting with (a) round (Ro) and (b) rhombic (Rh) inserts. The corresponding cutting parameters are listed in Table 1.
are predominantly tensile decreasing monotonically for all four trials with the depth and magnitude dependent on the condition. For the case of the round inserts, the hoop stresses are largest at the surface being slightly larger (1520 MPa) for the longer 23 pass trail compared to that (1280 MPa) for the shorter 10 pass trial with the affected depth in the region of 60 m. The radial stresses are tensile (∼300 MPa) near surface (<20 m) becoming compressive at larger depths reaching a level of around −400 MPa. By comparison, for the rhombic insert the hoop stresses are more modest both in magnitude and extent. The radial stresses for Rh2 are −650 MPa at a depth of 30 m or so. The profile for the rhombic insert, Rh1, on the other hand shows a much larger compressive stress (about −1000 MPa) extending to a much greater depth (∼180 m). Furthermore in contrast to the others, the surface stress is compressive (∼−150 MPa) in this case. The changes in X-ray diffraction peak FWHM for the tool insert trial specimens are summarized in Fig. 7. By comparing the measured peak widths against those for uniaxially tensile deformed
Fig. 7. FWHM as a function of depth measured for insert trials.
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calibration samples (Li et al., 2006), the surface FWHM levels correspond to around 22% and 30% plastic work for the round and rhombic inserts, respectively. The FWHM values fall to that representative of the parent disc material at a depth similar to the corresponding affected depths in Fig. 6 for residual stress, decreasing rapidly with depth. The plastically worked zone extends to around 60 m depth for the round inserts and 70 m for the rhombic insert Rh2. The FWHM profile for Rh1 on the other hand has a higher value at the surface and varies over a larger depth than the other three profiles. Taken together there is evidence that the round inserts introduce less plasticity. Further, the residual stress and the FWHM data show no consistent variation with time over the short timescales studied here. Rather it appears that the plastic work profile for Rh1 on the other hand has a higher value at the surface and varies over a much greater depth than the other three profiles, and that this is responsible for the increased magnitude and extent of the compressive stress field. It implies that for the rhombic tool machining the work hardening is higher for the smaller number of cuts. This seems to contradict our intuitive understanding that prolonged cutting will increase tool wear which leads to a rubbing/ploughing effect on the work and thus introduce more plastic deformation into the surface layer. However, the relationship between the plastic deformation in the metal and tool wear is not necessarily straightforward. Factors other than tool wear might play a role in changing the level of work hardening. 3.2. Comparison of coated and uncoated tools Two cutting speeds (medium: 3982 rpm and low: 2292 rpm at the same radial position) were used to evaluate the effect of tool coating on the ring specimens and the results are shown in Fig. 8. Comparing the results with those of Fig. 6 for the disc samples it is clear that the stresses near surface are smaller and the previously observed compressive subsurface peak hardly evident. As before the hoop stresses are tensile, in these cases decreasing rapidly to
Fig. 9. FWHM depth profiles for cutting with coated and uncoated tools.
zero at a depth around of 30 m indicative of a less aggressive machining arrangement than for the disc samples. The medium and low speed samples all show similar profiles for both the coated and uncoated conditions except for the uncoated sample at the medium speed which does show slightly larger stresses. The radial stresses show significantly lower tensile stresses near surface than for the hoop direction in accordance with the disc samples. As for the hoop measurements only the face turned with the uncoated tool at medium speed shows any notable difference both in terms of tensile magnitude and depth (∼40 m). The observed higher surface stresses associated with the uncoated tool at 50 m/min seem to contradict the results reported by Sharman et al in machining Inconel 718 (Sharman et al., 2006) that the coating gives rise to higher tensile stresses than for uncoated tools. Given that the Al3 O2 coating is a thermal barrier which prevents heat from dissipating into the tool bulk, greater heat localization and thus high tensile stresses might be expected. Whether such discrepancy lies in the difference between the thermomechanical properties of RR1000 and Inconel 718 or the machining conditions in the two experiments needs to be further explored. As expected the FWHM profiles in Fig. 9 confirm the lower depths and lower levels of plastic work (corresponding to a level around 15%) introduced into the ring samples compared to the disc samples studied above (Fig. 7). This may suggest less aggressive conditions, perhaps due to better tool conditions. Though only slight, the results do corroborate the conclusion that the uncoated tool introduces plastic work and residual stress to higher levels and depths than the other conditions. 3.3. Comparison of new and worn tools Fig. 10 compares the residual stress profiles for cutting with new and worn tools on the washer samples. In common with all
Fig. 8. Residual stresses measured in tool coating trials on ring specimens: (a) hoop stresses and (b) radial stresses.
