【机械类毕业论文中英文对照文献翻译】硬齿面精加工看作是磨料磨损过程
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机械类毕业论文中英文对照文献翻译
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【机械类毕业论文中英文对照文献翻译】硬齿面精加工看作是磨料磨损过程,机械类毕业论文中英文对照文献翻译,机械类,毕业论文,中英文,对照,文献,翻译,硬齿面,精加工,看作,磨料,磨损,过程
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Wear 258 (2005) 6269Hard gear finishing viewed as a process of abrasive wearE. Brinksmeier, A. GiwerzewFoundation Institute of Materials Science, Bremen, GermanyReceived 2 December 2003; accepted 7 December 2003Available online 14 November 2004AbstractAbrasivemachiningisoneofthekeytechnologiesingearmanufacturing.Amongstabrasivefinishingprocessesofhardenedgears,operationscharacterised by cutting speeds as low as 0.35m/s have been becoming more and more important during the last 20 years. A lot of effort hasbeen spent to increase the productivity of such low speed grinding processes. A part of this effort is the development of new abrasive tools,which involves much empirical investigation and is, therefore, very time and cost intensive.The present study is aimed at a deeper understanding of differences in the process behaviour of various grinding wheels. For this purpose,the grinding process has been analysed as a process of abrasive wear. The chosen research approach implied a combination of grindingexperiments and analytical calculations. To characterise the process quantities and work results in grinding, an abrasive wear map was used.Furthermore, calculations of the maximum single grain chip thickness, the mean single grain normal forces and X-ray measurements ofresidual subsurface stresses were made. Results of the investigations show that at constant grinding conditions, the maximum single grainchip thickness determines the tool performance. Additionally, the essential work results in grinding have been observed to correlate withthe overall friction coefficient in the contact area regardless of the used tool. These results can be advantageously used for both industrialdevelopment of new abrasive tools and further fundamental research. 2004 Elsevier B.V. All rights reserved.Keywords: Abrasive wear; Low speed grinding; Single grain chip thickness; Friction1. IntroductionAbrasive processes using cutting speeds as low as0.35m/sandchangingworkingdirectionhavebeenincreas-ingly applied in hard gear finishing operations since the mid1980s. Terms like “gear honing”, “power honing”, “spherichoning”, “shave-grinding” and others describe all the samekinematic principles of a bonded abrasives technology basedonameshingcontactofgear-shapedtoolswiththeworkgearin the form of a helical gear pair 1. Compared to profile orgeneratinggrindingtraditionallyusedingearmanufacturing,the technology is applied with approximately 1060 timeslower cutting speeds and 2030 times lower specific mate-Correspondingauthor.Presentaddress:IWT,StiftungInstitutfurWerk-stofftechnik, Badgasteiner Str. 3, 28359 Bremen, Germany.Tel.: +49 421 218 5488; fax: +49 421 218 3272.E-mail address: bleiliwt.uni.bremen.de (E. Brinksmeier).rial removal rates, see Table 1. Due to low cutting speeds,maximum temperatures in the tribological contact area areestimated to be in the range of 100300C 2. Such lowtemperatures in combination with a contact duration in themagnitudeofsomemicro-secondsexplainthefact,thatwork-piece thermal damage in these processes has never been ob-served in the past. Subsurface layers of the low speed groundgear flanks show high compressive residual stresses, whichremain detectable at a somewhat lower amplitude over theentire gear lifetime 3,4. The technology improves surfacefinish and offers the possibility of geometrical tooth flankcorrections. Furthermore, a characteristic surface texture ofthe finished gears reduces the noise emission of gearboxes5,6.A variety of different abrasive tool types are available forcarrying out hard gear finishing operations at low cuttingspeeds 7,8. Depending on the machine tool type and work-piece quality requirements (stock removal, surface rough-0043-1648/$ see front matter 2004 Elsevier B.V. All rights reserved.doi:10.1016/j.wear.2004.09.032E. Brinksmeier, A. Giwerzew / Wear 258 (2005) 626963Table 1Comparison of grinding technologies used in gear manufacturingProfile grinding/generating grindingShave-grinding (gear honing)Cutting speed (m/s)High (20 . 60)Very low (0.3 . 5.0)Specific material removal rate(mm3/mms)High (1 . 10)Low (0.05 . 