an analysis of the residual stresses generated in inconel 718™ when turning

Journal of Materials Processing Technology 173 (2006) 359–367

An analysis of the residual stresses generated in Inconel718TM when turning

A.R.C. Sharman ∗, J.I. Hughes, K. RidgwayAdvanced Manufacturing Research Centre with Boeing, Department of Mechanical Engineering, University of Sheffield, Sheffield S1 3JD, United Kingdom

Received 7 February 2005; received in revised form 8 December 2005; accepted 8 December 2005

Abstract

Inconel 718 is one of a family of nickel based superalloys that are used extensively by the aerospace industry in the hot sections of gas turbineengines. The literature detailing the effects of varying operating parameters on tool life when machining nickel based superalloys is comprehensive,however, relatively little of this data refers to their effects on machined workpiece surface integrity and residual stress generation. Greater knowledgeof the effects of operating parameters on surface integrity is critical to the acceptance of new cutting tool materials, tool geometries and strategies.The paper initially reviews prior work on the surface integrity achieved when turning Inconel 718. Following on from this a series of experimentsevaluating the effects of varying cutting tool material, geometry, wear level and operating parameters are detailed. The results show that the largestihu©

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nfluence on surface integrity was tool wear. Cutting with a worn tool resulted in greater microstructural deformation, microhardness changes andigh surface tensile stresses. High tensile stresses were also formed in the surface layer when cutting with a coated tool, while cutting with anncoated tungsten carbide insert at the same operating parameters produced deep compressive stresses beneath a reduced tensile layer.

2006 Elsevier B.V. All rights reserved.

eywords: Surface integrity; Residual stress; Nickel based superalloys; Turning

. Introduction

Inconel 718 is a high strength, heat resistant superalloyHRSA) that is used extensively by the aerospace industry forhe hot sections of gas turbine engines for components such as,urbine disks, blades, combustors, casings, etc. It is one of the

ost widely used nickel base superalloys accounting for around5% of all production [1]. The microstructure of Inconel 718s comprised of an austenitic face centred cubic (FCC) matrixhase, which is a solid solution of Fe, Cr and Mo in nickelogether with other secondary phases. The main strengtheninghase is the precipitate gamma double prime (denoted �′′). Thishase consists of uniformly distributed body centred tetragonalBCT) disc-shaped particles (of composition Ni3Nb) that areoherent with the parent matrix. Inconel 718 is often used insolution treated and aged condition, this involves a solution

reatment at 970–1175 ◦C, followed by a precipitation treatmentt 600–815 ◦C [2]. The heat treatment results in a microstructuref large grains containing the �′′ precipitated phase and a heavy

concentration of carbides at the grain boundaries. The difficultyof dislocation motion through this microstructure is responsi-ble for the high tensile and yield strength of the material. Themicrostructure only degrades significantly when held at temper-atures higher than its ageing temperature for extended periods.During long-term exposure to moderate temperatures (∼650 ◦C)the �′′ particles increase in size but coherency is not lost. As thetemperature is further increased strength begins to decrease withtime due to growth of the �′′ particles and consequential loss ofcoherency [2].

The properties that make Inconel 718 an important engineer-ing material are also responsible for its generally poor machin-ability. Low thermal conductivity (11.4 W/m/K) leads to highcutting temperatures being developed in the cutting zone. Thesehave been shown to rise from around 900 ◦C at a relativelylow cutting speed of 30 m/min up to 1300 ◦C at 300 m/min [3].The cutting forces generated are also very high, around doublethat found when cutting medium carbon alloy steels. Literaturedetailing the effects of operating parameters on tool life whenmachining nickel based superalloys is comprehensive, however,relatively little of this data refers to the effects of machining on

∗ Corresponding author.E-mail address: [email protected] (A.R.C. Sharman).

workpiece surface integrity. The main problems reported are sur-face tearing, cavities, cracking, metallurgical recrystalisation,

924-0136/$ – see front matter © 2006 Elsevier B.V. All rights reserved.oi:10.1016/j.jmatprotec.2005.12.007

360 A.R.C. Sharman et al. / Journal of Materials Processing Technology 173 (2006) 359–367

plastic deformation, microhardness increases and the formationof residual stresses [4–10].