Fig. 10. Residual stress depth profiles from worn and new tool trials.
W. Li et al. / Journal of Materials Processing Technology 209 (2009) 4896–4902
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Fig. 11. FWHM depth profiles for cutting with new and worn tools.
the previous results the hoop stresses are more tensile than the radial although in this case there is some evidence of a compressive trough not just for the radial stress but also for the axial direction. Unsurprisingly, the worn tool has introduced stresses that are larger (more tensile near surface) with the compressive maxima at larger depths for both radial and hoop curves. From the FWHM data (Fig. 11) it is evident that the levels of plastic work introduced near surface (∼22%) and affected depths (∼60 m) are rather similar in the two cases. Larger tensile stresses are thought to be related to an increased thermal effect due to the rubbing between the worn tool flank and the work. Moreover, Sharman et al. (2006) also reported an increase in compressive stress underneath the tensile stress layer, associated with a measured significant increase in work hardening when turning with worn tools. However, the FWHM measurements in this study do not point to any notable increase of work hardening subsurface (see Fig. 11). As the prior cutting time for the worn tool was 640 s, only a third of the full tool life (1930s), the tool wear was probably not sufficiently severe to cause a perceivable “ploughing”. Therefore no greater plastic deformation was introduced into the subsurface layer by the worn tool. This may explain why no increase in compressive stress was observed in the subsurface with the worn tool. 3.4. Chipped tool On one occasion, tool breakage occurred in the first few passes during turning a disc specimen at 60 m/min cutting speed and 0.15 mm/rev feed. Uncharacteristically high tensile stress was found at the surface in the hoop direction, reaching a peak value of 1300 MPa 7 m beneath the surface and decreasing to zero around 100 m below the surface (see Fig. 12a). In the radial direction the surface stress was found to be moderately tensile at 240 MPa, below which a significantly compressive stress was found extending from 20 m to a depth in excess of 350 m. This is much deeper and larger in magnitude than observed for all the faces cut using conventional tool conditions. The pronounced anisotropy observed in the hoop and radial directions highlights the dominant role that the mechanical effect plays. Under normal condition for metal cutting, the plastic deformation is thought to take place primarily in three regions: the plastic shear in between the cut and uncut chip, the compressive deformation ahead of the cutting edge and the tensile deformation behind the passing edge. The generated residual stress is expected to be a result of the contest between the compressive and tensile strains (Hua et al., 2005). As such, a chipped edge can lead to a drastic change in the strain field in the vicinity of the cutting edge. Axinte et al. (2006) considered that when a tool chip occurs the plastic deformation caused by a severe “material drag” effect might prevail
Fig. 12. (a) Residual stress depth profiles measured from turning with chipped tool and (b) the associated full width half maximum diffraction peak variation.
Fig. 13. Measured cutting force in the radial direction (Fy ) as a function of cutting time. Each value was an average of the cutting force in a single pass.
over the shearing of the chip. In such a case the plastic deformation is believed to take place to a much greater depth. This hypothesis is partly supported by the FWHM results obtained in this study. The FWHM profile in Fig. 12b would suggest the work hardened region extends to a depth of nearly 300 m. Furthermore, an anomalous increase of the cutting force in the radial direction was observed in the trial (see Fig. 13), suggesting that the observed large subsurface radial stress resulted from a significant mechanical effect. 4. Conclusions The effects of insert type, tool coating, tool wear and tool breakage on residual stress distribution beneath the turned surface of RR1000 nickel-based superalloy have been investigated. Residual stress depth profiles were obtained using X-ray diffraction. Peak width analysis was used to evaluate the work hardening. The effect of tool damage has also been investigated. The following conclusions can be drawn: 1. Compared with the rhombic insert, the round insert generated a slightly higher tensile stress up to 1500 MPa. Despite the rela-
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2. 3.
4.
5.
