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Influences of grinding with Toric CBN grinding tools on surface and subsurface of 1.3344 PM steel
Journal of Materials Processing Technology 229 (2016) 541–548
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Journal of Materials Processing Technology journal homepage: www.elsevier.com/locate/jmatprotec
Influences of grinding with Toric CBN grinding tools on surface and subsurface of 1.3344 PM steel B. Denkena, T. Grove, H. Lucas ∗ Leibniz Universität Hannover, Institute for Production Engineering and Machine Tools, An der Universität 2, 30823, Garbsen, Germany
a r t i c l e
i n f o
Article history: Received 15 April 2015 Received in revised form 22 September 2015 Accepted 23 September 2015 Available online 9 October 2015 Keywords: Residual stresses Grinding Toric tools
a b s t r a c t Residual stresses have an important role concerning service life of forming tools. The grinding of the tool is one of the major manufacturing steps inducing residual stresses into the subsurface of the tool. Tool and process parameters for grinding with toric grinding pins have been investigated for their influences on residual stresses and surface finish. Residual stresses were measured using the sin2 method, while the surface finish was analyzed using white light confocal microscopy. Results show that the grinding strategy has a major influence on residual stress generation for both principle directions of the process. In addition cutting grain size has the major impact on residual stresses transverse to cutting direction, while feed rate has the main influence on residual stresses in cutting direction. A bigger grain size results in more compressive residual stresses, while a higher feed rate shifts stresses towards the tensile regime. A good surface finish is achieved with small cutting grain size, low feed rates and frontal grinding strategy. © 2015 Elsevier B.V. All rights reserved.
1. Introduction Sheet-bulk metal forming is a new production process which combines the advantages of deep drawing, upsetting and forging to manufacture complex parts out of sheet metal up to a thickness of 3 mm. Complex parts with e.g., small cogs can be manufactured without the need for intermediate heating. The lack of recrystallization or recovery leads to improved mechanical properties compared to hot worked parts, due to work hardening occurring in the process. As described by Merklein et al. (2011) the downside is the high process forces of up to 2100 kN which are needed to fill the different functional geometries of the designated parts. They further showed that during forming operations complex biaxial and triaxial stress and strain states occur inside the tools for sheet-bulk metal forming, which therefore must have especially high wear resistance. It is known that residual stresses in the subsurface have a great influence on crack initiation as well as propagation. Compressive residual stresses delay tool failure due to fatigue and enhance tool lifetime (Lange et al., 2008). Hence, it is necessary to optimize subsurface residual stresses to interact favorably with tool workloads to stop or at least slow down crack initiation and propagation. This is especially difficult since stress and strain states differ locally in
∗ Corresponding author. E-mail address: lucas [email protected] (H. Lucas). http://dx.doi.org/10.1016/j.jmatprotec.2015.09.039 0924-0136/© 2015 Elsevier B.V. All rights reserved.
the forming tool. Thus, a manufacturing process is needed which can safely apply the complex geometries needed for the forming tool while inducing locally adjusted residual stresses at the same time. Additionally, surface roughness may not be neglected. Surface roughness has a strong impact on material flow in sheet-bulk metal forming operations as it was shown by Hense et al. (2015). A manufacturing process, which induces locally optimized residual stresses after hardening of the tool and simultaneously generates an appropriate surface quality, is needed for that reason. Grinding with toric tools is a promising process for this application. Denkena and van der Meer (2009) described how constant contact conditions can be realized due to the two different radii of the toric tool while grinding ceramic free form surfaces. Because of the adapted contact conditions, material removal rates stay constant and therefore guarantee a consistent surface quality. However, residual stress generation was not investigated. Grinding of ductile workpiece materials like steel with toric tools was not yet investigated. Especially identifying tool and process influences to accurately predict residual stress generation while grinding steel with toric tools is the main goal of the ongoing research presented in this article. Main influences of tool properties (bonding type, grain size, grain concentration), process parameters (cutting velocity and feed rate), as well as grinding strategy influencing subsurface residual stresses and surface roughness will be identified. To determine the main influences of these parameters on the subsurface residual stresses and surface quality a fractional factorial design of experiment is applied. All parameters are varied
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Fig. 1. Frontal and lateral grinding strategies while grinding with toric tools.
Fig. 2. Different contact conditions as considered by Denkena et al. (2013) and adapted for makro roughness much smaller than the depth of cut.
in two steps, which are taken from both edges of reasonable parameter ranges. The results are evaluated statistically assuming a Student’s t-distribution, to identify the main influences. 2. Grinding with toric grinding tools Toric grinding tools have a small minor ring radius and a bigger torus diameter (see Fig. 1). The tools enable the use of two different tool radii without the need for tool changing. Small grooves can be ground with the minor ring radius nearly the same way as with ball end tools, while maintaining much higher cutting speed along the radius. Consequently, lower rotational speeds are necessary to achieve the same cutting speed. For high cutting speeds like 40 m/s, more than 76,000 min−1 would be needed for a ball end grinding pin with a radius of 5 mm. A toric tool with a main diameter of
T = 30 mm and a minor ring radius of r = 5 mm achieves this cutting speed already at about 25,500 min−1 . This enables the use of regular machine tools without special high speed spindles. Because of the geometry of the tools, two strategies must be distinguished and will be further described as frontal and lateral grinding. The strategies are depicted in Fig. 1. Using the frontal grinding strategy cutting direction and feed direction are transverse to each other. In comparison, if cutting direction and feed direction are parallel to each other, it is referred as lateral grinding. For both strategies, the tool may be inclined. For frontal grinding inclination is realized in feed direction via the lead angle ˇf , while for lateral grinding inclination is done transversely to feed direction via the tilt angle ˇfN . Contact conditions of the tool with the surface as well as force distribution are very different for the two strategies. For the
B. Denkena et al. / Journal of Materials Processing Technology 229 (2016) 541–548 Table 1 Contact conditions for lateral and frontal grinding strategies at ˇf = fN = 90◦ .
Lateral grinding (S) Frontal grinding (F)
Contact width bg [m]
Contact length lg [m]
224 474
387 273
lateral strategy, cutting force and feed force have the same direction, while for the frontal strategy they are transverse to each other. The contact length of the tool along the main diameter is considerably higher than along the minor radius of the torus. Denkena et al. (2013) described how to calculate contact length and width of the torus with the workpiece for both lateral and frontal grinding (see Eqs. (1) and (2) for ˇf = ˇfN = 90◦ ). Contact length and width are necessary factors to determine the theoretical macro roughness they also described. The theoretical macro roughness model includes only the geometry of the tool and the workpiece, as well as the width and depth of cut, but does not consider the tool microstructure. Depending on the overlapping of the single tool paths with each other the theoretical macro roughness is calculated between the lowest point in the valley of the path and the peaks in between two paths. bgS = 2lgF = 2lgS = bgF =
r arccos 90
r − a
T arccos 90
e
(1)
r
T − a e
(2)
T
However, the case of small path distances is not covered in their work. This additional case occurs if the path distance is small enough so that the theoretical macro roughness is smaller than the depth of cut (see Fig. 2). For the experiments in this article a depth of cut of ae = 5 m and a theoretical macro roughness of Rthmac = 0.25 m was chosen. Consequently, this case is applicable here. Denkena et al. (2013) considered for lateral grinding that the torus minor radius would be immerged for the full depth of cut on both sides of the tool path. However, for small tool path distances the previous path already cut most of the material on one side away. The same is applicable for the torus main diameter while using the frontal grinding strategy. The formulas used by them are still valid (see Eqs. (1) and (2) for ˇf = ˇfN = 90◦ ). Instead of doubling the circle arc for the contact width of frontal grinding and the contact length for lateral grinding, it is necessary to calculate the contact conditions on the already cut side of the tool path with a depth of cut equal to the theoretical macro roughness Rthmac = 0.25 m. Considering this, contact length and width for both strategies were calculated and are presented in Table 1. The contact length in cutting direction for both strategies is about twice as high as transverse to cutting direction. For tilt or lead angles ˇf /ˇfN <90◦ the difference gets bigger, since the circle arc over the torus main diameter distorts to a longer ellipse segment, while the circle arc over the minor radius does not change. The considerations concerning contact conditions and process force directions lead to the hypothesis, that the grinding strategy will have a major influence not only on surface quality but also on subsurface residual stresses.