0.3)Subsurface thermal damageMay occurNever observedQuality of surface finish (?m)High (Rz=5 . 10)High (Rz=2 . 5)Profile/form correctionPossible ?Possible ?Noise emission of assembledgearingLower as compared to profile/generating grindingness, profile accuracy), electroplated diamond and cubicboron nitride (cBN) superabrasives or conventional corun-dum abrasives in a resin or vitrified bond can be used. Forcorundum abrasive materials, mixed-type bonding systemsconsisting of vitrified conglomerates in a resin matrix, aswell as resin-infiltrated vitrified bonds are available. Choos-inganabrasivetoolforaspecialmachiningtaskreliesmostlyon practical experience and less on the basic technologicalknowledge, since the tool design alone contains no directinformation on the tool performance.The objective of the present study was to find physicalquantities allowing an explicit characterisation of the toolperformance and work result for tools commonly used in thementioned low speed grinding operations. For this purpose,the machining technology of interest has been viewed as aprocess of abrasive wear. The process analysis was carriedout based on experimental data obtained from grinding tests.Evaluating the experimental data, an abrasive wear map, achip thickness model and X-ray measurements of the work-piece residual stresses were applied.2. Scope of experimental investigationsThe scope of the experimental investigation includedgrindingtestsoftheeightcommerciallyavailablegearhoningtoolslistedinTable2.Thetoolshaddifferenttypesofconven-tional corundum abrasive materials (ruby, seeded gel, whiteand monocrystalline corundum); some of them contained amixture of two abrasive types. The grain size of abrasivesvaried from F80 to F220 according to the Federation of Eu-ropean Producers of Abrasives (FEPA) standard correspond-ingto180and59?mmeangraindiameter,respectively.Toolbonds investigated belonged to different resin, vitrified andmixed-type systems. Besides the type of the matrix material,the systems were characterised by different relative pore vol-umes, being comparably high for vitrified and low for resinbonds.The investigation was carried out on case-hardened steel1.7131 (16MnCr5E). Cylindrical workpieces of an outer di-ameter of approximately 45mm and a width of 4mm wereheat-treated to ensure a material structure similar to that ofcase-hardened gears. The heat treatment consisted of gas-carburisation to a depth of 1.1mm, direct-hardening in oiland annealing in circulated air to a surface hardness of ap-proximately 700HV. Prior to experimental investigations,eachworkpiecehasbeenpre-machinedbycylindricalplungegrinding to remove an oxidised subsurface layer and to es-tablish the necessary concentricity of the outer diameter andthe pickup hole.Grinding tests were performed in an analogous processfor gear honing. The analogous process is designed as an ex-ternal face plunge grinding one modified by an axial offsetbetween the tool and workpiece axis as shown in Fig. 1. Byintroductionoftheaxialoffset,peripheralworkspeed ? vftandwheel speed ? vsare no longer parallel vectors. In the shownconfiguration, the cutting speed ? vcis a vector difference ofthe wheel and work speeds as the absolute value of the ra-dial feed speed vector ? vfris negligible as compared to theabsolute values of ? vsand ? vft. A detailed description of theprocess kinematics is given in former publications 1,2. Thetest bench used was equipped with a piezoelectric force dy-namometer for an in-process measurement of the grindingforce components and by an acoustic emission (AE) sen-sor for an automatic recognition of the first contact betweentool and workpiece. Grinding tests were carried out usingthe machining conditions specified in Figs. 25, which wereheld constant throughout the experiments. The only param-eter varied was the grinding wheel. Before carrying out agrinding experiment, the grinding wheel was dressed by adiamond roller. The dressing operation was performed underthe conditions listed in Table 3.Each grinding experiment consisted of a series of 18 uni-fiedgrindingcyclesperformedinonesetupwithoutaninter-mediate tool dressing. The grinding cycle comprised a quickmotion of the tool towards the workpiece 30?m deep be-yond the initial contact point (initial preload), followed bya 26.4s long continuous feed under preload and a systemrelease again by a quick motion of the tool reverse to thefeed