W. Li et al. / Journal of Materials Processing Technology 209 (2009) 4896–4902
tively small difference in cutting times (1008 and 1344 s), large difference in stress profile and depth were observed for the rhombic inserts, consistent with a much more worn tool. There is some evidence from the FWHM data that the rhombic insert introduces lower levels of work that affects a smaller depth than the rhombic insert. The uncoated insert did not perform less well than the coated. A moderate increase in tool wear led to a higher tensile surface stresses (by between 200 and 400 MPa) but did not appear to introduce more plastic deformation. The work hardened subsurface layer for most samples was found to have a depth of around 50 m, regardless of tool coating or tool wear, except in the case of the damage tool for which a much larger depth was measured (∼300 m). Turning with a chipped tool introduced a large compressive radial stress field with the maximum value reaching −1000 MPa and the penetration depth 400 m. The tool breakage was found to introduce additional plastic deformation into the subsurface and the work-hardened zone extended to 300 m. This is five to eight times deeper than for the normal cases.
The information obtained in this study would help determine the optimized tool conditions for face finish turning of RR1000. The beneficial compressive residual stress can be promoted in the turning process using coated tool and new tool. Excessive tool wear must be avoided because it may introduce high level of tensile stress in both hoop and radial directions. The plastically deformed depth of the machined components was found to be less than 100 m irrespective of the machining conditions (except for tool breakage). This is important for the application of post-machining shot-peening process. Peening can overwhelm the influence of the machining residual stress and thus render uniformly distributed compressive stress in a machined surface, which will improve the fatigue life of the component. Acknowledgements The authors would like to thank London South Bank University for provision of specimens. WL is grateful to Rolls-Royce plc for financial support.
References Arunachalam, R.M., Mannan, M.A., Spowage, A.C., 2004. Surface integrity when machining age hardened Inconel 718 with coated carbide cutting tools. International Journal of Machine Tools & Manufacture 44, 1481–1491. Axinte, D.A., Andrews, P., Li, W., Gindy, N., Withers, P.J., 2006. Turning of advanced Ni based alloys obtained via powder metallurgy route. In: Proceedings of CIRP 56, Kobe. Cullity, B.D., Stock, S.R., 2001. Elements of X-ray Diffraction. Prentice-Hall, Inc., New York. El-Wardany, T.I., Kishawy, H.A., Elbestawi, M.A., 2000. Surface integrity of die material in high speed machining. Part 2. Microhardness variations and residual stress. Journal of Manufacturing Science & Engineering 122, 632–641. Henriksen, E.K., 1951. Residual stresses in machined surfaces. Journal of Applied Mechanics 73, 69–76. Hua, J., Shivpuri, R., Cheng, X., Bedekar, V., Matsumoto, Y., Hashimoto, F., Watkins, T.R., 2005. Effect of feed rate, workpiece hardness and cutting edge on subsurface residual stress in the hard turning of bearing steel using chamfer+hone cutting edge geometry. Materials Science & Engineering 394, 238–248. Jacobus, K., Devor, R.E., Kapoor, S.G., 2000. Machining-induced residual stress; experimentation and modeling. Transaction of the ASME 122, 20–31. Jang, D.Y., Watkins, T.R., Kozaczek, K.J., Hubbard, C.R., 1996. Surface residual stress in machined austenitic stainless steel. Wear 194, 168–173. Li, W., Preuss, M., Withers, P.J., Axinte, D., Andrews, P., 2006. Characterisation of residual stresses in machined surfaces of a high strength nickel-base superalloy. Materials Science Forum 524–525, 587–592. Liu, C.R., Barash, M.M., 1976a. The mechanical state of the sublayer of a surface generated by chip-removal process. Part 1. Cutting with a sharp tool. Transactions of the ASME, 1192–1201. Liu, C.R., Barash, M.M., 1976b. The mechanical state of the sublayer of a surface generated by chip-removal process. Part 2. Cutting with a tool with flank wear. Transactions of the ASME, 1202–1208. Ordas, N., Penalva, M.L., Fernandez, J., Garcia-Rosales, C., 2003. Residual stresses in tool steel due to hard-turning. Journal of Applied Crystallography 36, 1135–1143. Ozel, T., Zeren, E., 2006. Finite element modeling the influence of edge roundness on the stress and temperature fields induced by high-speed machining. International Journal of Advanced Manufacturing Technology Published online. Prevey, P.S., 2000. The effect of cold work on the thermal stability of residual compression in surface enhanced IN718. In: Proceedings of the 20th ASM Materials Solutions conference & Exposition, St. Louis, MO. Sasahara, H., Obikawa, T., Shirakashi, T., 2004. Prediction model of surface residual stress within a machined surface by combining two orthogonal plane models. International Journal of Machine Tools & Manufacture 44, 815–822. Sharman, A.R.C., Hughes, J.I., Ridgeway, K., 2006. An analysis of the residual stresses generated in Inconel 718 when turning. Journal of Materials Processing Technology 173, 359–367.