When a workpiece is subjected to deformation that is notuniform throughout its section, residual stresses are developed.These can be thought of as locked-in stresses that are not sub-ject to external forces. In machining the mechanical, thermaland metallurgical effects associated with chip formation leadto inhomogeneous plastic deformation of the workpiece. Thematerial directly ahead of the advancing cutting tool experiencescompressive plastic deformation whilst the material behind it isin tension. Additional tensile deformation occurs due to rub-bing from the flank face of the tool. If the amount of tensiledeformation produced is greater than the level of compressivedeformation then compressive residual stresses will be producedand vice verse. The heat generated during chip formation pro-duces compressive plastic deformation of the surface due tolocalised thermal expansion, this results in tensile stresses uponcooling. The interaction of all these factors and the thermo-mechanical properties of the workpiece material being machinedwill determine the final residual stress state. It is stated thatthe thermal factors are the most damaging to the integrity ofthe component, producing high tensile stresses at or near to theworkpiece surface [11].

Studies on machining Inconel 718 have shown that the resid-ual stress profile produced when turning is tensile at the work-piece surface followed by a reduction with increasing depthbiptacavhddto

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is expected that this would correspond to some stress relief. Ingeneral, the use of a cutting fluid was shown to reduce workpiecesurface damage [8].

It has been shown that the use of ceramic tools can lead tothe production of greater levels of tensile residual stress whencompared to that obtained with uncoated tungsten carbide (WC)inserts [14]. This result is most likely due to the high cuttingspeeds utilised with ceramic tools and thus greater tempera-tures developed in the cutting zone. In addition, surface integritywas poor due to a high level of workpiece smearing. Ezugwuand Tang [15] reported high levels of work hardening, plasticdeformation and tearing of the workpiece surface when cuttingwith alumina based ceramic tools. Arunachalam and Mannan [4]stated that finishing operations on flight critical HRSA aerospaceparts are currently conducted using uncoated WC tools at rel-atively low cutting speeds due to concerns about the surfaceintegrity produced with other tool materials and their associatedoperating parameters.

The following experimental work was undertaken to evaluatethe effects of using coated and uncoated WC tools at variouscutting conditions, on the surface integrity and residual stressobtained when turning Inconel 718.

2. Experimental procedure

2.1. Workpiece materials and equipment

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eneath it [6–9,12,13]. Sadat and Reddy [8,9] found that anncrease in cutting speed (from 6 to 60 m/min) reduced work-iece surface damage (in the form of microstructural deforma-ion, surface tearing, etc.) by reducing the cutting forces gener-ted (by ∼400 N) due to an increase in cutting temperature andorresponding drop in workpiece mechanical strength. However,t the higher cutting speed the maximum tensile residual stressalue was increased. Although the use of low cutting speedsas been shown to reduce the level of tensile residual stresseveloped, the workpiece surface suffered from greater plasticeformation, tearing, pull out and microhardness increases dueo the development of higher cutting forces and a built up edgen the tool [6].

Work on tools with a controlled contact length has shownhat the level of tensile residual stress generated was lower dueo a reduction in sliding friction between the tool edge and theorkpiece surface [6]. Similar results have also been seen when

omparing the surfaces produced with new and worn tools [5].hen cutting with new tools little subsurface microstructural

lteration was seen, however after continued cutting the work-iece contained significant plastic deformation and increasedicrohardness. This was attributed to the combination of higher

utting temperatures and forces developed when cutting with aorn tool.Gorsler [7] stated that turning HRSA without the use of a

utting fluid produced a higher and deeper tensile residual stressrofile than was seen when cutting wet due to the higher tem-eratures generated. In contrast, Sadat and Reddy [8,9] foundhat machining dry actually reduced the peak residual stress

easured when compared to machining wet. However, the work-iece surfaces produced in the study were highly cracked and it

The workpiece material used was Inconel 718 with a chemical compositionf 53.8% Ni, 18.1% Cr, 5.5% Nb, 2.9% Mo, 1% Ti, 0.55% Al, 0.25% C, 0.04%i, 0.06% Mn and balance Fe (weight percent). This material was solution treatednd aged to a nominal bulk hardness of 38 HRC.

