Influence of dressing strategy on tool wear and performance behavior in grinding of forming tools with toric grinding pins

The powder metallurgical high-speed steel AISI M3:2 (ASP^®2023) was selected as experimental material. The chemical composition is shown in Table 1. The high-speed steel was heat treated in vacuum at an austenitizing temperature of 1400 °C and tempered three times at 560 °C for 1 h each time. After heat treatment, the material has a hardness of 62 ± 2 HRC. The specimens used have a cuboid geometry with dimensions of 43 × 30 × 12 mm³.

The measurements of the geometry of the toric grinding pins and of their wear were performed using a Helicheck measuring machine (Walter AG, Germany) and an Infinity Focus XL200 G5 topography measuring device (Alicona Imaging GmbH, Austria). The topography measurements were evaluated using the software “µsoft analysis premium 7.11”^® (NanoFocus AG, Germany).

The residual stresses are measured by using the sin²ψ-method in ω-mode as described by Macherauch and Müller [15]. A Seifert 3003 TT dual-circuit diffractometer system (GE Inspection Technologies, Germany) with a Cr-anode in combination with a V-filter and a collimator with a diameter of d_c = 2 mm is used for the measurement. The measurement is performed with an anode current of I_a = 35 mA and an accelerating voltage of U_a = 30 kV, with an information depth of τ_max = 5.5 µm. The Bragg reflections and elastic constants used for the workpiece material AISI M3:2 are listed in Table 4.

The condition of the toric grinding pins after dressing and after grinding is measured in the surfaces in the contact area and in the torus radius. In the case of the toric grinding pin surfaces, a qualitative evaluation of the initial state and the clogging after the grinding process is carried out. When the toric grinding pins are adjusted to β_fN = 30°, the contact geometry in the contact area can be described by the torus radius r, which makes it an important parameter. Furthermore, this parameter is an indicator for the form accuracy of the dressing process, which allows the description of the influences of the contact geometry between toric grinding pin and workpiece. The description of the wear behavior of the toric grinding pins in dependence on the dressing strategy and grinding strategy is done by the G-ratio in addition to the consideration of the toric grinding pin surfaces. Due to the rotational symmetry of the toric grinding pins, the calculation is based on the ratio of the torus radius before and after the grinding process. The higher the G-ratio, the more favorable is the wear behavior of the toric grinding pins.

A target value of r = 5 mm is aimed for the torus radius after dressing. The contour measurement after the dressing process results in a torus radius of r = 5.04 ± 0.008 mm for dressing combination A1 (crushed, q_d = 1) which means a tolerance of about 0.16% to the target value. For dressing combination A2 (q_d = 0.8), a torus radius of r = 4.89 ± 0.007 mm is measured which means a tolerance of about 0.14% to the target size. Due to the small deviations, it can be assumed that the form tolerance has no influence on the operational and wear behavior. The initial condition of the toric grinding pin surfaces after dressing is shown in Fig. 4 for both dressing processes. Both dressing processes lead to a high degree of porosity of the surfaces, which is common for toric grinding pins with a vitrified bond. Cracks, clogging or other contaminations are not visible. Overall, the surfaces show qualitatively comparable dressing results.

The surfaces of the toric grinding pins after machining a volume of V_Wst = 804 mm³ are shown in Fig. 5. The qualitative representation of the tool surfaces shows the influence of different grinding and dressing strategies on clogging. Toric grinding pins used with the frontal grinding strategy (Fig. 5, top) show qualitative differences in the clogging of the surface depending on the two dressing processes. Tools dressed by crushing (dressing strategy A1) show a porous, almost clog-free surface, which is comparable to the initial surface. The almost clog-free condition of the crushed tools can be explained by the operating principle of the dressing process and by the additional transverse forces acting on the pin due to the process kinematics in the case of the frontal grinding strategy [7]. In crushing, as described in Sect. 2, dressing is performed by crushing the bond, which leads to microcracks in the bond, and facilitates crack propagation and allow the bond layers to break out at lower process forces. This promotes the self-sharpening effect. The additional transverse forces that are a result of the frontal grinding strategy also promote this. However, the early self-sharpening leads to faster tool wear and thus to lower G-ratios (see Fig. 6). In contrast, clogging occurs on tools dressed with dressing strategy A2 and used with the frontal grinding strategy. In dressing strategy A2, the dressing process is performed by grinding off the bond and abrasive grit, which does not promote self-sharpening. Toric grinding pins used in the lateral grinding strategy (Fig. 5, bottom) show clogging on the surface for both dressing processes. Significant differences are not evident.

The different amounts of clogging depending on the grinding strategy can be explained in terms of the different geometric contact variables (cf. Fig. 3; Table 5). For the lateral grinding strategy the contact length l_g is about twice as great for the first path and greater by a factor of 4.7 for all other paths, compared to the frontal strategy. The contact width b_g for the frontal grinding strategy, however, is more than twice as great for the first path compared with the lateral grinding strategy, and by a factor of 2.5 greater for all further paths. This results in a higher contact cross section A_g for the frontal grinding strategy, which is about twice as large for the first path and 50% larger for all other paths than for the lateral grinding strategy. Due to the small cross section with a simultaneously large contact length for the lateral grinding strategy, the thermal tool load is distributed over a small annular area of the toric grinding pin. The higher temperatures and smaller chip thicknesses can lead to chip adhesion and clogging being favored in the lateral grinding strategy.

