Possibility to Control Surface Integrity in Grinding

Transcription

1 Possibility to Control Surface Integrity in Grinding Bogdan Kruszyński, Technical University of Łódź, Łódź, Poland Shinichi Togo, Tohoku Gakuin University, Tagajo, Japan Ryszard Wójcik, Technical University of Łódź, ŁódźPoland Summary Results of investigations on surface layer generated in surface grinding are presented in the paper. Good correlation was found between surface layer properties like residual stress and microhardness and grinding coefficient developed by combining grinding power density and wheel/workpiece contact time. Also, the influence of process parameters on the grinding coefficient proposed is experimentally evaluated. The experiments were carried out for SKS3 Japanese tool steel and Polish grinding wheels of different parameters in the wide range of grinding conditions. Experimental set-up is also described in the paper. 1. INTRODUCTION Grinding is a manufacturing method combining high material removal rates and good surface roughness. Unfortunately, increase of material removal rate leads, especially in grinding with aluminium oxide grinding wheels, to a disadvantageous increase of heat generation and consequently to an increase of grinding temperatures, temperature gradients and temperature rates in a workmaterial. This in turn causes structural changes, changes in hardness distribution and generation of tensile residual stresses in the surface layer of the workpiece, [1,2]. All these changes of the surface integrity are highly disadvantageous because they lead to a significant deterioration of functional properties of the workpiece, these being of the tribological, fatigue and environmental nature [3,4]. Finding a compromise between high productivity and advantageous surface layer properties is extremely difficult due to, among others, the lack of relatively simple and universal routines. Several general approaches are observed attempting to solve this problem. The first one, which can be considered as strictly analytical e.g. [5,6], is based on the analysis and mathematical description of the physical processes involved in the surface layer creation. In grinding it is the thermal effects that are usually described. On the basis of the calculations of temperature distribution in the workpiece such changes in the surface layer like microhardness, residual stress, microstructure, etc. are estimated, e.g. [5]. Such an approach is very promising, but due to its complexity it is limited to scientific investigations. The other problem is still limited knowledge about materials behaviour in extreme grinding conditions which include very high strain rates, high temperature rates and gradients, etc. The purely experimental approach, e.g. [7], aims at finding a correlation between grinding conditions and surface layer parameters. This relatively simple method is usually time- and capital consuming with limited possibility to extrapolate the experimental results on different grinding methods or even on different workmaterials of different grinding conditions. There is also a third approach to the problem of control of the surface layer creation, which involves a search for such grinding coefficients which are strongly correlated with surface layer properties, e.g. [8,9]. There are many coefficients existing, for example equivalent chip thickness (h eq ), which was proved to be useful in surface roughness control and in predicting residual stress in grinding ceramics, or power density (P``) useful in predicting residual stress in grinding aluminium oxide wheels. The main disadvantage of both coefficients mentioned above is that to calculate them it is necessary to estimate the effective grinding depth or wheel/workpiece contact length. Unfortunately, both values are very difficult to estimate on-line. Moreover, the determination of the effective wheel/workpiece contact length is possible only with very limited accuracy. Below are presented the results of investigations on the relations between the grinding coefficient combining power density and wheel/workpiece contact time and the surface integrity in surface grinding. 2. GRINDING COEFFICIENT

2 It was proved in many publications that high temperatures as well as high temperature gradients and temperature rates are the main reason for generating of unfavourable tensile residual stress and microstructure changes in the surface layer of ground steels. From the equation for grinding temperatures e.g.[1]: ( -.37 Z) 2P α.53 Tmax = 3.1L exp -.69 L (1) πkvft where: - P - heat flux entering the workpiece, - α - thermal diffusivity, - k - thermal conductivity, - v ft - velocity of moving heat source (tangential feed speed), - L - dimensionless length of heat source (L= v ft l c /4 α), - Z - dimensionless depth below surface (Z= v ft z/2 α), - l c - effective wheel/ workpiece contact length (length of the heat source), - z - depth below surface, it emerges that besides the properties of the workmaterial, wheel and coolant it is the heat flux (power density) entering the workpiece that is the parameter which highly influences the temperature distribution in grinding. It was mentioned above that power density is also well correlated with residual stress generated in the surface layer. But it can also be seen from the eq. (1) that there is also another parameter influencing temperature distribution. This is the wheel/workpiece contact time, which is represented in eq.(1) by the heat source velocity v ft. This is obvious that the longer contact is between the heat source and the particular point of the workmaterial is, the higher temperatures are generated. In [11] the parameter B p was developed which is a product of the power density and wheel/workpiece contact time. For surface grinding, coefficient B p can be expressed as follows: P lc P B p = P tc = = (2) bd lc vft bd vft where: P - grinding power b D - active grinding width All quantities in eq. (2), i.e. grinding power, tangential table speed and active grinding wheel width are easy to estimate on-line during the grinding process. The grinding power can be measured either as a product of tangential grinding force and wheelspeed or as a power consumption of the main wheel drive. The active grinding width is a workpiece or wheel width (whichever is less) in plunge grinding or the axial feed in longitudinal grinding. Also the tangential feed speed is easy to measure with simple measurement equipment. This means that coefficient Bp can be estimated on-line with the aid of simple measurement devices installed on the grinding machine, personal computer equipped with measurement cards and a relatively simple software for collection and evaluation of measurement data. At the present stage of research the energy partition was not taken into consideration in calculations of coefficient B p, eq. (2). Actually, energy partition ratio may vary with grinding wheel properties (particularly abrasive and bond materials and wheel structure), workmaterial properties and with kinematics e.g.[11,12,13,14]. Because investigations were carried out for the same workmaterial and conventional (aluminium oxide and silicon carbide) grinding wheels this simplification should not introduce significant errors. For other basic grinding operations, i.e. cylindrical and internal grinding, the equations to calculate coefficient B p are equally simple [15] and there should be expected no problems with on-line estimation. 3. EXPERIMENTS Set-up Surface plunge grinding was performed in the following range of grinding conditions: - workmaterial: SKS3 Japanese tool steel, 6±2HRC - grinding wheels (Polish brands): - silicon carbide grinding wheel,