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The grinding tools do not resemble a perfect torus shape, but the surface is composed of cutting grains sticking out of the bonding and the bonding matrix in between. Depending on cutting grains size, bonding type and concentration of cutting grains the surface topography of a grinding tool differs drastically. This surface reproduces itself on the surface of the workpiece. Therefore, they also described a theoretical micro roughness (Denkena et al., 2013). The micro roughness cannot be calculated easily. It either can be determined empirically through grinding experiments or calculated from the grinding tool surface topography and the kinematics of the grinding operation. Since it is time-consuming to measure the surface topography of every grinding tool used, while the topography may change fast during actual grinding experiments, this approach is not used here. Denkena et al. (2013) showed that for grinding ceramics with diamond tools a smaller grain size (dg = 15 m) and a ceramic bonding result in a better micro roughness compared to a grain size of dg = 46 m and electroplated tools. The same results can be expected to be valid for the experiments presented in this article.
3. Experimental setup To determine the main influences on surface generation and subsurface residual stresses the following six parameters were investigated: tool bonding type (B), cutting grain size (dg ), cutting grain concentration (C), grinding strategy (S), cutting speed (vc ) and feed rate (vf ). For each parameter two steps at both ends of the plausible parameter value range were investigated. For electroplated tools different cutting grain concentrations could not be investigated, since these are commercially not available. Manufacturers configure their electrolytic bathes to achieve a high concentration of cutting grains on the tool surface and cannot change these settings without disturbing the manufacturing process. A fractional factorial experimental design of the type 26–1 was applied in assumption that the grinding strategy and cutting grain size do not interact with each other. The interaction of cutting grain size and strategy is therefore substituted by the factor bonding type. In order to gain further insight into residual stress generation and to improve statistical data, experiments done with the frontal grinding strategy were later repeated with lateral strategy. Consequently, a full factorial experiment was conducted with use of all six different tools in lateral grinding.
3.1. Sample preparation Samples were cut from 1.3344 PM flat steel rods, plane ground and heat treated. The heat treatment was done for 2 h at 620 ◦ C with low heating and cooling rates, to relieve the samples of any preliminary residual stresses due to the preparation. On an exemplary specimen, surfaces scale was removed via electro polishing and residual stresses were measured via X-ray diffraction to confirm that no preliminary residual stresses were present in the samples before the grinding experiments.
Table 2 Grinding tool parameters. Bonding type
Cutting grain size dg [m]
Cutting grain concentration C
Acronym
Ceramic Ceramic Electro plated Ceramic Ceramic Electro plated
15 15 15 91 91 91
C50 C125 – C50 C125 –
15-C50 15-C125 15-g 91-C50 91-C125 91-g
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3.2. Machine tool and toric grinding pins Six different toric grinding tools were used. The tools are summarized in Table 2. The acronym of each tool will be used further in the article. All tools had a diameter of T = 30 mm and a radius of R = 5 mm. To take manufacturing variances into account each tool’s diameter and radius was measured individually via silhouette measurement with a Walter Helicheck WWM. The individual tool diameters and radii were taken into account for NC-programming of the experiments. The grinding experiments were conducted on a Röders RFM 600 DS machine tool. Cutting speed was varied on two levels of 15 m/s and 35 m/s, while feed rate was varied on levels of 50 mm/min and 2000 mm/min. The tool was cooled externally with five percent emulsion at two bar pressure. For each set of cutting parameters an area of approximately 4 mm × 20 mm was ground which conducts to four areas for each grinding pin. To take different initial surface heights into account each area was ground twice. The first passage to generate a new, flat surface with a depth of cut of approximately ae = 5 m, the second passage with an actual depth of cut of ae = 5 m from the newly generated surface. Path distance was calculated considering the theoretical macro roughness described by Denkena et al. (2013). A path distance of ab,F = 250 m for the frontal grinding strategy and ab,S = 100 m for the lateral grinding strategy result in a theoretical macro roughness of about 0.25 m for both strategies. The theoretical macro roughness of 0.25 m was chosen so that the micro roughness described by Denkena et al. (2013) would be the dominant factor for the actual roughness of the ground surface. 3.3. Surface and subsurface analysis Surface roughness was measured using a NanoFocus -surf white light confocal microscope. The measured area was analyzed with the software package “soft analysis premium 7.1” of Digital Surf SARL. An area of 2 × 2 mm is too short to fit five segments of 0.8 mm into it to calculate Rz roughness parameters in accordance to DIN 4760. The segments have therefore been shortened and the arithmetic mean out of five profiles has been calculated to be statistically verified. Residual stresses in the subsurface were measured by means of X-ray diffractometry with the sin2 -technique described in detail by Eigenmann and Macherauch (1995). Experiments were conducted using a General Electric XRD 3003 TT with Cr-anode and V-filter. The anode current was Ia = 35 mA with an acceleration voltage of Ua = 30 kV. The penetration depth of X-ray radiation at this energy is approximately 5.5 m depending on the inclination angle (Eigenmann and Macherauch, 1995). The measuring spot was 2 mm in diameter. The software package “Analyze” from General Electric is used to calculate the residual stresses in the subsurface. Background noise of the X-ray diffraction pattern was subtracted, the peaks were smoothed and parabolas were fitted into the peaks above a threshold of 70% of its height. Measurements were done parallel, transversely and diagonally to feed direction to be able to calculate the principal residual stresses in the subsurface from the three measurements via the method described by Macherauch and Müller (1961). 4. Experimental results 4.1. Surface roughness Surface roughness parameters for frontally ground samples were calculated parallel and transversely to feed direction. For laterally ground specimen the roughness parameters were only
Fig. 3. Roughness parameters Rz for surfaces ground with frontal grinding strategy. Table 3 Feed for all four experiment parameter combinations. Feed fF
vc1 = 15 m/s
vc2 = 35 m/s
vf1 = 50 mm/min Overlap vf2 = 2000 mm/min Overlap
0.005 mm 54.6 0.209 mm 1.3
0.002 136.5 0.090 3.0
calculated transversely to feed direction, since grinding grooves occur parallel to feed direction. Fig. 3 shows roughness values Rz for all frontally ground specimen. Comparing the different tools, roughness for the ceramic bonding tool with the small grain size of 15 m is minimal while the electroplated tool with a 91 m grain size results in high roughness values. For three experiments with a higher feed rate, chatter marks are visible on the surface. Li et al. (2002) showed that in contour grinding, in which cutting and feed direction are simmilar to the frontal grinding strategy, chatter is more likely to occur at high feed rates. Higher feed rates result in higher material removal rates, which lead to higher process forces and therefore a less stable process. Chatter also only occurred for tools with the bigger grain size of dg = 91 m and especially the electroplated tools. Bigger grains and higher grain protrusion also result in higher process forces and therefore more likely result in the observed chatter. The occurance of chatter induces very high roughness values for the three experiments. These experiments are not representative for general roughness generation. For each tool, the feed rate increase from vf = 50–2000 mm/min results in an at least fourfold increase of roughness values. This effect is dominant parallel to feed direction. The frontal grinding strategy generally results in a good surface quality, with the best results for the ceramically bonded tool with small grain size. In Fig. 4 the surfaces ground with the 91-C125 tool with cutting speed vc1 = 15 m/s and feed rate of vf1 = 50 m/min as well as the surface ground with the 91-g tool with vc2 = 35 m/s and vf2 = 2000 m/min are shown. For the low feed rate (Fig. 4a) the single tool path is clearly visible. Grinding grooves are oriented transversely to the tool path, while the tool moves over the surface in feed direction with no actual cutting in this orientation. For the high feed rate (Fig. 4b) the tool path is not visible anymore. Dominant features are feed marks at a distance of roughly 0,09 mm and overlapping chatter marks with a distance of roughly 1 mm. The higher roughness of the stable experiments with higher feed rates can be explained by calculating the feed (per revolution) for the four parameter combinations using Eq. (3). Results are presented in Table 3. The different ratios of the contact length
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Fig. 4. Exemplary surfaces for frontal grinding strategy: (a) 91-C125 vc1 = 15 m/s, vf1 = 50 mm/min (b) 91-C125 vc1 = 15 m/s, vf2 = 2000 mm/min.