Sections of the machined workpiece were cut out of the bar using wirelectro-discharge machining (WEDM). These samples were used for microstruc-ural and microhardness evaluation. Sections were hot mounted in Bakelite,round using SiC paper and polished with diamond grit. After polishing theyere immersion etched in Callings No. 2 reagent for around 10 s. Microhard-ess measurements were conducted using a Knoop indenter at a load of 50 gor 15 s. One of the problems associated with microhardness measurementsoncerns its sensitivity to hard particles just below the workpiece surface. Tovercome this a series of measurements were taken and an average obtained.ubsurface microstructural analysis was conducted with a Leica optical micro-cope up to a maximum of 1500× magnification, measurements of the depth oficrostructural deformation were conducted by taking a number of measure-ents at random positions across each section.

Additional specimens were WEDM’ed from the bar for residual stress eval-ation. Residual stress measurements were made using the blind hole drillingechnique by Stresscraft Ltd. Gauge positions were selected at the centre ofhe specimens and were prepared by swab etching with glyceregia and thencetone. The gauges were attached using glue. Drilling was performed using anrbital drilling technique and a three-axis PC controlled drilling stage. A 0.6 mmiameter inverted cone-shaped WC drill was used with an orbit eccentricity of.15 mm. Gauge installation and drilling was conducted in accordance with thePL good practice guide [16].

.2. Experimental procedure

All machining trials were conducted on a Cincinnati Hawk 300 turning centremploying a continuously variable spindle speed up to a maximum of 3000 rpmnd a drive motor rated up to 42 kW. Machining parameters fixed throughouthese trials were depth of cut (0.25 mm) and cutting fluid, which was a 5%olution of semi-synthetic emulsion supplied using the turning centres standardood system at 5 bar/30 l/min. All inserts were held in a Sandvik Coromantapto C5 modular tool holder corresponding to DCLNL configuration. Tests

A.R.C. Sharman et al. / Journal of Materials Processing Technology 173 (2006) 359–367 361

Table 1Test matrix

Tool material Feed rate(mm/rev)

Cutting speed (m/min)

40 80 120

Tool S (TiCN/Al2O3/TiN coated WC) 0.15√ √ √

0.25√ √ √

Tool H (uncoated WC) 0.15√

– –0.25

√– –

NB: symbol (√

) denotes test undertaken.

were carried out to investigate the effect of tool material, cutting speed, feed rateand tool wear on the residual stress and workpiece surface integrity obtained.Sections from the workpiece were taken at positions corresponding to around30 s tool life and from surfaces cut with a tool worn to ∼0.25 mm average flankwear, in this way the effects of both cutting parameters and the level of tool wearcould be examined. Tool wear was measured with a toolmakers microscope fittedwith a digital camera and image analysis software. During machining trials, ISO3685 – specification for tool-life testing with single point cutting tools – 1993,was followed as closely as possible [17]. The tool was classified as worn ifthe average flank wear reached 0.25 mm or the maximum flank wear reached0.5 mm. The variable cutting parameters and test matrix are detailed in Table 1.

Cutting force was measured with a Kistler 9121 three-component piezoelec-tric dynamometer and associated 5070 multichannel charge amplifier connectedto a PC employing Kistler Dynoware force measurement software. Measure-ments were taken within the first 10 s of cut with a new tool and with a worntool. For the cutting force trials a DCLNL shank tool holder was used. Twoinserts were used in this work, Tool H was a K10 grade uncoated WC insertand Tool S had a multilayer TiCN/Al2O3/TiN coating over a K10 grade WCsubstrate. Details of the tool geometries used are given in Fig. 1.

3. Results and discussion

The microstructural damage observed in this work was thesubject of a previous paper therefore only a brief descriptionwill be given here to assist in understanding the residual stressresults, for full details the reader is referred to ref. [10].

The subsurface damage caused by machining consisted ofdeformed grain boundaries in the direction of cutting (see Fig. 2),caaa[os

Fig. 1. Geometries of the tools used.

severe plastic flow in the workpiece surface region as the tooladvances leading to surface tearing. Although an adhering layerof workpiece material was present on the tool, in this work themajority of surface cavities were associated with carbide par-ticles that were directly in the path of the cutting tool edge, asseen in Fig. 3. These hard particles are unable to deform withthe plasticised layer and so crack to relieve strain. They are thenremoved from the surface with the chip, leaving behind a cavity(see Fig. 4). Bailey [20] suggested that the parent matrix aroundhard particles would also crack due to incompatible strain fieldsacross the particle/matrix boundary, which would cause parti-cle delamination, evidence for this can be seen in Fig. 3. Oncethe particle has been removed the tool advances through free airand as it re-enters the cut, the material tears due to the suddenincrease in shear stress, this tearing can be seen in Fig. 4. Ingeneral, the number of surface cavities was minimal and theirposition on the workpiece random, this did not change across thelevel of operating parameters examined in these trials. This pro-vides further evidence that it is a microstructural phenomenon,as the microstructure is expected to be consistent throughout thetrials.