3 - aluminium oxide grinding wheel - (aluminium oxide and chromium oxide), - grinding fluid: emulsion, - wheelspeed v s : 25.5 m/s (constant), - tangential table feed speed v ft :.8 m/s,.3 m/s and.5 m/s, - effective depth of cut a e : from.2 mm to.17 mm - active grinding width b D : 12 mm (constant, equal to the width of samples) In each grinding test tangential grinding force F t and tangential table feed speed were measured. Then grinding power was calculated as a product of F t and wheelspeed, power density as well as coefficient B p from eq. (2). After each test the residual stress and microhardness were measured. Residual stress measurements were carried out by using a well-known material removal method. An example of typical residual stress distribution with depth below the surface is shown in Fig. 1. In all the tests a local maximum of tensile residual stresses was observed close to the surface followed by a constant decrease of stress to zero and low compressive values. The maximum tensile stresses were evaluated and adopted as representative for each test. The changes of microhardness in the surface layer were measured on ZWICK 3212 hardness tester, using the Knoop method. The load of.1 kg was applied. The hardness at the depth of.3 mm below the surface were adopted as representative for each test. 1 Residual stress [MPa] Surface plunge grinding Workmaterial: SKS3, 6HRC Japanese tool steel Wheel: Coolant: emulsion RESULTS,1,2,3,4,5,6 In Figures 2 and 3 the influence of grinding Depth below parameters surface on z coefficient [mm] B p is shown. It can be seen from these figures that both grinding depth and tangential feed speed highly influence coefficient B p. In Figure 2 the correlation between Fig.1. effective An example grinding of residual depth and stress coefficient distribution B p is vs. presented depth below for different surface tangential feed speeds and different grinding wheels. The coefficient increases with the increase of the depth of cut but the intensity of these changes depends on the range of depths of cut and on tangential feed speed. Higher intensity of changes is observed in the range of lower grinding depths and for lower feeds. The latter is confirmed in Fig. 3 where the correlation between tangential feed speed and coefficient B p is shown for different grinding wheels. It is clearly visible that the intensity of changes is much higher in the range of feeds from.8 m/s to.3 m/s than for faster feed rates. The increase of grinding depth and of the tangential feed speed increases the grinding power but also changes the wheel workpiece contact time. The increase of grinding depth increases wheel/workpiece contact length and, consequently, the contact time. The increase of the tangential feed rate directly decreases the contact time. Diagrams presented in Figures 2 and 3 show that in the case of the grinding depth a dominating factor is the increase of grinding power, whereas in the case of tangential feed rate the shortening of the contact time dominates. The influence of the grinding wheel on coefficient B p also can be seen from these figures. The highest values of the coefficient Bp are obtained when grinding with silicon carbide grinding wheel which is the hardest of all the wheels applied. The wheel with the chromium oxide content () produced the lowest level of coefficient B p in the range of kinematical parameters investigated. This wheel was the softest one. The aluminium oxide-grinding wheel () gave the intermediate values between silicon carbide and chromium grinding wheels.