calculated in chapter 2 (lgF = 0273 mm) and the different feeds vary in the order of two magnitudes (see overlap in Table 3). For frontal grinding different cutting grain topographies of the tool overlap on the surface of the workpiece, since the tool moves transversely to cutting direction. For the low feed rate, these topographies overlap at least 54 times on the surface. For the high feed rate they overlap three times at maximum. fF =
vf × × T ve
(3)
These results coincide with the findings of Denkena et al. (2013). They also showed the dependency of micro roughness and feed rate for the frontal grinding strategy as well as higher roughness values for electroplated tools in comparison to ceramically bonded ones. Reported values for the micro roughness are a lot higher than the ones presented in this article though. Values from Denkena et al. (2013) range from Rz = 4 m for tools with a cutting grains size of dg = 15 m to Rz = 10 m for tools with a cutting grain size of dg = 46 m. Not considering experiments where chatter occurred, the worst roughness values shown in this article are still better than the best values for the brittle ceramic workpieces presented by Denkena et al. (2013). For experiments conducted with the lateral strategy no chatter occurred. For the frontal grinding strategy the tool was “hopping” over the surface in feed direction, were no actual cutting occurred. For lateral grinding feed and cutting direction are the same. Consequently, there is no tool movement outside cutting direction over the surface in contrast to the frontal grinding strategy. The cutting speed for these experiments is at least three orders of magnitude higher than the feed rate and therefore not influencing cutting conditions. Roughness values for all tools show no clear connection to either feed rate or cutting speed (see Fig. 5). It has to be noted that roughness values for lateral grinding in general are much higher than for frontal grinding. Experiments with electroplated tools for both grain sizes show significantly higher roughness values than experiments conducted with ceramically bonded tools. For the lateral strategy tool topographies do not overlap each other, but the envelope of the tool depicts itself on the surface of the work-
Fig. 5. Roughness parameters Rz for surfaces ground with lateral grinding strategy.
piece. The higher protrusion of the grains in the electroplated tools therefore results in higher grinding grooves on the surface. Some experiments show outstanding high roughness values which result from single grains emerging especially high from the tool surface. Fig. 6 shows some examples of surface topography after lateral grinding. Cutting grooves are oriented along the tool path and as depicted in Fig. 6(b) especially deep grooves repeat themselves with tool path distance. These deep grooves are the result of single grains emerging comparatively high out of the bonding of their tool. For the lateral strategy these grains cut along the tool path direction, which results into a single, deep groove. This effect is mainly dependent on bonding type. In ceramic bonding cutting grains are deeply immerged into the bonding. Cutting grains of electroplated tools on the other hand emerge up to half of their size out of the bonding matrix. The roughness of surfaces ground with ceramic bonding and the higher grain size of 91 m is therefore lower than the roughness of surfaces ground with the electroplated tool with 15 m cutting grain size. The effect of the
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Fig. 6. Exemplary surfaces for lateral grinding strategy: (a) 15-C50 vc1 = 15 m/s, vf2 = 2000 mm/min, (b) 91-g vc1 = 15 m/s, vf2 = 2000 mm/min.
grain size for electroplated tools is not present for every parameter combination, but can exemplarily be determined by comparing the roughness values Rz for the 15-g vc1 = 15 m/s; vf1 = 50 mm/min grinding experiment with the 91-g vc2 = 35 m/s; vf1 = 50 mm/min experiment. While the other Rz values for both experiments were around 5–6 m, these two are considerably higher. In this case, a very prominent cutting grain that is highly emerged from the surface has cut especially deep grooves. This results in a roughness Rz of Rzg-15/vc1/vf1 = 8.96 m for the tool with 15 m cutting grain size and a roughness of Rzg-91/vc2/vf1 = 30.4 m. Single grains of 91 m size can emerge considerably higher from the bonding matrix and therefore worsen the surface quality considerably. Dependencies of surface roughness with feed, bonding type and cutting grain size for frontal grinding and bonding type and cutting grain size for lateral grinding could be verified for CBN tools in machining steel workpieces. Roughness values are lower for the ductile steel workpieces than for the brittle ceramic workpieces investigated by them. For a small cutting grain size and ceramic bonding, roughness values of up to Rz = 0.5 m could be achieved. Roughness values for lateral strategy are at least four times as high as for the frontal strategy. This factor is bigger as reported by Denkena and van der Meer. The ductile material removal allows a better smoothening of the surface from overlapping grinding paths in the frontal grinding strategy. 4.2. Subsurface residual stresses Principle residual stresses parallel to the workpiece surface have been calculated from three measurements with directions parallel 1 = 0◦ , diagonal 2 = 45◦ and transverse 3 = 90◦ to feed direction as described by Macherauch and Müller (1961). The values of the calculated principle residual stresses are all in 10 MPa range of the two measurements in feed direction and transverse to feed direction for each individual experiment. This means that the principle residual stress directions are the same as the principle directions of the grinding process. If this occurs the calculation of the angle in between original measurement and the principle residual stresses is largely defective. Very small differences in residual stresses
Fig. 7. Residual stresses parallel and transversal to feed direction for frontally ground specimen.
result in very large angles between the measured and calculated directions. Therefore, the original measurements are plotted and discussed in the following section. The first 24 experiments of the fractional factorial design were analyzed assuming a Student’s t-distribution for the measured residual stress values. Results showed that only the grinding strategy had a significant (99% confidence interval) effect on the residual stresses. For every other parameter the data was not sufficient. The fractional factorial experimental design was extended for the lateral strategy to a full factorial design, to further investigate relevant process and tool parameters. While analyzing the data it was discovered that considering the cutting direction instead of the feed direction as the principle orientation for residual stresses is more promising. The results for all experiments are summarized in Figs. 7–9 . All but one experiment generated compressive residual stresses in the subsurface for both principle directions. The exception
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Fig. 8. Residual stresses parallel and transversal to feed direction for specimen laterally ground with 15 m cutting grain size tools.