When cutting with a new tool relatively little plastic deforma-tion of the grain boundaries occurred (average 12 �m; see Fig. 6).All the samples were strain hardened in the near surface layer

of In

racked carbide particles (see Fig. 3), surface cavities (see Fig. 4)nd microhardness increases (see Fig. 5). These types of defectre commonly reported for the machining of Inconel 718 and canffect the subsequent mechanical properties of the workpiece5,6,8,9,15,18]. Sadat and Bailey [19] attributed the formationf surface cavities to the presence of a sticking layer on the toolurface, adhesion between the tool and the workpiece causes

Fig. 2. Typical microstructural deformation
conel 718. (a) New tool and (b) worn tool.


Fig. 3. Cracked carbide particles in the deformed layer.

to an average of 440 HK0.05 before dropping to bulk hardness(385 HK0.05) within 50 �m depth (see Fig. 5). In all cases, thesurfaces produced with worn tools were harder (maximum 500HK0.05) with a depth penetration greater (∼200 �m) than thoseproduced with a new tool. Chou [21] considered that the primaryhardening mechanism in machining was the rapid heating andcooling cycle with mechanical deformation being secondary. ForInconel 718 a very high temperature in the near surface regioncould cause overaging leading to the drop in microhardness lev-

Fig. 4. Typical surface tearing and cavities observed (note carbide particles atsurface).

els has often seen when cutting with worn tools (see Fig. 5b andd). It is also expected that the high temperatures would causesome strain recovery to occur. In addition to significant harden-ing, the surfaces produced with worn tools also had higher levelsof grain boundary deformation (up to ∼36 �m). This can beseen when comparing Fig. 2a and b. As the tool wears, its clear-ance angle is reduced leading to an increase in tool/workpiececontact area and thus greater rubbing of the workpiece surface.

Ft

ig. 5. Microhardness depth profiles (bulk hardness is 385 ± 15 HK0.05). (a) New toools at 0.25 mm/rev feed rate and (d) worn tools at 0.25 mm/rev feed rate.

ols at 0.15 mm/rev feed rate, (b) worn tools at 0.15 mm/rev feed rate, (c) new


Fig. 6. Depth of microstructural deformation with operating parameters.

This causes an increase in plastic deformation of the workpiecesurface region and increases the temperature generated by fric-tion. The increase in rubbing is highlighted by the considerablyhigher cutting forces that were developed when cutting witha worn tool. Fig. 7 shows that the radial, axial and tangentialforces all increased when the tool was worn, with the radialforce increasing by up to a factor of 10. Fig. 8 is a plot of aver-age microstructural deformation and maximum microhardnesschange (i.e. the increase over bulk levels) that occurred in eachtest compared against the cutting forces recorded for that test.This graph shows that both depth of deformation and maximummicrohardness have good correlation with cutting force, specif-ically as the resultant force increased so did the level of strainhardening and depth of deformation. In addition, the influenceof cutting forces on surface damage is further highlighted by thefact that higher depths of deformation were seen when cutting

utting

Fig. 8. Effect of cutting force on depth of deformation and microhardness.

at the higher feed rate which also produced higher forces (seeFig. 7).

A number of authors have reported that when cutting speedis increased the cutting forces drop due to a reduction in work-piece mechanical properties in the shear zone caused by thehigher temperatures generated [8,9,22]. No such drop in forcewas seen in these experiments. Also it has been stated that whenhigher cutting speeds are used the near surface microhardnesslevel decreases due to thermal softening [5]. Again, in these tri-als there did not appear to be any noticeable difference betweenthe microhardness profile or the depth of plastic deformation atthe various cutting speeds examined when new tools were used.It is likely that within the range of cutting speeds investigated,

Fig. 7. C
forces.


the increasing cutting temperature was balanced by the increasein strain rates. In contrast to the results seen with new tools,there was a large increase in both the level of subsurface hard-ness and the depth of microstructural deformation produced at80 and 120 m/min cutting speed when a worn tool was used (seeFigs. 5b and d and 6). It is suggested that the higher strain ratesencountered with higher cutting speeds along with the highercutting forces developed when cutting with a worn tool areresponsible for this result. Indeed very high radial cutting forceswere developed when cutting under those operating conditions(see Fig. 7).