Contents lists available at ScienceDirect
Journal of Materials Processing Technology journal homepage: www.elsevier.com/locate/jmatprotec
Residual stresses in face finish turning of high strength nickel-based superalloy W. Li a , P.J. Withers a,∗ , D. Axinte b , M. Preuss a , P. Andrews c a
School of Materials, University of Manchester, Grosvenor St M1 7HS, UK School of Mechanical, Materials and Manufacturing Engineering, The University of Nottingham, University Park, Nottingham NG7 2RD, UK c Rolls-Royce plc PO Box 31, Derby DE24 8BJ, UK b
a r t i c l e
i n f o
Article history: Received 16 July 2008 Received in revised form 22 December 2008 Accepted 15 January 2009 Keywords: Superalloy X-ray diffraction Plastic work Tooling condition Turning
a b s t r a c t This paper investigates the residual stresses distributions introduced in a new generation nickel-based superalloy RR1000 by surface finish turning. The residual stresses introduced as a function of depth have been analysed for a series of machining trials with round and rhombic inserts, coated and uncoated inserts, new and worn tools, and chipped tool. The residual stress depth profiles obtained by X-ray diffraction, and layer removal show that the tool type, tool coating, tool wear and tool breakage influence the residual stress. The extent of plastic deformation for different cutting conditions has been inferred from the different peak width. Overall, residual stresses tend to have a tensile character at all depths in the hoop direction, but exhibit a significant compressively stressed zone in the radial direction. © 2009 Elsevier B.V. All rights reserved.
1. Introduction Recent metallurgical developments in the gas turbine industry have been directed at a new generation of nickel-based superalloys to meet the demand of higher operating temperatures. These contain even higher volume fractions of strengthening ␥ precipitates and refractory elements than conventional superalloys, thereby achieving higher strength at high temperatures. This combined with low thermal conductivity makes them difficult to machine. In order to achieve acceptable performance of safety critical components, optimized tool material and cutting parameters are needed. In this regard residual stress is an influential factor determining fatigue life and surface integrity. As a result, process optimization needs to consider the dependency of surface and near surface residual stress upon cutting and tooling conditions. To date, numerous studies have been conducted to establish the relationship between the tooling condition and residual stresses. Henriksen (1951) in 1950s made the earliest attempt to examine the effect of varying rake angles on residual stress and found that residual stresses decreased with increasing rake angle. Liu and Barash (1976a,b) carried out a notable study in 1970s on the mechanical state of the machined layer affected by tool wear. The authors attempted to describe the deformation in the machined layer by
∗ Corresponding author. Tel.: +44 161 306 8872; fax: +44 161 306 5066. E-mail addresses: [email protected], [email protected] (P.J. Withers). 0924-0136/$ – see front matter © 2009 Elsevier B.V. All rights reserved. doi:10.1016/j.jmatprotec.2009.01.012
considering stress–strain curves and the associated impact of thermal effects due to the tool wear. Later, El-Wardany et al. (2000) used finite element models to identify the separate effects of the thermal and mechanical loads suggesting that the thermal load promotes tensile stress whereas the mechanical load suppresses it. Ordas et al. (2003) also carried out a research on the effect of tool wear on the subsurface residual stress profiles. They used the diffraction peak full width at half maximum (FWHM) to identify the significant microstructural change in the machined layer due to the high temperature caused by the worn tool. Other tool geometry parameters have also been of interest. Arunachalam et al. (2004) established that negative rake angle tends to generate compressive residual stress whereas positive rake angle tends to induce tensile residual stress. Creating a chamfer on the cutting edge has a similar effect as creating a negative rake angle, by which the maximum compressive residual stress in the subsurface can be increased. Interestingly, several research groups including Jang et al. (1996), Jacobus et al. (2000) and Hua et al. (2005) found increasing edge radius also has a positive effect on the maximum compressive residual stress. Sasahara et al. (2004) further revealed that geometrically, increasing edge radius to a point that the edge roundness is larger than the uncut chip thickness actually decreases the effective rake angle towards the negative value. Hence a large edge radius has a similar effect to a negative rake angle. Ozel and Zeren (2006) studied the effect of a large radius via FE modelling. The FEM results indicate that a large edge radius is associated with a deep temperature and stress field compared with small edge radius. Moreover, Ordas et al. (2003) and Sharman et al. (2006) reported that increasing flank