Fig. 9. Residual stresses parallel and transversal to feed direction for specimen laterally ground with 91 m cutting grain size tools.
are residual stresses in cutting direction for the experiment with the lateral grinding strategy the 91-C50 tool, vc2 = 35 m/s and vf2 = 2000 mm/min (see Fig. 9). Compressive residual stresses occur because of plastic deformations in the subsurface. Compressive residual stresses transverse to the cutting direction are generally higher than the stresses in cutting direction. In cutting direction the deformed material is removed as chips. Transverse to cutting direction the material is deformed while being pushed out of the cutting path of the tool, resulting in compressive residual stresses. Comparing the experiments with the same parameters except for cutting grain size from Figs. 8 and 9, compressive residual stresses for both directions are higher for each experiment with the higher grain size. Effects of the grain size on these residual stresses can be understood considering the single grain cutting depth hcu established by Lierse (1998). At the same concentration, a bigger grain size results in a lower amount of cutting grains at the tool surface. The single grain cuts much deeper into the surface and consequently deforms the subsurface to a much higher degree. The highest compressive residual stresses in between −400 MPa and −450 MPa for both strategies were found transverse to cutting direction with the 91 m grain size and either electroplated tools or ceramic tools with the higher concentration of cutting grains. Comparing residual stresses in cutting direction for both strategies with different feed rates, it can be found that with higher feed
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rate less compressive residual stresses occur. This can either be caused by less mechanical deformation of the subsurface or more generation of heat, which is involved in the generation of tensile residual stresses. A higher feed rate is accompanied by higher material removal rates. Choi (1986) showed that these higher material removal rates for both conventional aluminium oxide and CBN grinding wheels result in higher temperatures in the workpiece. Temperatures in the contact zone of CBN grinding wheels are generally lower, due to the higher thermal conductivity of CBN compared to aluminium oxide. This mostly results in tensile residual stresses in the subsurface for workpieces ground with aluminium oxide, while CBN ground workpieces mostly show compressive residual stresses. Mohlfeld (2000) reported the same for the grinding of cermets and cemented carbides. A higher feed rate results in less compressive stresses in the subsurface due to the higher amount of active cutting edges, resulting in higher thermal loads on the workpiece. He also reported lower values for compressive residual stresses for cermets in comparison to the cermented carbides, because of the lower thermal conductivity of cermets, resulting in even higher thermal stresses. Kruszynski and Wójcik proposed a grinding coefficient B, which is a factor of the power density and the contact time of the tool with the surface (Kruszynski and Wójcik, 2001). In contrast to the findings of Mohlfeld, the coefficient B decreases with higher feed rate resulting in lower thermal loads and therefore lower tensile residual stresses. However, their feed rates varied from 600 mm/min to 3000 mm/min and all of their experiments resulted in tensile residual stresses, which comes from a high heat generation during the process. For their experiments, a higher feed rate leads to a shorter contact time resulting in less heat energy transferred to the surface. Different cooling rates resulting from different feed rates were not considered. For the experiments conducted in this paper high temperature generation is unlikely, because of the use of CBN tools and the very small contact zone of the grinding pin with the surfaces. The cooling rate is likely to have a more prominent effect. At a feed rate of vf = 50 mm/min the tool moves slow enough to give the workpiece material time to cool while still in contact with the tool. On the other hand at the high feed rate of vf = 2000 mm/min the workpiece is cooled by the cutting fluid, while the tool has moved on. Investigations of the development of temperature in the subsurface zone will be necessary to confirm this hypothesis. After completion of the full factorial design the main influences were analyzed statistically for the lateral strategy, separately at first. The significance was therefore plotted in correlation to the cutting direction instead of the feed direction. For the lateral strategy, feed and cutting direction are parallel. For the frontal strategy, the cutting direction is transversal to the feed direction. Confidence intervals for 95%, 99% and 99.9% were calculated and t-values of each parameter were plotted against them to identify the significant and highly significant ones (see Fig. 10). The t-distribution has a degree of freedom of df = 18 and measured data was distributed close to normally. For residual stresses in cutting direction the feed rate is clearly the only highly significant parameter. While bonding type is close to the significant regime (99.0% confidence interval) not enough data was collected to clearly show a dependency between the residual stresses in cutting direction and the bonding type. For the lateral strategy, the feed rate is the main and only highly significant influence (t-value above 99.9% confidence interval) on residual stresses in cutting direction. Transversal to cutting direction the feed rate on the other hand has no significant influence. The only parameter identified in the significant regime (outside 99.0% confidence interval) is the cutting grain size of the tool. It is close to being highly significant. Based on these findings the significance for all experimental results have been plotted in Fig. 11, leaving out the grinding
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5. Conclusion
Fig. 10. Significance of tool and process parameters on residual stresses for lateral grinding strategy.
The grinding strategy has the highest impact on surface roughness as well as subsurface residual stresses. Since the kinematic contact is very different for both strategies, they need to be considered separately. Depending on the workpiece geometry both grinding strategies are needed, so while the frontal grinding strategy always results in better surface qualities, the lateral strategy cannot be neglected. For the surface roughness ceramic tools with a small grain size and high grain concentration are favorable. The small grains are immerged deeply in the vitrified bond resulting in a very smooth tool topography, which displays itself on the workpiece surface. Feed rate and cutting speed must be adjusted so that the tool feed is small enough that several tool topographies overlap and no chatter occurs. Otherwise the tool parameters have the dominant influence on surface roughness over the process parameters cutting speed and feed rate. For residual stresses two major influencing parameters have been identified, the feed rate and the cutting grain size. If considering the residual stresses main directions, parallel and transversal to cutting direction, these parameters are equally important for both strategies. Bigger grains result in a higher single grain cutting depth and therefore more plastic deformation transverse to cutting direction. A higher feed rate shifts residual stresses towards tensile values. To quantify the impact of the most influential parameters on residual stresses further investigations will be carried out while measuring workpiece temperature and cutting forces. Acknowledgement The work presented in this article was carried out in the subproject B8 “Grinding strategies for local and stress orientated subsurface-modification of sheet-bulk metal forming tools” of the Transregional Collaborative Research Centre on sheet-bulk metal forming (SFB/TR 73). The TR 73 is funded by the German Research Foundation (DFG).
Fig. 11. Significance of tool and process parameters on residual stresses for both grinding strategies.
strategy which was already established as highly significant. The t-distribution for all data has a degree of freedom of df 30. The correlation between feed rate and residual stresses in cutting direction stays highly significant for all experiments. For both strategies the feed rate must therefore be considered for residual stress development in cutting direction. The cutting speed comes close to a significant level, while the bonding type stays slightly below significance. For the residual stresses transversal to cutting direction the cutting grain size becomes even more significant by considering all experiments and is slightly above the threshold for the 99.9% regime. The cutting speed becomes more prominent for the residual stresses in this direction too, while its t-value is still slightly too low for the 99% significance level. For both strategies feed rate and cutting grain size show clear correlation to the residual stresses in cutting direction and transversal to cutting direction respectively. A large cutting grain size leads to higher compressive residual stresses, while a high feed rate shifts residual stresses towards the tensile area. For all experiments but one, stresses stay compressive though. Considering all data available, the cutting speed might have an impact on residual stresses for both strategies, only the frontal strategy or not at all. Further investigations are needed to clearly identify or dismiss an impact.