In all cases when cutting with an unworn tool the residualstress profile was tensile at the surface with values ranging from14 to 747 MPa depending upon the cutting parameters employed.With increasing depth beneath the workpiece surface the ten-sile stress rapidly drops and quickly reaches compressive levels(within 50 �m) before slowly returning to bulk values (within200 �m). The profiles follow the same trend when measuredin both the direction of feed and the direction of cutting butin all cases the feed direction results were slightly lower (seeFigs. 9–12).

When cutting using the same operating parameters the sur-face produced with the coated tool (Tool S) had higher surfacetensile stress than the corresponding surface cut with Tool H,which was uncoated (compare Figs. 9 and 12). This increase intensile stress is due to the fact that Tool S has a multilayer coat-itottttruttI

Fig. 10. Residual stress at 80 m/min cutting speed with a new Tool S.


Fig. 12. Residual stress at 40 m/min cutting speed with a new Tool H.

ng containing a thermal barrier layer of Al3O2 which preventshe heat generated during cutting from dissipating into the bulkf the tool. Although the amount of heat that conducts into theool bulk is considered to be a small percentage, relative to theotal amount generated [23], it would still be sufficient to raisehe temperature of the workpiece surface to a higher level andhus cause the higher tensile stresses seen. A similar result waseported by Arunachalam et al. [22], when turning Inconel 718sing CBN and mixed alumina tools. The lower conductivity ofhe alumina tool led to the production of much higher surfaceensile stresses, the effect on the depth profile was not measured.n the present work, the compressive stress band found beneath



the thin tensile surface layer was similar for both tool types.Given that compressive stresses are generated by plastic defor-mation of the workpiece surface this result is understandable, asboth the depth of deformation and microhardness increase arealso similar for both tool types (see Figs. 5 and 6). As stated inthe literature review the resulting stress state after machining isa combination of the thermal and mechanical effects, with thethermal effects producing tensile stress.

In general, when the cutting speed was increased the peaktensile stress reduced with the greatest reduction occurring at40–80 m/min (see Figs. 9 and 10). At the highest cutting speedused in these trials (120 m/min) the residual stress profiles arevery similar to those obtained with Tool H at 40 m/min. It must beemphasised that Tool H was not used at the higher cutting speedsdue to its relatively low tool life as reported in a previous study[10]. Schlauer and Oden [18] found that when turning Inconel718 using an Al2O3–SiC ceramic tool, at 0.5 mm/rev feed rateand 0.2 mm depth of cut, increasing the cutting speed led tohigher surface tensile stress and an increase in the depth andmagnitude of the compressive zone beneath this. They also con-ducted experiments at different feed rates (0.1 and 0.3 mm/rev)and noted that when a lower feed rate was used the tensile stressat the surface was lower and the effects of increasing cuttingspeed were reduced, although no explanation was given. Thedrop in surface tensile residual stress seen in this work can beexplained when considering the increased chip flow rate associ-artstwi

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Fig. 13. Residual stress at 40 m/min cutting speed with a worn Tool S.



ted with increasing cutting speed. Higher chip flow rates willeduce the length of time available for the heat generated inhe shear zone to diffuse into the workpiece surface and con-equently increase the amount of thermal energy evacuated inhe chip [23]. This reduces the effect of the thermal load on theorkpiece and therefore mechanical effects dominate, resulting

n a more compressive residual stress regime, as seen.As feed rate was increased, tensile stress at the surface and

he depth of the compressive layer both increased slightly (seeigs. 9–12). A number of workers have found that increasing theeed rate results in a trend towards higher and deeper compres-ive residual stresses, due to the generation of higher cuttingorces regardless of the workpiece material being machined18,24–28]. El-Wardany et al. [28] found that an increase in feedate also allows more heat to be dissipated within the (thicker)hip bulk. This reduces the tensile stresses generated in the work-iece by the thermal load, again contributing to higher levels ofompressive residual stress. In the present work the increase ineed rate resulted in an increase in both the depth of deformationf the microstructure and the cutting forces (see Figs. 6 and 7),hich led to the increased stress levels seen.As found for the other measures of surface integrity, the