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wear also alters the cutting edge geometry and increases the tensile residual stress at the surface as well as the compressive residual stress in the depth. Recently, several studies on machining residual stress of nickelbased superalloy have appeared in the literature. Axinte et al. (2006) reported that the tensile residual stress at the machined surface of a nickel disc alloy RR1000 tended to increase with the spiral length that the tool passes. When the alloy was machined with a chipped tool a drastic metallurgical alteration (white etching layer) occurred near the surface Li et al. (2006) presented a method to determine the distribution of plastic work using calibrated diffraction peak width and also reported that that the characteristic residual stress depth profiles are different in the two principal in-plane directions for the machined RR1000. Sharman et al. (2006) examined the effect of using coated and uncoated tungsten carbide (WC) tools at various cutting speeds, feeds and levels of tool wear for turning of Inconel 718. It was found that a worn tool could cause a significant increase in both the levels and the depth of the stressed layer, while the use of uncoated WC tools can lead to greater tensile residual stress at the surface than for the coated tools. The present study aims to evaluate the effect of varying tooling conditions on the residual stress distribution in the machined layer for the new generation nickel-base turbine disc alloy RR1000. The targeted tooling conditions include insert type, tool wear, tool coating and tool breakage. Here we focus on the face finish turning as the final surface quality and surface state are largely determined by the finishing operation. 2. Experimental method In this work the X-ray diffraction technique was chosen for the residual stress measurement because of its accuracy and its ability to concurrently acquire the information about the changes in plasticity. In order to evaluate the residual stress distribution with the depth and the variation with the surface locations, the X-ray diffraction stress measurements were carried out in the substrates as well as the surfaces of the samples. 2.1. Materials and specimens RR1000 is a nickel-base superalloy manufactured via powder metallurgy route having a chemical composition of 15% Cr, 14–19% Co, 5% Mo, 3% Al, 4% Ti, 2% Ta, 0.06% Zr, 0.0012–0.033% C, 0.01–0.025% B and Ni balance. The billet used for this study was solution treated at 1120 ◦ C for 4 h, removed from the furnace and then oil quenched. The material was then aged at 760 ◦ C for 8 h. Finish face turning trials were carried out at the Rolls-Royce UTC in University of Nottingham. Real-geometry turbine disc forgings were also provide by Rolls-Royce Plc for the turning trials carried out at the London South Bank University and University of Nottingham in order to assess the effect of tool wear and tool coating, respectively. The forgings had undergone a heat-treatment route slightly different from the billet material: the cooling after the solution treatment was fan air quench instead of oil quench. After the surface integrity assessment all the samples were passed to the University of Manchester for the residual stress study. Three types of samples were produced. 2.1.1. Disc specimens The disc specimens (thickness = 5 mm, diameter = 120 mm ) were produced from the RR1000 billet by the Rolls-Royce UTC, Nottingham University. The principal motion and cutting directions (axial, hoop and radial) are defined in Fig. 1. In each pass the disc surface was cut along the radius from the outer diameter proceeding towards the inner diameter. Two types of coated carbide inserts (round and rhombic) were studied in conjunction with two differ-
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Fig. 1. The geometrical setup of surface finish turning on disc specimen.
ent tool holders to assess the effects of the insert type (Fig. 2). The conditions for the turning trials are summarized in Table 1. 2.1.2. Ring specimen Two 300 mm diameter ring samples made of the RR1000 disc forgings were face finish turned using four different machining conditions at London South Bank University before radial sections were cut from them for examination (Fig. 3). Four surface areas produced on two rings were designated A, B, C and D to test various cutting parameters. As shown in Fig. 3 surfaces A and D were turned using an uncoated rhombic insert, whereas C and B were face turned using a similar rhombic insert having multilayer coating (see Table 1 for cutting conditions). 2.1.3. Washer specimen The washer was produced from the RR1000 disc forgings by the Rolls-Royce UTC at Nottingham University (Fig. 4). The gauge surface on one side of the washer was machined by a relatively new tool which experienced a total cutting time of 480 s, while on the other side it was machined by a relatively worn tool which experienced a total cutting time of 840 s. Residual stresses were measured at the middle of remaining gauge. The cutting conditions are given in Table 1. 2.2. X-ray diffraction The principles of X-ray diffraction residual stress measurement can be found in Cullity and Stock (2001). In this study residual stresses were measured on a portable ProtoXRD unit using manganese K␣ radiation ( = 2.1031 Å) at 20 kV 4 mA to acquire the (3 1 1) diffraction peak at a 2 angle of about 152◦ , using a spot size of 1 mm in diameter. Sin2 measurements were performed at eleven ± angles (ˇ angles with ProtoXRD) using a 3◦ oscillation at each ˇ angle. The Young’s modulus and Poisson’s ratio used for the (3 1 1) reflection were 204 GPa and 0.27, respectively. Stresses in the hoop and radial direction were evaluated using the usual assumption that because the penetration depth is limited the residual stress in the direction normal to the free machined surface to be zero. Considering the accumulation of heat which might affect residual stress during machining as a function of radial positions, all stress measurements were carried out at the outer radius of the disc surface, about 3 mm from the edge. The residual stress
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Fig. 2. Inserts and tool holders used for finishing trials.