References Choi, Hon-Zong, 1986. Beitrag zur Ursachenanalyse der Randzonenbeeinflussung beim Schleifen. In: Dr. -Ing. Dissertation. Universität Hannover. Denkena, B., van der Meer, M., 2009. Diamond tools for the grinding of complex ceramic implant surfaces. Adv. Mater. Res. 76–78, 33–37. Denkena, B., Köhler, J., van der Meer, M., 2013. A roughness model for the machining of biomedical ceramics by toric grinding pins. CIRP J. Manuf. Sci. Technol. 6, 22–33. Eigenmann, B., Macherauch, E., 1995. Röntgenographische Untersuchungen von Spannungszuständen in Werkstoffen, Teil I. Materialwissenschaft und Werkstoffstechnik 26 (3), 148–160. Hense, R., Wels, C., Kersting, P., Vierzigmann, U., Löffler, M., Biermann, D., Merklein, M., 2015. High-feed milling of tailored surfaces for sheet-bulk metal forming tools. Product. Eng 9 (2), 215–223. Kruszynski, B.W., Wójcik, R., 2001. Residual stress in grinding. J. Mater. Process. Technol. 109, 254–257. Lange, K., Kammerer, M., Pöhlandt, K., Schöck, J., 2008. Fließpressen, 1. Edn. Springer, Berlin. Li, Y., Gracewski, S.M., Funkenbusch, P.D., Ruckman, J., 2002. Analysis of chatter in contour grinding of optical materials. Int. J. Mach. Tools Manuf. 42 (10), 1095–1103. Lierse, T., 1998. Mechanische und thermische Wirkungen beim Schleifen keramischer Werkstoffe. In: Dr. -Ing. Dissertation. Leibniz Universität Hannover. Macherauch, E., Müller, P., 1961. Das sin2 -Verfahren der röntgenographischen Spannungsmessung. Zeitschrift för angewandte Physik 13 (7), 305. Merklein, M., Koch, J., Opel, S., Schneider, T., 2011. Fundamental investigations on the material flow at combined sheet and bulk metal forming processes. CIRP Annals-Manuf. Technol. 60, 283–286. Mohlfeld, A., 2000. Trockenbohren mit PVD-beschichteten Hartmetallwerkzeugen. In: Dr. -Ing. Dissertation. Universität Hannover.
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Journal of Materials Processing Technology journal homepage: www.elsevier.com/locate/jmatprotec
Influences of grinding with Toric CBN grinding tools on surface and subsurface of 1.3344 PM steel B. Denkena, T. Grove, H. Lucas ∗ Leibniz Universität Hannover, Institute for Production Engineering and Machine Tools, An der Universität 2, 30823, Garbsen, Germany
a r t i c l e
i n f o
Article history: Received 15 April 2015 Received in revised form 22 September 2015 Accepted 23 September 2015 Available online 9 October 2015 Keywords: Residual stresses Grinding Toric tools
a b s t r a c t Residual stresses have an important role concerning service life of forming tools. The grinding of the tool is one of the major manufacturing steps inducing residual stresses into the subsurface of the tool. Tool and process parameters for grinding with toric grinding pins have been investigated for their influences on residual stresses and surface finish. Residual stresses were measured using the sin2 method, while the surface finish was analyzed using white light confocal microscopy. Results show that the grinding strategy has a major influence on residual stress generation for both principle directions of the process. In addition cutting grain size has the major impact on residual stresses transverse to cutting direction, while feed rate has the main influence on residual stresses in cutting direction. A bigger grain size results in more compressive residual stresses, while a higher feed rate shifts stresses towards the tensile regime. A good surface finish is achieved with small cutting grain size, low feed rates and frontal grinding strategy. © 2015 Elsevier B.V. All rights reserved.
1. Introduction Sheet-bulk metal forming is a new production process which combines the advantages of deep drawing, upsetting and forging to manufacture complex parts out of sheet metal up to a thickness of 3 mm. Complex parts with e.g., small cogs can be manufactured without the need for intermediate heating. The lack of recrystallization or recovery leads to improved mechanical properties compared to hot worked parts, due to work hardening occurring in the process. As described by Merklein et al. (2011) the downside is the high process forces of up to 2100 kN which are needed to fill the different functional geometries of the designated parts. They further showed that during forming operations complex biaxial and triaxial stress and strain states occur inside the tools for sheet-bulk metal forming, which therefore must have especially high wear resistance. It is known that residual stresses in the subsurface have a great influence on crack initiation as well as propagation. Compressive residual stresses delay tool failure due to fatigue and enhance tool lifetime (Lange et al., 2008). Hence, it is necessary to optimize subsurface residual stresses to interact favorably with tool workloads to stop or at least slow down crack initiation and propagation. This is especially difficult since stress and strain states differ locally in
∗ Corresponding author. E-mail address: lucas [email protected] (H. Lucas). http://dx.doi.org/10.1016/j.jmatprotec.2015.09.039 0924-0136/© 2015 Elsevier B.V. All rights reserved.
the forming tool. Thus, a manufacturing process is needed which can safely apply the complex geometries needed for the forming tool while inducing locally adjusted residual stresses at the same time. Additionally, surface roughness may not be neglected. Surface roughness has a strong impact on material flow in sheet-bulk metal forming operations as it was shown by Hense et al. (2015). A manufacturing process, which induces locally optimized residual stresses after hardening of the tool and simultaneously generates an appropriate surface quality, is needed for that reason. Grinding with toric tools is a promising process for this application. Denkena and van der Meer (2009) described how constant contact conditions can be realized due to the two different radii of the toric tool while grinding ceramic free form surfaces. Because of the adapted contact conditions, material removal rates stay constant and therefore guarantee a consistent surface quality. However, residual stress generation was not investigated. Grinding of ductile workpiece materials like steel with toric tools was not yet investigated. Especially identifying tool and process influences to accurately predict residual stress generation while grinding steel with toric tools is the main goal of the ongoing research presented in this article. Main influences of tool properties (bonding type, grain size, grain concentration), process parameters (cutting velocity and feed rate), as well as grinding strategy influencing subsurface residual stresses and surface roughness will be identified. To determine the main influences of these parameters on the subsurface residual stresses and surface quality a fractional factorial design of experiment is applied. All parameters are varied
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Fig. 1. Frontal and lateral grinding strategies while grinding with toric tools.
Fig. 2. Different contact conditions as considered by Denkena et al. (2013) and adapted for makro roughness much smaller than the depth of cut.
in two steps, which are taken from both edges of reasonable parameter ranges. The results are evaluated statistically assuming a Student’s t-distribution, to identify the main influences. 2. Grinding with toric grinding tools Toric grinding tools have a small minor ring radius and a bigger torus diameter (see Fig. 1). The tools enable the use of two different tool radii without the need for tool changing. Small grooves can be ground with the minor ring radius nearly the same way as with ball end tools, while maintaining much higher cutting speed along the radius. Consequently, lower rotational speeds are necessary to achieve the same cutting speed. For high cutting speeds like 40 m/s, more than 76,000 min−1 would be needed for a ball end grinding pin with a radius of 5 mm. A toric tool with a main diameter of
T = 30 mm and a minor ring radius of r = 5 mm achieves this cutting speed already at about 25,500 min−1 . This enables the use of regular machine tools without special high speed spindles. Because of the geometry of the tools, two strategies must be distinguished and will be further described as frontal and lateral grinding. The strategies are depicted in Fig. 1. Using the frontal grinding strategy cutting direction and feed direction are transverse to each other. In comparison, if cutting direction and feed direction are parallel to each other, it is referred as lateral grinding. For both strategies, the tool may be inclined. For frontal grinding inclination is realized in feed direction via the lead angle ˇf , while for lateral grinding inclination is done transversely to feed direction via the tilt angle ˇfN . Contact conditions of the tool with the surface as well as force distribution are very different for the two strategies. For the
B. Denkena et al. / Journal of Materials Processing Technology 229 (2016) 541–548 Table 1 Contact conditions for lateral and frontal grinding strategies at ˇf = fN = 90◦ .