reatest influence on the residual stress profile was caused byool wear. In all cases when cutting with a worn tool the sur-ace tensile stress increased dramatically, up to a maximumf 1043 MPa when cutting at 40 m/min and 0.25 mm/rev withool S (see Fig. 13). In addition, the stress beneath the sur-ace layer became much more compressive and penetrated todeeper depth (see Figs. 13–16). In a number of cases a sig-

ificant level of compressive stress still remained up to a depthf 0.5 mm beneath the workpiece surface. A number of work-


Fig. 16. Residual stress at 40 m/min cutting speed with a worn Tool H.

ers have noted increased surface tensile stress when machiningwith worn tools but little data is available on the effects on thedepth profile for Inconel 718 [8–9,22]. Increased tensile stressin the near surface layer was caused by the higher cutting tem-peratures developed from rubbing/ploughing of the workpiecesurface by the worn tool flank. An increase in rubbing/ploughingalso leads to greater plastic deformation of the workpiece sur-face, as demonstrated by the increase in depth of deformationand microhardness changes (see Figs. 5 and 6). This leads tothe very high compressive stresses found beneath the thermallyaffected surface layer. The same trends, with respect to changesin machining parameters, noted for the unworn tools, were alsoseen in the surfaces produced with worn tools. The highest ten-sile stress was seen when cutting with a worn Tool S at 40 m/mincutting speed and 0.25 mm/rev feed rate, while the largest com-pressive stress was seen with a worn Tool H at 40 m/min and0.15 mm/rev. It is clear from the results seen in this study that,in order to minimise residual stress changes in a machined com-ponent, the level of tool wear must be kept to a minimum.

Due to the cost and complexity of residual stress measure-ments, one of the aims of this work was to determine whether thestress profile could be ascertained from other more simple mea-surements of surface integrity. The results show that althoughthere are some trends relating changes in surface integrity param-eters to the final residual stress state obtained, the trends are notconsistent. In general, an increase in surface damage results in agtv

4

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pressive levels (within 50 �m depth) and then levelling out withincreasing depth (by 200 �m) below the workpiece surface.

Under the same operating parameters, the surfaces producedwith the coated insert (Tool S) were more tensile due to thegreater heat input into the workpiece surface resulting from theuse of a ceramic coating. As cutting speed was increased thesurface residual tensile stress dropped while an increase in feedrate resulted in a slight increase in both the surface tensile stressand the depth of the compressive stress layer.

The largest influence on the surface integrity generated wascaused by tool wear. An increase in wear led to greater plasticdeformation of the microstructure (up to 36 �m deep), higherlevels of strain hardening (up to 500 HK0.05) and increasedlevels of residual stresses. When cutting with a worn tool thesurface tensile stress increased dramatically followed by a largerand deeper compressive stress layer (up to 0.5 mm). This wasattributed to the increase in plastic deformation and frictioninduced temperature rise caused by rubbing of the worn toolflank on the workpiece surface. It is clear from these resultsthat controlling the level of tool wear is critical and to produceconsistent surface integrity wear should be kept to a minimum.

Although the changes in the residual stress followed simi-lar trends to those seen for microhardness and microstructuraldeformation a large amount of scatter was present in the data.These measurements, along with those of cutting force, cannotbe relied upon to give an accurate prediction of the residual stresssmr

A

Cm

R

eneral increase in the level of residual stress obtained howeverhe measurements could not be used to predict the actual stressalue with any degree of certainty.

. Conclusions

When cutting with a new tool relatively little plastic deforma-ion of the grain boundaries occurred (average 12 �m) and all theamples were strain hardened in the near surface layer to an aver-ge of 440 HK0.05 before dropping to bulk hardness (385 HK0.05)ithin 50 �m depth. Furthermore, the residual stress profile was

ensile in the near surface layer before rapidly dropping to com-

tate but can used to identify the response caused when changingachining parameters and thereby reduce the total number of

esidual stress measurements required.

cknowledgments

The authors would like to thank Chris Mills of AB Sandvikoromant and Ken Williams and Andy Smith of Sandvik Coro-ant UK for the provision of funding and technical support.

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an analysis of the residual stresses generated in inconel 718™ when turning

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