Table 1 Cutting and tooling parameters for the turning trials carried out under the same feed rate and depth of cut. Trial type
Sample type
Insert type
Insert radius (mm)
Cutting speed (rpm)
Cutting time (s)
Tool Type Ro1 Tool Type Ro2 Tool Type Rh1 Tool Type Rh2 Coating 1 Coating 2 Coating 3 Coating 4 Wear 1 Wear 2 Breakage
Disc Disc Disc Disc Ring Ring Ring Ring Washer Washer Disc
Coated round Coated round Coated rhombic Coated rhombic Coated rhombic Coated rhombic Uncoated rhombic Uncoated rhombic Coated round Coated round Coated round
6 6 0.8 0.8 1.2 1.2 1.2 1.2 6 6 6
7955 7955 7955 7955 3982 2292 3982 2292 6818 6818 9549
840 1932 1008 1344 NA NA NA NA 420 840 1260
vs. depth profiles were measured by removing successive layers of material to a depth in excess of 100 m by electro-polishing. Considering the relatively large thickness of the disc sample (∼5 mm) in comparison with the stressed layer (<100 m), the difference between the redistributed stress due to layer removal and the true stress state was indeed negligible. Therefore no correction was made for layer removal.
The levels of plastic deformation were estimated from the variation in diffraction peak width (Prevey, 2000). Peak width is usually represented by the (in degrees) FWHM. Peak broadening can be affected by low-angle grain boundaries, very small particle or grain size, plastic deformation, compositional inhomogeneity, as well as variations in macro- or micro-lattice strain, etc. In this case the grain size and other broadening factors besides plastic work are taken
Fig. 3. The specimen for evaluation of the tool coating effect on residual stress.
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Fig. 4. Turned washer with gauge surfaces.
to be insignificant across the sample and FWHM is considered to be a good indicator of plastic work. Li et al. (2006) correlated the measured plastic work with FWHM and thus obtained a calibration curve (Fig. 5). The calibration curve can also be used in this study to infer the plastic work distribution in the subsurface layer. 3. Results and discussion 3.1. Comparison of round and rhombic inserts The residual stress depth profiles measured on the disc sample for cutting with round (Ro) and rhombic (Rh) inserts are shown in Fig. 6. It is evident that the subsurface residual stresses differ dramatically between the hoop and radial directions. The hoop stresses
Fig. 5. Diffraction peak width broadening with plastic work: (a) broadening of (3 1 1) peak width with increasing plastic work (equivalent uniaxial tensile true strain); (b) percent plastic work plotted against FWHM.
Fig. 6. Residual stress depth profiles measured on the disc sample for cutting with (a) round (Ro) and (b) rhombic (Rh) inserts. The corresponding cutting parameters are listed in Table 1.
are predominantly tensile decreasing monotonically for all four trials with the depth and magnitude dependent on the condition. For the case of the round inserts, the hoop stresses are largest at the surface being slightly larger (1520 MPa) for the longer 23 pass trail compared to that (1280 MPa) for the shorter 10 pass trial with the affected depth in the region of 60 m. The radial stresses are tensile (∼300 MPa) near surface (<20 m) becoming compressive at larger depths reaching a level of around −400 MPa. By comparison, for the rhombic insert the hoop stresses are more modest both in magnitude and extent. The radial stresses for Rh2 are −650 MPa at a depth of 30 m or so. The profile for the rhombic insert, Rh1, on the other hand shows a much larger compressive stress (about −1000 MPa) extending to a much greater depth (∼180 m). Furthermore in contrast to the others, the surface stress is compressive (∼−150 MPa) in this case. The changes in X-ray diffraction peak FWHM for the tool insert trial specimens are summarized in Fig. 7. By comparing the measured peak widths against those for uniaxially tensile deformed
Fig. 7. FWHM as a function of depth measured for insert trials.