Lateral grinding (S) Frontal grinding (F)
Contact width bg [m]
Contact length lg [m]
224 474
387 273
lateral strategy, cutting force and feed force have the same direction, while for the frontal strategy they are transverse to each other. The contact length of the tool along the main diameter is considerably higher than along the minor radius of the torus. Denkena et al. (2013) described how to calculate contact length and width of the torus with the workpiece for both lateral and frontal grinding (see Eqs. (1) and (2) for ˇf = ˇfN = 90◦ ). Contact length and width are necessary factors to determine the theoretical macro roughness they also described. The theoretical macro roughness model includes only the geometry of the tool and the workpiece, as well as the width and depth of cut, but does not consider the tool microstructure. Depending on the overlapping of the single tool paths with each other the theoretical macro roughness is calculated between the lowest point in the valley of the path and the peaks in between two paths. bgS = 2lgF = 2lgS = bgF =
r arccos 90
r − a
T arccos 90
e
(1)
r
T − a e
(2)
T
However, the case of small path distances is not covered in their work. This additional case occurs if the path distance is small enough so that the theoretical macro roughness is smaller than the depth of cut (see Fig. 2). For the experiments in this article a depth of cut of ae = 5 m and a theoretical macro roughness of Rthmac = 0.25 m was chosen. Consequently, this case is applicable here. Denkena et al. (2013) considered for lateral grinding that the torus minor radius would be immerged for the full depth of cut on both sides of the tool path. However, for small tool path distances the previous path already cut most of the material on one side away. The same is applicable for the torus main diameter while using the frontal grinding strategy. The formulas used by them are still valid (see Eqs. (1) and (2) for ˇf = ˇfN = 90◦ ). Instead of doubling the circle arc for the contact width of frontal grinding and the contact length for lateral grinding, it is necessary to calculate the contact conditions on the already cut side of the tool path with a depth of cut equal to the theoretical macro roughness Rthmac = 0.25 m. Considering this, contact length and width for both strategies were calculated and are presented in Table 1. The contact length in cutting direction for both strategies is about twice as high as transverse to cutting direction. For tilt or lead angles ˇf /ˇfN <90◦ the difference gets bigger, since the circle arc over the torus main diameter distorts to a longer ellipse segment, while the circle arc over the minor radius does not change. The considerations concerning contact conditions and process force directions lead to the hypothesis, that the grinding strategy will have a major influence not only on surface quality but also on subsurface residual stresses.
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The grinding tools do not resemble a perfect torus shape, but the surface is composed of cutting grains sticking out of the bonding and the bonding matrix in between. Depending on cutting grains size, bonding type and concentration of cutting grains the surface topography of a grinding tool differs drastically. This surface reproduces itself on the surface of the workpiece. Therefore, they also described a theoretical micro roughness (Denkena et al., 2013). The micro roughness cannot be calculated easily. It either can be determined empirically through grinding experiments or calculated from the grinding tool surface topography and the kinematics of the grinding operation. Since it is time-consuming to measure the surface topography of every grinding tool used, while the topography may change fast during actual grinding experiments, this approach is not used here. Denkena et al. (2013) showed that for grinding ceramics with diamond tools a smaller grain size (dg = 15 m) and a ceramic bonding result in a better micro roughness compared to a grain size of dg = 46 m and electroplated tools. The same results can be expected to be valid for the experiments presented in this article.
3. Experimental setup To determine the main influences on surface generation and subsurface residual stresses the following six parameters were investigated: tool bonding type (B), cutting grain size (dg ), cutting grain concentration (C), grinding strategy (S), cutting speed (vc ) and feed rate (vf ). For each parameter two steps at both ends of the plausible parameter value range were investigated. For electroplated tools different cutting grain concentrations could not be investigated, since these are commercially not available. Manufacturers configure their electrolytic bathes to achieve a high concentration of cutting grains on the tool surface and cannot change these settings without disturbing the manufacturing process. A fractional factorial experimental design of the type 26–1 was applied in assumption that the grinding strategy and cutting grain size do not interact with each other. The interaction of cutting grain size and strategy is therefore substituted by the factor bonding type. In order to gain further insight into residual stress generation and to improve statistical data, experiments done with the frontal grinding strategy were later repeated with lateral strategy. Consequently, a full factorial experiment was conducted with use of all six different tools in lateral grinding.
3.1. Sample preparation Samples were cut from 1.3344 PM flat steel rods, plane ground and heat treated. The heat treatment was done for 2 h at 620 ◦ C with low heating and cooling rates, to relieve the samples of any preliminary residual stresses due to the preparation. On an exemplary specimen, surfaces scale was removed via electro polishing and residual stresses were measured via X-ray diffraction to confirm that no preliminary residual stresses were present in the samples before the grinding experiments.
Table 2 Grinding tool parameters. Bonding type
Cutting grain size dg [m]
Cutting grain concentration C
Acronym
Ceramic Ceramic Electro plated Ceramic Ceramic Electro plated
15 15 15 91 91 91
C50 C125 – C50 C125 –
15-C50 15-C125 15-g 91-C50 91-C125 91-g
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3.2. Machine tool and toric grinding pins Six different toric grinding tools were used. The tools are summarized in Table 2. The acronym of each tool will be used further in the article. All tools had a diameter of T = 30 mm and a radius of R = 5 mm. To take manufacturing variances into account each tool’s diameter and radius was measured individually via silhouette measurement with a Walter Helicheck WWM. The individual tool diameters and radii were taken into account for NC-programming of the experiments. The grinding experiments were conducted on a Röders RFM 600 DS machine tool. Cutting speed was varied on two levels of 15 m/s and 35 m/s, while feed rate was varied on levels of 50 mm/min and 2000 mm/min. The tool was cooled externally with five percent emulsion at two bar pressure. For each set of cutting parameters an area of approximately 4 mm × 20 mm was ground which conducts to four areas for each grinding pin. To take different initial surface heights into account each area was ground twice. The first passage to generate a new, flat surface with a depth of cut of approximately ae = 5 m, the second passage with an actual depth of cut of ae = 5 m from the newly generated surface. Path distance was calculated considering the theoretical macro roughness described by Denkena et al. (2013). A path distance of ab,F = 250 m for the frontal grinding strategy and ab,S = 100 m for the lateral grinding strategy result in a theoretical macro roughness of about 0.25 m for both strategies. The theoretical macro roughness of 0.25 m was chosen so that the micro roughness described by Denkena et al. (2013) would be the dominant factor for the actual roughness of the ground surface. 3.3. Surface and subsurface analysis Surface roughness was measured using a NanoFocus -surf white light confocal microscope. The measured area was analyzed with the software package “soft analysis premium 7.1” of Digital Surf SARL. An area of 2 × 2 mm is too short to fit five segments of 0.8 mm into it to calculate Rz roughness parameters in accordance to DIN 4760. The segments have therefore been shortened and the arithmetic mean out of five profiles has been calculated to be statistically verified. Residual stresses in the subsurface were measured by means of X-ray diffractometry with the sin2 -technique described in detail by Eigenmann and Macherauch (1995). Experiments were conducted using a General Electric XRD 3003 TT with Cr-anode and V-filter. The anode current was Ia = 35 mA with an acceleration voltage of Ua = 30 kV. The penetration depth of X-ray radiation at this energy is approximately 5.5 m depending on the inclination angle (Eigenmann and Macherauch, 1995). The measuring spot was 2 mm in diameter. The software package “Analyze” from General Electric is used to calculate the residual stresses in the subsurface. Background noise of the X-ray diffraction pattern was subtracted, the peaks were smoothed and parabolas were fitted into the peaks above a threshold of 70% of its height. Measurements were done parallel, transversely and diagonally to feed direction to be able to calculate the principal residual stresses in the subsurface from the three measurements via the method described by Macherauch and Müller (1961). 4. Experimental results 4.1. Surface roughness Surface roughness parameters for frontally ground samples were calculated parallel and transversely to feed direction. For laterally ground specimen the roughness parameters were only
Fig. 3. Roughness parameters Rz for surfaces ground with frontal grinding strategy. Table 3 Feed for all four experiment parameter combinations. Feed fF
vc1 = 15 m/s
vc2 = 35 m/s
vf1 = 50 mm/min Overlap vf2 = 2000 mm/min Overlap
0.005 mm 54.6 0.209 mm 1.3
0.002 136.5 0.090 3.0
calculated transversely to feed direction, since grinding grooves occur parallel to feed direction. Fig. 3 shows roughness values Rz for all frontally ground specimen. Comparing the different tools, roughness for the ceramic bonding tool with the small grain size of 15 m is minimal while the electroplated tool with a 91 m grain size results in high roughness values. For three experiments with a higher feed rate, chatter marks are visible on the surface. Li et al. (2002) showed that in contour grinding, in which cutting and feed direction are simmilar to the frontal grinding strategy, chatter is more likely to occur at high feed rates. Higher feed rates result in higher material removal rates, which lead to higher process forces and therefore a less stable process. Chatter also only occurred for tools with the bigger grain size of dg = 91 m and especially the electroplated tools. Bigger grains and higher grain protrusion also result in higher process forces and therefore more likely result in the observed chatter. The occurance of chatter induces very high roughness values for the three experiments. These experiments are not representative for general roughness generation. For each tool, the feed rate increase from vf = 50–2000 mm/min results in an at least fourfold increase of roughness values. This effect is dominant parallel to feed direction. The frontal grinding strategy generally results in a good surface quality, with the best results for the ceramically bonded tool with small grain size. In Fig. 4 the surfaces ground with the 91-C125 tool with cutting speed vc1 = 15 m/s and feed rate of vf1 = 50 m/min as well as the surface ground with the 91-g tool with vc2 = 35 m/s and vf2 = 2000 m/min are shown. For the low feed rate (Fig. 4a) the single tool path is clearly visible. Grinding grooves are oriented transversely to the tool path, while the tool moves over the surface in feed direction with no actual cutting in this orientation. For the high feed rate (Fig. 4b) the tool path is not visible anymore. Dominant features are feed marks at a distance of roughly 0,09 mm and overlapping chatter marks with a distance of roughly 1 mm. The higher roughness of the stable experiments with higher feed rates can be explained by calculating the feed (per revolution) for the four parameter combinations using Eq. (3). Results are presented in Table 3. The different ratios of the contact length
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Fig. 4. Exemplary surfaces for frontal grinding strategy: (a) 91-C125 vc1 = 15 m/s, vf1 = 50 mm/min (b) 91-C125 vc1 = 15 m/s, vf2 = 2000 mm/min.