An analysis of the G-ratios for the frontal grinding strategy shows that crushed toric grinding pins have a lower G-ratio by a factor of 4.8. For the lateral grinding strategy the opposite behavior can be observed. These differences can be explained by the grain load in the respective grinding strategy. In the case of the frontal grinding strategy, there is an additional transverse load on the abrasive grain as a result of the orthogonally aligned cutting and feed movements. The resulting shear forces are not present in the lateral grinding strategy. The transverse load acting on the toric grinding pin, in combination with the active principle of crushing, leads to premature breakout of the bond and the abrasive grains. Why the dressing strategy does not seem to have any influence on the clogging in the case of the lateral strategy and why the opposite behavior is observed with the G-ratios could not be conclusively clarified within the context of the present study and requires further investigations.

For each specimen, XRD sin²ψ-measurements were made after a defined machined volume. The results of these residual stress measurements as a function of the machined volume and the dressing strategy are shown in Fig. 7. Higher compressive residual stresses are induced orthogonal to the cutting direction. The difference between the parallel and orthogonal induced residual stresses is on average ∆σ = 269 MPa for the frontal strategy for dressing combination A1 (crushed, q_d = 1) and σ = 353 MPa for dressing combination A2 (q_d = 0.8). For the lateral strategy, there is an average difference of σ = 246 MPa for dressing combination A1 and σ = 263 MPa for dressing combination A2. The different residual stress values in the parallel and orthogonal directions for both strategies can be explained by plastic deformation of the material due to machining. When the toric grinding pin exits the cutting path, residual stresses are induced into the subsurface due to the plastic deformation of the material. Parallel to the cutting direction, part of the subsurface and thus part of the already plastically deformed material is removed as chips by further machining. In contrast, the generated subsurface orthogonal to the cutting direction remains completely unaffected [7].

First of all, a rapid increase in the compressive residual stresses can be observed for the toric grinding pins used up to a machined volume of approx. V_Wst = 110 mm³ for the frontal grinding strategy and up to approx. V_Wst = 40 mm³ for the lateral grinding strategy. For toric grinding pins dressed with combination A2 (red line), a drop in compressive residual stresses to the level of the blue curves can be seen for both grinding strategies after a machined volume of 110 mm³ and 40 mm³ respectively. For the frontal grinding strategy, there is a drop in compressive residual stresses of 129 MPa for the orthogonal residual stresses and a drop of 78 MPa for the parallel residual stresses. This increase and subsequent decrease of the compressive residual stresses indicates a significant run-in phase. During this run-in phase, the grains blunt due to the mechanical stress, and grains reaching relatively far out of the bond break out. As a result, the contact area between the toric grinding pin and the material increases. The increasing contact area and the blunted grains lead to an increasing thermal load due to higher friction, which causes the compressive residual stresses to decrease again. For toric grinding pins used with the lateral grinding strategy, this run-in phase is unincisive. For toric grinding pins dressed with strategy A1 (blue), a steady increase in compressive residual stresses can be observed. Due to the principle of action during crushing and the occurring load during the grinding process, the self-sharpening effect is accelerated. The mechanical and thermal stress on the active partners is reduced by the breaking of the bond and the widely protruding grains. The stepwise release of new grain layers increases the degree of plastic deformation, which results in increasing residual compressive stresses. After the run-in phase, with the frontal grinding strategy from a machined volume of approx. 340 mm³ a constant residual stress state is established. With the lateral grinding strategy from a machined volume of approx. 110 mm³, a constant increase of compressive residual stress can be seen. This shows that the dressing strategy has a direct influence on the application behavior during the run-in phase.

Further consideration of the development of residual stresses as a function of the machined volume shows for the frontal strategy that the residual stresses remain at a constant level with increasing machined volume. Here, the crushed toric grinding pin induces on average 13% less compressive residual stresses orthogonal to the cutting direction and 9% less compressive residual stresses parallel to the cutting direction. After a machined volume of V_Wst = 804 mm³, there is a difference of σ = 119 MPa for the orthogonal residual stresses and σ = 77 MPa for parallel residual stresses between the two dressing strategies. Therefore, from the different levels of the induced residual stresses, a correlation with the selected dressing process is evident. For the lateral strategy, after the run-in phase, the compressive residual stresses increase linearly with increasing machined volume for both dressing strategies. The crushed toric grinding pin induces on average 4.5% less compressive residual stresses orthogonal as well as parallel to the cutting direction. The maximum difference is σ = 92 MPa for residual stresses orthogonal to the cutting direction and σ = 53 MPa for residual stresses parallel to the cutting direction. It is evident that the influence of the dressing process on the level of induced residual stresses is not as pronounced for the lateral grinding strategy as it is for the frontal grinding strategy. The influence of the dressing strategy on the course of the induced compressive residual stresses can be explained by tool wear. The frontal strategy more often results in self-sharpening and thus in a comparatively consistent change of the geometry of the toric grinding pin. This in turn results in a change of the mechanical stress on the workpiece surface. This does not apply for the lateral grinding strategy. The observed clogging of these toric grinding pins indicates that self-sharpening occurs less frequently and the tools are therefore more blunt until self-sharpening occurs. With a blunting of the abrasive grains, the mechanical load on the workpiece becomes higher, which is reflected in the development of higher compressive residual stresses. This coincides with the different characteristics of the occurred clogging after grinding. The lower residual compressive stresses induced by crushed toric grinding pins can also be attributed to the self-sharpening effect, which reduces the load between tool and workpiece.