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calibration samples (Li et al., 2006), the surface FWHM levels correspond to around 22% and 30% plastic work for the round and rhombic inserts, respectively. The FWHM values fall to that representative of the parent disc material at a depth similar to the corresponding affected depths in Fig. 6 for residual stress, decreasing rapidly with depth. The plastically worked zone extends to around 60 m depth for the round inserts and 70 m for the rhombic insert Rh2. The FWHM profile for Rh1 on the other hand has a higher value at the surface and varies over a larger depth than the other three profiles. Taken together there is evidence that the round inserts introduce less plasticity. Further, the residual stress and the FWHM data show no consistent variation with time over the short timescales studied here. Rather it appears that the plastic work profile for Rh1 on the other hand has a higher value at the surface and varies over a much greater depth than the other three profiles, and that this is responsible for the increased magnitude and extent of the compressive stress field. It implies that for the rhombic tool machining the work hardening is higher for the smaller number of cuts. This seems to contradict our intuitive understanding that prolonged cutting will increase tool wear which leads to a rubbing/ploughing effect on the work and thus introduce more plastic deformation into the surface layer. However, the relationship between the plastic deformation in the metal and tool wear is not necessarily straightforward. Factors other than tool wear might play a role in changing the level of work hardening. 3.2. Comparison of coated and uncoated tools Two cutting speeds (medium: 3982 rpm and low: 2292 rpm at the same radial position) were used to evaluate the effect of tool coating on the ring specimens and the results are shown in Fig. 8. Comparing the results with those of Fig. 6 for the disc samples it is clear that the stresses near surface are smaller and the previously observed compressive subsurface peak hardly evident. As before the hoop stresses are tensile, in these cases decreasing rapidly to
Fig. 9. FWHM depth profiles for cutting with coated and uncoated tools.
zero at a depth around of 30 m indicative of a less aggressive machining arrangement than for the disc samples. The medium and low speed samples all show similar profiles for both the coated and uncoated conditions except for the uncoated sample at the medium speed which does show slightly larger stresses. The radial stresses show significantly lower tensile stresses near surface than for the hoop direction in accordance with the disc samples. As for the hoop measurements only the face turned with the uncoated tool at medium speed shows any notable difference both in terms of tensile magnitude and depth (∼40 m). The observed higher surface stresses associated with the uncoated tool at 50 m/min seem to contradict the results reported by Sharman et al in machining Inconel 718 (Sharman et al., 2006) that the coating gives rise to higher tensile stresses than for uncoated tools. Given that the Al3 O2 coating is a thermal barrier which prevents heat from dissipating into the tool bulk, greater heat localization and thus high tensile stresses might be expected. Whether such discrepancy lies in the difference between the thermomechanical properties of RR1000 and Inconel 718 or the machining conditions in the two experiments needs to be further explored. As expected the FWHM profiles in Fig. 9 confirm the lower depths and lower levels of plastic work (corresponding to a level around 15%) introduced into the ring samples compared to the disc samples studied above (Fig. 7). This may suggest less aggressive conditions, perhaps due to better tool conditions. Though only slight, the results do corroborate the conclusion that the uncoated tool introduces plastic work and residual stress to higher levels and depths than the other conditions. 3.3. Comparison of new and worn tools Fig. 10 compares the residual stress profiles for cutting with new and worn tools on the washer samples. In common with all
Fig. 8. Residual stresses measured in tool coating trials on ring specimens: (a) hoop stresses and (b) radial stresses.
Fig. 10. Residual stress depth profiles from worn and new tool trials.
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Fig. 11. FWHM depth profiles for cutting with new and worn tools.