calculated in chapter 2 (lgF = 0273 mm) and the different feeds vary in the order of two magnitudes (see overlap in Table 3). For frontal grinding different cutting grain topographies of the tool overlap on the surface of the workpiece, since the tool moves transversely to cutting direction. For the low feed rate, these topographies overlap at least 54 times on the surface. For the high feed rate they overlap three times at maximum. fF =
vf × × T ve
(3)
These results coincide with the findings of Denkena et al. (2013). They also showed the dependency of micro roughness and feed rate for the frontal grinding strategy as well as higher roughness values for electroplated tools in comparison to ceramically bonded ones. Reported values for the micro roughness are a lot higher than the ones presented in this article though. Values from Denkena et al. (2013) range from Rz = 4 m for tools with a cutting grains size of dg = 15 m to Rz = 10 m for tools with a cutting grain size of dg = 46 m. Not considering experiments where chatter occurred, the worst roughness values shown in this article are still better than the best values for the brittle ceramic workpieces presented by Denkena et al. (2013). For experiments conducted with the lateral strategy no chatter occurred. For the frontal grinding strategy the tool was “hopping” over the surface in feed direction, were no actual cutting occurred. For lateral grinding feed and cutting direction are the same. Consequently, there is no tool movement outside cutting direction over the surface in contrast to the frontal grinding strategy. The cutting speed for these experiments is at least three orders of magnitude higher than the feed rate and therefore not influencing cutting conditions. Roughness values for all tools show no clear connection to either feed rate or cutting speed (see Fig. 5). It has to be noted that roughness values for lateral grinding in general are much higher than for frontal grinding. Experiments with electroplated tools for both grain sizes show significantly higher roughness values than experiments conducted with ceramically bonded tools. For the lateral strategy tool topographies do not overlap each other, but the envelope of the tool depicts itself on the surface of the work-
Fig. 5. Roughness parameters Rz for surfaces ground with lateral grinding strategy.
piece. The higher protrusion of the grains in the electroplated tools therefore results in higher grinding grooves on the surface. Some experiments show outstanding high roughness values which result from single grains emerging especially high from the tool surface. Fig. 6 shows some examples of surface topography after lateral grinding. Cutting grooves are oriented along the tool path and as depicted in Fig. 6(b) especially deep grooves repeat themselves with tool path distance. These deep grooves are the result of single grains emerging comparatively high out of the bonding of their tool. For the lateral strategy these grains cut along the tool path direction, which results into a single, deep groove. This effect is mainly dependent on bonding type. In ceramic bonding cutting grains are deeply immerged into the bonding. Cutting grains of electroplated tools on the other hand emerge up to half of their size out of the bonding matrix. The roughness of surfaces ground with ceramic bonding and the higher grain size of 91 m is therefore lower than the roughness of surfaces ground with the electroplated tool with 15 m cutting grain size. The effect of the
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Fig. 6. Exemplary surfaces for lateral grinding strategy: (a) 15-C50 vc1 = 15 m/s, vf2 = 2000 mm/min, (b) 91-g vc1 = 15 m/s, vf2 = 2000 mm/min.
grain size for electroplated tools is not present for every parameter combination, but can exemplarily be determined by comparing the roughness values Rz for the 15-g vc1 = 15 m/s; vf1 = 50 mm/min grinding experiment with the 91-g vc2 = 35 m/s; vf1 = 50 mm/min experiment. While the other Rz values for both experiments were around 5–6 m, these two are considerably higher. In this case, a very prominent cutting grain that is highly emerged from the surface has cut especially deep grooves. This results in a roughness Rz of Rzg-15/vc1/vf1 = 8.96 m for the tool with 15 m cutting grain size and a roughness of Rzg-91/vc2/vf1 = 30.4 m. Single grains of 91 m size can emerge considerably higher from the bonding matrix and therefore worsen the surface quality considerably. Dependencies of surface roughness with feed, bonding type and cutting grain size for frontal grinding and bonding type and cutting grain size for lateral grinding could be verified for CBN tools in machining steel workpieces. Roughness values are lower for the ductile steel workpieces than for the brittle ceramic workpieces investigated by them. For a small cutting grain size and ceramic bonding, roughness values of up to Rz = 0.5 m could be achieved. Roughness values for lateral strategy are at least four times as high as for the frontal strategy. This factor is bigger as reported by Denkena and van der Meer. The ductile material removal allows a better smoothening of the surface from overlapping grinding paths in the frontal grinding strategy. 4.2. Subsurface residual stresses Principle residual stresses parallel to the workpiece surface have been calculated from three measurements with directions parallel 1 = 0◦ , diagonal 2 = 45◦ and transverse 3 = 90◦ to feed direction as described by Macherauch and Müller (1961). The values of the calculated principle residual stresses are all in 10 MPa range of the two measurements in feed direction and transverse to feed direction for each individual experiment. This means that the principle residual stress directions are the same as the principle directions of the grinding process. If this occurs the calculation of the angle in between original measurement and the principle residual stresses is largely defective. Very small differences in residual stresses
Fig. 7. Residual stresses parallel and transversal to feed direction for frontally ground specimen.
result in very large angles between the measured and calculated directions. Therefore, the original measurements are plotted and discussed in the following section. The first 24 experiments of the fractional factorial design were analyzed assuming a Student’s t-distribution for the measured residual stress values. Results showed that only the grinding strategy had a significant (99% confidence interval) effect on the residual stresses. For every other parameter the data was not sufficient. The fractional factorial experimental design was extended for the lateral strategy to a full factorial design, to further investigate relevant process and tool parameters. While analyzing the data it was discovered that considering the cutting direction instead of the feed direction as the principle orientation for residual stresses is more promising. The results for all experiments are summarized in Figs. 7–9 . All but one experiment generated compressive residual stresses in the subsurface for both principle directions. The exception
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Fig. 8. Residual stresses parallel and transversal to feed direction for specimen laterally ground with 15 m cutting grain size tools.