direction. This design provided the total grinding feedrealised within an experiment being lower than the depth ofhardenedcase,sothatthegrindingoperationwasappliedonly64E. Brinksmeier, A. Giwerzew / Wear 258 (2005) 6269Table 2Composition of the investigated wheelsWheelAbrasive typeGrain size (FEPAa-standard)Bond typeCommentARuby corundumF180ResinCasting bondBSeeded gel/white corundumF150/F180ResinCMonocrystalline/white corundumF150/F180ResinDMonocrystalline/white corundumF100/F120ResinHollow balls covered with abrasives substitute poresEMonocrystalline/white corundumF80/F100ResinInclusions of vitrified agglomeratesFMonocrystalline/white corundumF80/F100/F150VitrifiedGMonocrystalline/white corundumF80/F100/F150VitrifiedPores infiltrated with resin bondHMonocrystalline/white corundumF220/F220VitrifiedaFEPA: Federation of European Producers of Abrasives.Table 3Process parameter settings in dressingParameterValueDressing roller typeD181Dressing modeUp-cutPeripheral roller speed vr(m/s)0.63Peripheral wheel speed vsd(m/s)1.10Radial feed rate vfrd(?m/s)3.75tothehardmaterial.Toensurethestatisticrepeatability,threegrinding experiments of the described type have been carriedout for each tool investigated. For each experiment, processforces, workpiece roughness, effective material removal andtool wear have been measured.3. Experimental results3.1. Grinding wheel testsFig. 2 shows grinding performance of the investigatedtools. It was evaluated with respect to the specific (relatedto the contact width) normal grinding force F?n, mean peak-Fig. 1. Kinematics of the grinding process.to-valley workpiece roughness Rz, specific (related to thecontact width) effective material removal V?wand grindingratio G (ratio of the effective material removal Vw,effto thewear volume of the wheel Vs). In the diagram, the scatter baron the each bargraph indicates the standard deviation fromthe appropriate arithmetic mean value.As can be seen, resin-bonded grinding wheels AE werecharacterised by high specific normal forces F?n(approxi-mately 6090N/mm), small values of the workpiece rough-ness Rz (approximately 2.55.5?m) and relatively low val-ues of the specific effective material removal V?w(approxi-mately 4060mm3/mm). These tools showed comparativelyhigh values of the grinding ratio G lying between approxi-mately 6 and 20mm3/mm3. In contrast to that, vitrified toolsFH showed low specific forces (F?n=3.58.5N/mm), roughworkpiece surfaces (Rz=815?m), high specific materialremoval (V?w=95120mm3/mm) and low grinding ratio val-ues (G=14mm3/mm3).AssessmentofgrindingbehaviourundertakeninFig.2re-vealssignificantdifferencesamongstthewheelsinvestigated.However, it does not explain physical phenomena of the ob-served differences. To clarify them, the results achieved havebeen analysed by means of a wear map and a chip thicknessmodel.Fig. 2. Performance of the grinding wheels under study.E. Brinksmeier, A. Giwerzew / Wear 258 (2005) 626965Fig. 3. Experimental data plotted on the abrasive wear map 10.3.2. Interpretation of material removal mechanisms bymeans of abrasive wear mapAs already mentioned, thermal effects in the investigatedlow speed grinding technology are considered to be negli-gible compared to mechanical interactions between contactsurfaces. For this reason, a type of mechanical wear maporiginallyproposedbyChilds9hasbeentakenintoconsid-eration to interpret material removal mechanisms. The wearmap was used in a generalised form established by Williamsfor the case of a soft body abrasion by a hard rough counter-face 10.In analogy to true wear processes, the grinding process ofinterest was viewed as a process of abrasive wear. For thegrinding conditions used in this study, grinding behaviour ofthe each tool was characterised by the non-dimensional wearcoefficientKandgrindingforceratio.ThewearcoefficientK was determined according to the Archards wear equationFig. 