the previous results the hoop stresses are more tensile than the radial although in this case there is some evidence of a compressive trough not just for the radial stress but also for the axial direction. Unsurprisingly, the worn tool has introduced stresses that are larger (more tensile near surface) with the compressive maxima at larger depths for both radial and hoop curves. From the FWHM data (Fig. 11) it is evident that the levels of plastic work introduced near surface (∼22%) and affected depths (∼60 m) are rather similar in the two cases. Larger tensile stresses are thought to be related to an increased thermal effect due to the rubbing between the worn tool flank and the work. Moreover, Sharman et al. (2006) also reported an increase in compressive stress underneath the tensile stress layer, associated with a measured significant increase in work hardening when turning with worn tools. However, the FWHM measurements in this study do not point to any notable increase of work hardening subsurface (see Fig. 11). As the prior cutting time for the worn tool was 640 s, only a third of the full tool life (1930s), the tool wear was probably not sufficiently severe to cause a perceivable “ploughing”. Therefore no greater plastic deformation was introduced into the subsurface layer by the worn tool. This may explain why no increase in compressive stress was observed in the subsurface with the worn tool. 3.4. Chipped tool On one occasion, tool breakage occurred in the first few passes during turning a disc specimen at 60 m/min cutting speed and 0.15 mm/rev feed. Uncharacteristically high tensile stress was found at the surface in the hoop direction, reaching a peak value of 1300 MPa 7 m beneath the surface and decreasing to zero around 100 m below the surface (see Fig. 12a). In the radial direction the surface stress was found to be moderately tensile at 240 MPa, below which a significantly compressive stress was found extending from 20 m to a depth in excess of 350 m. This is much deeper and larger in magnitude than observed for all the faces cut using conventional tool conditions. The pronounced anisotropy observed in the hoop and radial directions highlights the dominant role that the mechanical effect plays. Under normal condition for metal cutting, the plastic deformation is thought to take place primarily in three regions: the plastic shear in between the cut and uncut chip, the compressive deformation ahead of the cutting edge and the tensile deformation behind the passing edge. The generated residual stress is expected to be a result of the contest between the compressive and tensile strains (Hua et al., 2005). As such, a chipped edge can lead to a drastic change in the strain field in the vicinity of the cutting edge. Axinte et al. (2006) considered that when a tool chip occurs the plastic deformation caused by a severe “material drag” effect might prevail
Fig. 12. (a) Residual stress depth profiles measured from turning with chipped tool and (b) the associated full width half maximum diffraction peak variation.
Fig. 13. Measured cutting force in the radial direction (Fy ) as a function of cutting time. Each value was an average of the cutting force in a single pass.
over the shearing of the chip. In such a case the plastic deformation is believed to take place to a much greater depth. This hypothesis is partly supported by the FWHM results obtained in this study. The FWHM profile in Fig. 12b would suggest the work hardened region extends to a depth of nearly 300 m. Furthermore, an anomalous increase of the cutting force in the radial direction was observed in the trial (see Fig. 13), suggesting that the observed large subsurface radial stress resulted from a significant mechanical effect. 4. Conclusions The effects of insert type, tool coating, tool wear and tool breakage on residual stress distribution beneath the turned surface of RR1000 nickel-based superalloy have been investigated. Residual stress depth profiles were obtained using X-ray diffraction. Peak width analysis was used to evaluate the work hardening. The effect of tool damage has also been investigated. The following conclusions can be drawn: 1. Compared with the rhombic insert, the round insert generated a slightly higher tensile stress up to 1500 MPa. Despite the rela-
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tively small difference in cutting times (1008 and 1344 s), large difference in stress profile and depth were observed for the rhombic inserts, consistent with a much more worn tool. There is some evidence from the FWHM data that the rhombic insert introduces lower levels of work that affects a smaller depth than the rhombic insert. The uncoated insert did not perform less well than the coated. A moderate increase in tool wear led to a higher tensile surface stresses (by between 200 and 400 MPa) but did not appear to introduce more plastic deformation. The work hardened subsurface layer for most samples was found to have a depth of around 50 m, regardless of tool coating or tool wear, except in the case of the damage tool for which a much larger depth was measured (∼300 m). Turning with a chipped tool introduced a large compressive radial stress field with the maximum value reaching −1000 MPa and the penetration depth 400 m. The tool breakage was found to introduce additional plastic deformation into the subsurface and the work-hardened zone extended to 300 m. This is five to eight times deeper than for the normal cases.
The information obtained in this study would help determine the optimized tool conditions for face finish turning of RR1000. The beneficial compressive residual stress can be promoted in the turning process using coated tool and new tool. Excessive tool wear must be avoided because it may introduce high level of tensile stress in both hoop and radial directions. The plastically deformed depth of the machined components was found to be less than 100 m irrespective of the machining conditions (except for tool breakage). This is important for the application of post-machining shot-peening process. Peening can overwhelm the influence of the machining residual stress and thus render uniformly distributed compressive stress in a machined surface, which will improve the fatigue life of the component. Acknowledgements The authors would like to thank London South Bank University for provision of specimens. WL is grateful to Rolls-Royce plc for financial support.
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