Fig. 9. Residual stresses parallel and transversal to feed direction for specimen laterally ground with 91 m cutting grain size tools.
are residual stresses in cutting direction for the experiment with the lateral grinding strategy the 91-C50 tool, vc2 = 35 m/s and vf2 = 2000 mm/min (see Fig. 9). Compressive residual stresses occur because of plastic deformations in the subsurface. Compressive residual stresses transverse to the cutting direction are generally higher than the stresses in cutting direction. In cutting direction the deformed material is removed as chips. Transverse to cutting direction the material is deformed while being pushed out of the cutting path of the tool, resulting in compressive residual stresses. Comparing the experiments with the same parameters except for cutting grain size from Figs. 8 and 9, compressive residual stresses for both directions are higher for each experiment with the higher grain size. Effects of the grain size on these residual stresses can be understood considering the single grain cutting depth hcu established by Lierse (1998). At the same concentration, a bigger grain size results in a lower amount of cutting grains at the tool surface. The single grain cuts much deeper into the surface and consequently deforms the subsurface to a much higher degree. The highest compressive residual stresses in between −400 MPa and −450 MPa for both strategies were found transverse to cutting direction with the 91 m grain size and either electroplated tools or ceramic tools with the higher concentration of cutting grains. Comparing residual stresses in cutting direction for both strategies with different feed rates, it can be found that with higher feed
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rate less compressive residual stresses occur. This can either be caused by less mechanical deformation of the subsurface or more generation of heat, which is involved in the generation of tensile residual stresses. A higher feed rate is accompanied by higher material removal rates. Choi (1986) showed that these higher material removal rates for both conventional aluminium oxide and CBN grinding wheels result in higher temperatures in the workpiece. Temperatures in the contact zone of CBN grinding wheels are generally lower, due to the higher thermal conductivity of CBN compared to aluminium oxide. This mostly results in tensile residual stresses in the subsurface for workpieces ground with aluminium oxide, while CBN ground workpieces mostly show compressive residual stresses. Mohlfeld (2000) reported the same for the grinding of cermets and cemented carbides. A higher feed rate results in less compressive stresses in the subsurface due to the higher amount of active cutting edges, resulting in higher thermal loads on the workpiece. He also reported lower values for compressive residual stresses for cermets in comparison to the cermented carbides, because of the lower thermal conductivity of cermets, resulting in even higher thermal stresses. Kruszynski and Wójcik proposed a grinding coefficient B, which is a factor of the power density and the contact time of the tool with the surface (Kruszynski and Wójcik, 2001). In contrast to the findings of Mohlfeld, the coefficient B decreases with higher feed rate resulting in lower thermal loads and therefore lower tensile residual stresses. However, their feed rates varied from 600 mm/min to 3000 mm/min and all of their experiments resulted in tensile residual stresses, which comes from a high heat generation during the process. For their experiments, a higher feed rate leads to a shorter contact time resulting in less heat energy transferred to the surface. Different cooling rates resulting from different feed rates were not considered. For the experiments conducted in this paper high temperature generation is unlikely, because of the use of CBN tools and the very small contact zone of the grinding pin with the surfaces. The cooling rate is likely to have a more prominent effect. At a feed rate of vf = 50 mm/min the tool moves slow enough to give the workpiece material time to cool while still in contact with the tool. On the other hand at the high feed rate of vf = 2000 mm/min the workpiece is cooled by the cutting fluid, while the tool has moved on. Investigations of the development of temperature in the subsurface zone will be necessary to confirm this hypothesis. After completion of the full factorial design the main influences were analyzed statistically for the lateral strategy, separately at first. The significance was therefore plotted in correlation to the cutting direction instead of the feed direction. For the lateral strategy, feed and cutting direction are parallel. For the frontal strategy, the cutting direction is transversal to the feed direction. Confidence intervals for 95%, 99% and 99.9% were calculated and t-values of each parameter were plotted against them to identify the significant and highly significant ones (see Fig. 10). The t-distribution has a degree of freedom of df = 18 and measured data was distributed close to normally. For residual stresses in cutting direction the feed rate is clearly the only highly significant parameter. While bonding type is close to the significant regime (99.0% confidence interval) not enough data was collected to clearly show a dependency between the residual stresses in cutting direction and the bonding type. For the lateral strategy, the feed rate is the main and only highly significant influence (t-value above 99.9% confidence interval) on residual stresses in cutting direction. Transversal to cutting direction the feed rate on the other hand has no significant influence. The only parameter identified in the significant regime (outside 99.0% confidence interval) is the cutting grain size of the tool. It is close to being highly significant. Based on these findings the significance for all experimental results have been plotted in Fig. 11, leaving out the grinding
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5. Conclusion
Fig. 10. Significance of tool and process parameters on residual stresses for lateral grinding strategy.
The grinding strategy has the highest impact on surface roughness as well as subsurface residual stresses. Since the kinematic contact is very different for both strategies, they need to be considered separately. Depending on the workpiece geometry both grinding strategies are needed, so while the frontal grinding strategy always results in better surface qualities, the lateral strategy cannot be neglected. For the surface roughness ceramic tools with a small grain size and high grain concentration are favorable. The small grains are immerged deeply in the vitrified bond resulting in a very smooth tool topography, which displays itself on the workpiece surface. Feed rate and cutting speed must be adjusted so that the tool feed is small enough that several tool topographies overlap and no chatter occurs. Otherwise the tool parameters have the dominant influence on surface roughness over the process parameters cutting speed and feed rate. For residual stresses two major influencing parameters have been identified, the feed rate and the cutting grain size. If considering the residual stresses main directions, parallel and transversal to cutting direction, these parameters are equally important for both strategies. Bigger grains result in a higher single grain cutting depth and therefore more plastic deformation transverse to cutting direction. A higher feed rate shifts residual stresses towards tensile values. To quantify the impact of the most influential parameters on residual stresses further investigations will be carried out while measuring workpiece temperature and cutting forces. Acknowledgement The work presented in this article was carried out in the subproject B8 “Grinding strategies for local and stress orientated subsurface-modification of sheet-bulk metal forming tools” of the Transregional Collaborative Research Centre on sheet-bulk metal forming (SFB/TR 73). The TR 73 is funded by the German Research Foundation (DFG).
Fig. 11. Significance of tool and process parameters on residual stresses for both grinding strategies.
strategy which was already established as highly significant. The t-distribution for all data has a degree of freedom of df 30. The correlation between feed rate and residual stresses in cutting direction stays highly significant for all experiments. For both strategies the feed rate must therefore be considered for residual stress development in cutting direction. The cutting speed comes close to a significant level, while the bonding type stays slightly below significance. For the residual stresses transversal to cutting direction the cutting grain size becomes even more significant by considering all experiments and is slightly above the threshold for the 99.9% regime. The cutting speed becomes more prominent for the residual stresses in this direction too, while its t-value is still slightly too low for the 99% significance level. For both strategies feed rate and cutting grain size show clear correlation to the residual stresses in cutting direction and transversal to cutting direction respectively. A large cutting grain size leads to higher compressive residual stresses, while a high feed rate shifts residual stresses towards the tensile area. For all experiments but one, stresses stay compressive though. Considering all data available, the cutting speed might have an impact on residual stresses for both strategies, only the frontal strategy or not at all. Further investigations are needed to clearly identify or dismiss an impact.
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