4. Correlation between G-ratio, wear coefficient and grinding forceratio.11:K = Vw,effHFnL(1)where Vw,effis the effective material removal, H the work-piece hardness, Fnthe normal component of the grindingforce and L is the grinding distance. The grinding force ra-tio was calculated as the ratio of tangential Ftto normalgrinding force Fn: =FtFn(2)For plotting the experimental data on the wear map 10,an estimation of the normalised interfacial shear strength /k(shear strength of the interface between contact bodies re-lated to the shear strength k of the workpiece material) forthe tools used had to be made. Deviating from the work ofChilds, the interfacial shear strength /k has been approxi-mated by the grinding force ratio . Using and K as inputparameters,theabsolutepositionofeachparticularcombina-tion toolworkpiece on the map has been quantified. Fig. 3shows the abrasive machining wear map section 10 withdata points indicating the appropriate grinding wheel.In the diagram, changes of the wear mechanism with cor-responding wear coefficient values are described dependingontheinterfacialshearstrength(grindingforceratio)andthemean inclination angle between sliding abrasives and work-piece surface (asperity attack angle). In abrasive machining,theasperityattackangleisproportionaltothegrainsdepthof cut or the effective chip thickness hcu,eff. By extrapolat-ing data point positions on the axis of abscissa, a conclusionon qualitative changes of the effective chip thickness can bemade by comparison of relative abscissa values correspond-ing to the different tools.Effectivechipthicknessvaluesoftheresin-bondedwheelsAEareofthesamemagnitudeandrelativelysmallcompared66E. Brinksmeier, A. Giwerzew / Wear 258 (2005) 6269to the vitrified wheels FH. Within the vitrified wheels, theeffective chip thickness increases in the sequence GHF.The absolute values of the wear coefficient K achieved withdifferenttoolstendtoincreasewheneitherhcu,eff-or-valuesare increasing. This observation is illustrated in Fig. 4 show-ing a good correlation between the quantities grinding forceratio and wear coefficient K on the one hand as well asgrindingforceratioandgrindingratioGontheotherhand.Inbothcases,powerlawfunctionshavebeenchosentofitthemeasurement points. Applying the method of smallest stan-darddeviation,anegativeexponentof2.7wasachievedforthe G curve and a positive exponent of 4.9 for the Kcurve.In Fig. 4, two important results should be pointedout, which seem to relate grinding performance of thetools with elemental physical processes taking place at thetoolworkpiece interface. The first one refers to chip forma-tion and the second to friction processes.For abrasive machining, the wear coefficient K can be in-terpreted as an indirect measure of the efficiency of mate-rial removal. Indeed, when the chip formation mechanismis ideal micro-ploughing, no material is removed (Vw,eff=0)and K=0 is considered to be the lower threshold value of theprocess efficiency. A rough estimation of the upper thresh-old value using the model 12 for common abrasive ma-chining with asperity attack angles in the range of 2060leads to K=2/tan, when the chip formation mechanismis ideal micro-cutting 13. Regarding the absolute K-valuesdetermined for the tools investigated, the fraction of micro-ploughing on chip formation is supposed to be relativelyhigh for the resin-bonded wheels AE, since the K-valuesdo not exceed 0.003. When vitrified wheels FH were used,thefractionofmicro-ploughingdecreasedsignificantlyastheK-values increased to 0.07.Treating the grinding force ratio as a quantity characteris-ing friction processes in the contact zone, for a given work-piece material and constant grinding conditions, the use oftools with higher friction is found to improve the efficiencyof chip formation. At the same time, severe friction increasestool wear and thus reduces the grinding ratio G. An expla-nation of this observation can be gained from the analysis ofexperimental data with a chip thickness model.3.3. Analysis of experimental data with a chip thicknessmodelTo quantify the qualitative differences in chip thicknessidentifiedbyplottingtheexperimentaldataonthemechanicalwear map, calculations of nominal chip thickness values forthe grinding wheels and process kinematics used have beenperformed. Furthermore, the number of active grains in thecontact zone and the mean single grain normal forces havebeenestimated.Theanalysiswasbasedonthechipthicknessmodel of Werner 14 assuming a non-uniform distributionof abrasive grains over the wheel periphery and a triangularshape of the chip cross-section.According to the model, the nominal value of the maxi-mum undeformed chip thickness hcu,maxis equal to:hcu,max= 0.695?2C1tan?1/3?1q?1/3?aedeq?1/3(3)with grain density C1, grain shape factor tan?, speed ratioq, working engagement aeand equivalent wheel diameterdeq. The quantities C1and tan? describing grinding wheeltopography were determined by roughness measurement ofthe wheel surface using a contact stylus method. For the pro-cess kinematics used, working engagement aewas equal toradial feed fr, whereas speed ratio q and equivalent wheeldiameter deqagreed with the work of Schneider 15:q =vsvftsin( + )sin,for ?