4 Coefficient B p [W s/mm 2 ] vft=.8 m/s vft=.5 m/s Effective depth of cut ae [mm] Fig.2 The influence of effective grinding depth on the coefficient B p 7 Coefficient B p [W s/mm 2 ] Increase of wheel hardness vft=.5 m/s vft=.8 m/s Tangential table feed speed vft [m/s] Fig.3 The influence of tangential table feed speed on the coefficient B p The changes of B p with the grinding wheel brand are generally within 2% of the values obtained for the 99A6K-grinding wheel. In Fig.4 the correlation between coefficient B p and maximum residual stresses in the surface layer is shown for different grinding wheels. A linear approximation was applied to the results obtained with good effect. For all the grinding wheels applied were obtained very similar coefficients of linear regression: for wheel, for wheel and for one. It indicates that the workmaterial is the main factor determining the inclination of the σ-b p lines. The shift among the lines plotted for different diagrams is caused, most probably, by differences of abrasive materials and wheel harnesses, which caused changes in grinding energy partition between workmaterial and the grinding wheel in particular cases. In Fig. 5 the relations between the B p coefficient and changes of microhardness are shown. In this case no substantial differences were observed for different grinding wheels and all the results were included in one diagram. It can be seen from this Figure that for the small values of B p (up to.5) there are almost no changes observed in the surface layer hardness. After the magnitude of B p =.5 is exceeded, a sharp drop in microhardness is visible. It indicates that the value of.5 for B p is a threshold value for structural changes in the surface layer of the SKS3 steel. A rather low threshold B p level indicates that this steel is difficult to grind and sensitive to elevated temperatures in grinding.

5 In the range from B p =1. to B p =4 the slow decrease in microhardness is observed followed by the next sharp drop for the higher magnitudes of B p. It indicates the very high rise of temperatures for the high values of B p. 12 Residual stress [MPa] Coefficient B p [W s/mm 2 ] Fig. 4 The influence of coefficient B p on maximum residual stress 14 Microhardness HK.1 at the depth.3 mm below surface Bp J very intensive burns HK burns start to appear Coefficient B p [W s/mm 2 ] Fig.5. The changes of microhardness vs. coefficient B p. There is a difference between the relation of residual stress and microhardness vs. coefficient B p. This is caused mainly by different phenomena and processes involved in changes of both parameters. In the case of microhardness it is mainly the level of temperature which determines the softening of the workmaterial outer layer. In the case of residual stress the process is more complicated. The changes of residual stress may be caused by the plastic deformation due to gradient of temperature and/or by changes of microstructure combined with changes of the volume of material caused by elevated temperatures. The results obtained indicate that coefficient B p may be used to predict and/or control surface layer parameters.

6 5. CONCLUSIONS 1. Experimental results proved good correlation between coefficient B p and surface layer parameters like residual stress and microhardness. 2. Linear dependence between B p and residual stress was found. The slope of the lines is characteristic for the workmaterial and seems to be independent of the grinding conditions (eg. grinding wheel, process parameters). 3. The hardness changes with coefficient B p show characteristic thresholds, which may indicate the increase of grinding temperatures over the phase transformation points. 4. The coefficient B p which is easy to estimate on-line may be useful in predicting and/or controlling the residual stress and microhardness changes in the surface layer in conventional grinding. The investigations were carried out with support of Polish National Committee for Scientific Research (KBN), Grant no. 7 T7D REFRENCES [1] Tonshoff H.K., Friemuth T., Becker J.C.: Process Monitoring in grinding. Annals of the CIRP, 51/2/22. [2] Rowe W.B., et al: A Simplified Approach to Control of Thermal Damage in Grinding, Annals of the CIRP 45/1/1996, pp [3] Kruszyński B., Luttervelt C. A. van, 1989, The influence of manufacturing processes on surface properties. Advanced Manufacturing Engineering, vol. 1/4: [4] Kruszyński B., 199, Basics of the surface layer. Scientific Bulletin of Technical University of Lodz, No.79 [5] Vansevenant E., 1987, A Subsurface Integrity Model in Grinding, PhD thesis, KU Leuven. [6] Tönshoff H, K., et al., 1992, Modelling and simulation of grinding process, Annals of the CIRP, 41/2: [7] Zheyun Y., Zhonghui H., 1989, Surface integrity of grinding of bearing steel GCr15 with CBN wheels, Annals of the CIRP, 38/1: [8] Brinksmeier E., et al., 1993, Basic parameters in grinding, Annals of the CIRP 42/1: [9] Kruszyński B., Luttervelt C.A. van, 1991, An attempt to predict residual stresses in grinding of metals with the aid of the new grinding parameter, Annals of the CIRP, 4/1: [1] Takazawa K., 1972, Thermal aspects of grinding operation. Industrial Diamond Review No.4. [11] Rowe W.B., et al.: Experimental Investigation of Heat Transfer in Grinding. Annals of the CIRP, 44/1/1995, pp [12] Black S.C.E., et al.: Experimental Energy Partitioning in Grinding. Proc. of the Eurometalworking Conference, EMW 94. Udine 1994, pp [13] Rowe W.B., Morgan M.N., Black S.C.E.: Validation of Thermal Properties in Grinding.: Annals of the CIRP, 47/1/1998, pp [14] Guo, C., Malkin, S.: Inverse Heat Transfer Analysis of Grinding, Transactions of the ASME, Journal of Engineering for Industry, February 1996, vol 18, pp [15] Kruszyński B.W, 1994, The conception of evaluation of the grinding process on surface integrity. (in Polish). Proceedings of the conference on Material Removal Processes, Koszalin, Poland.