= 0(4)deq=dwsin2,for ?= 0(5)with wheel speed vs, work speed vft, cutting angle , offsetangle and workpiece diameter dw(see Fig. 1). The modelof Werner 14 describes the total number of active grainsper unit wheel area dependent on the wheel topography andworkingconditionsbythedynamicgrainnumberNdyn,whichis defined asNdyn= 1.2?2C21tan?1/3?1q?1/3?aedeq?1/6(6)To calculate the number of active grains Nactaccomplish-ingtheactualchipformationinthecontactzone,thedynamicgrain number Ndynwas multiplied with the contact zone di-mensions width of cut apand effective contact length le:Nact= Ndynaple(7)The contact zone dimensions were calculated usingSchneiders procedure 15:ap bwcos,for ?= 90(8)le=?1 1qsinsin( + )?dwsin2(fr+ Rz),for ?= 0(9)with workpiece width bw. Having calculated the number ofactive grains Nactapplying Eqs. (4)(9) and neglecting a di-rect load bearing by the bond, the mean single grain forcecould easily be estimated by relating the total grinding forceto the number of active grains. Thus, the mean single grainnormal forceFn,gwas calculated as:Fn,g=FnNact(10)Table 4 shows the calculation results for grinding wheelsAC (resin bonds) and FH (vitrified bonds). Because ofrelatively inhomogeneous structures of the wheels D and E,E. Brinksmeier, A. Giwerzew / Wear 258 (2005) 626967Table 4Comparisonoftopographyconstants,maximumundeformedchipthickness,dynamicandactivegrainnumbersandmeansinglegrainnormalforcesfordifferentgrinding wheelsWheelParameterC1(mm3)Parameter tanDynamic grain numberNdyn(mm2)Number of active grainsin the contact zone NactMaximum undeformed chipthickness hcu,max(?m)Mean single grain normalforceFn,g(N)A203955.81031710.921.16B173605.6931570.981.61C172545.9921830.971.13DEF24242.4331302.510.08G55553.6511691.670.14H51393.0521481.760.09topography constants C1and tan could not be reliably de-termined for them using the contact stylus method. Thesewheels were, therefore, excluded from the analysis.As can be seen, nominal values of the maximum unde-formed chip thickness of the resin-bonded wheels AC areroughly the same and significantly lower than that of the vit-rifiedwheels.Withinthevitrifiedwheels,thehighesthcu,max-value shows the wheel F followed by the wheels H andGthese results exactly confirm the conclusions made inthe qualitative chip thickness analysis by abrasive wear map.Itisnoteworthy,thattheabsolutenominalhcu,max-valuescal-culated according to the model 14 are virtually larger thantheradialfeedfrusedasituation,whichisunlikelytooccurin practice. The reason of this conflict is seen in the simpli-fication of grain geometry with a taper cone, which is inap-propriate for fr-values much lower than the average cuttingedge radius of a grain 16. This model limitation, however,does not restrict the predicted relative tendencies.As the kinematic process parameter settings in grindingwere identical for the wheels tested, the differences in chipthickness values can be explained by the different wheel to-pography characteristics given in Table 4. Compared to theresin-bonded wheels, all vitrified wheels show relatively lowstatic grain density C1and sharp cutting edges (lower tan-values). As a result, for the vitrified wheels the number ofactive grains per unit wheel area (dynamic wheel numberNdyn) is calculated to be two to three times lower than the ap-propriate values of the resin-bonded wheels. Consequently,applying the vitrified wheels a unit material volume is re-moved by a less number of active grains having higher chipthickness values.Similar to chip thickness values, the calculated values ofthe mean single grain normal force show significant differ-encesbetweentheresin-bondedandvitrifiedwheels.Despitethe higher chip thickness, the single grain normal forces cal-culated for the vitrified wheels are more than tenfold lowerthan for the resin-bonded ones. This result can obviously beattributed to a lower bond strength of the vitrified wheels. Athigh chip thickness values, low bond strength is regarded toenhance grain breakout, which is supposed to be the crucialfactor for a higher wheel wear (lower G-ratio) of these tools.Different single grain forces should also differently affectthe workpiece material layer being in contact with the abra-sive. This influence has been investigated by residual stressmeasurements of the ground workpieces.3.4. Residual stress measurementsResidual stress measurements were carried out using anX-ray diffractometer and Cr K? radiation. For each of theworkpieces investigated, depth profiles of residual stressestangential to grinding direction were determined applyingstepwise material removal by electrolytic polishing. Depthprofiles of residual stresses induced by the different grindingwheels are shown in Fig. 5.Eachsingleresultshowninthediagramisameanvalueofaseriesofthreeresidualstressmeasurements.Standarddevi-ationvaluesofsuchaseriesvariedbetweenapproximately10and 30MPa. Upon the results of all measurements, the mea-surement uncertainty is estimated to be lower than50MPa.Fig. 5. Depth profiles of residual stresses depending on the grinding wheelused.68E. Brinksmeier, A. Giwerzew / Wear 258 (2005) 6269To enable a clearer comparison of the residual stress pro-files amongst the wheels, error bars of the single points areomitted.In good agreement with previously reported results3,4,15, only compressive residual stresses were measuredforallwheeltypes,indicatingthermaleffectstobenegligiblecompared to mechanical ones. The absolute stress values atthe surface and the depth profiles of residual stresses seemto depend on the type of grinding wheel and the mean singlegrain force.For the resin-bonded wheels (top diagram), depth pro-files of residual stresses show a steep gradient. The stressvalues have their minimum at the workpiece surface andamount to approximately 700MPa, with increasing depththey increase to the residual stress level of the case-hardenedlayerandremainedthenunchanged.Thepenetrationdepthofcompressive residual stresses induced by these tools variesbetween approximately 10 and 30?m. In contrast to that,grinding with the vitrified wheels only slightly changes theresidual stress values after heat treatment both at the surfaceand in deeper layers (bottom diagram).Theobserveddifferencesintheresidualstressprofilescanbe attributed to different single grain forces, which are addi-tionallylistedinthelegendsofbothdiagrams.Thecalculatedmean single grain normal forces of the resin-bonded wheelsare an order of magnitude higher compared to the vitrifiedones. Taking into account smaller mean grain sizes of theresin-bonded wheels, higher mean contact pressures beneaththe grain are supposed to be induced when grinding withthese wheels, causing higher compressive residual stressesand stronger depth gradients.An obvious correlation between single grain forces andcompressive residual stresses indicates mechanical actionsare responsible for this phenomenon. In parallel to chip for-mation, workpiece material beneath the abrasive experiencesplasticdeformation.Thedeformationincreaseswiththeratioof micro-ploughing on chip formation. This result also con-firmsthedeductiononagreatermicro-ploughingwhenusingresin-bonded wheels rather than vitrified ones.4. ConclusionsConventionalcorundumgrindingwheelstypicallyusedinthe technology of low speed hard gear finishing show a quitedifferentgrindingperformanceatconstantprocessparametersettings. Viewing the machining technology as a process ofabrasive wear, following conclusions can be drawn from theresults of this study:1. Grinding behaviour of the grinding wheels is found to beprimarily dependent on their chip thickness values. Forgiven kinematic conditions, the nominal chip thicknessvalue of a wheel is determined by the mean grain shape,grain size and grain spacing. The last parameter is addi-tionally influenced by the wheel porosity, which is alsodependent on the bonding system.2. For the wheels investigated, the influence of the bondingsystemonthechipthicknessisevidentlymuchhigherthanthe influence of the grain size. As a result of a relativelyhigh wheel porosity, chip thickness values of the vitrifiedwheels are higher compared to the appropriate quantitiesof the resin-bonded wheels.3. Duetosmallchipthickness,resin-bondedgrindingwheelsshow comparatively low effective material removal andlow tool wear at low friction values. They can be further-more characterised by a high fraction of micro-ploughingon chip formation causing high compressive residualstresses in the workpiece subsurface layer.4. Higheffectiv
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