Creep-feed grinding of tungsten carbide using small diameter electroplated diamond wheels

Creep-feed grinding of tungsten carbide using small diameter electroplated diamond wheels This paper, by Z. Shi, H. Attia, D. Chellan and T. Wang, is concerned with an experimental investigation into creep-feed grinding of tungsten carbide using electroplated diamond wheels smaller than 25.4 mm in diameter. This work was motivated by applying plated diamond wheels to the grinding of deep slots or grooves in tungsten carbide components accessible only by small wheels. The objective is to explore the feasibility to grind tungsten carbides at large depths of cut, reasonable removal rates, and wheel life with small plated diamond wheels and to obtain some practical process parameters. Straight surface grinding experiments were conducted on tungsten carbide bars over a wide range of grinding conditions with waterbased grinding fluid. Grinding power, forces, workpiece surface roughness, and radial wheel wear were measured. Depths of cut as high as 4 mm corresponding to a specific material removal rate of 3.2 mm 2 /s were achieved with a wheel of 19 mm in diameter. Experimental results for specific energy, forces, and surface roughness are discussed in detail. It was shown that small plated diamond wheels are capable of grinding deep slots or grooves in tungsten carbides efficiently. Tungsten carbide materials possess extraordinary physical and mechanical properties such as high hardness and rigidity, high compressive strength, and excellent resistance to wear and heat. They are commonly used as materials of cutting tools, punch-and-die sets, and wear resistance components. Due to their extremely high hardness, tungsten carbides are generally machined by grinding with diamond wheels [1]. Both resin bond and single layer plated diamond wheels have been used for grinding tungsten carbides, with the resin bond wheels mainly for plain or simple workpiece profiles [2-7] and the plated wheels for grinding of both simple and complex profiles [7-9]. Despite of being very brittle, tungsten carbides were found to exhibit extensive plastic deformation during grinding as revealed by flow-typed chips, large compressive residual stresses in ground surfaces, and high energy consumed on the plowing of diamond grains against workpieces [1, 2]. Typical wheel speeds with resin bond wheels reported were in the range of 15 to 3 m/s [3-5]. Creep-feed grinding of tungsten carbides with resin bond wheels showed benefits long time ago [6, 7]. A cost saving of about 83% was achieved in creep-feed grinding of solid tungsten carbide gear cutting hobs as compared with conventional pendulum grinding [6]. It should be noted that profile grinding with wheels that need truing such as vitrified or resin bond wheels is usually limited by the complexity of the ground profiles. If the corresponding wheel axial profile is dependent on the wheel diameter, as in the case of helical flute grinding, the workpiece profile cannot be ground precisely to the geometry without truing the wheel profile based on the wheel diameter, which becomes smaller and smaller due to the removal of wheel material in truing operations. Electroplated superabrasive wheels, either with diamond or cubic boron nitride (CBN) abrasives, are manufactured with a single layer of abrasive grains held on the wheel hub by an electroplated nickel bond. Unlike other types of wheels, plated wheels are not periodically trued or dressed and are more suitable for profile grinding. Grinding of tungsten carbides with plated diamond wheels was introduced long before the invention of Cubic Boron Nitride (CBN) abrasives [8] and was successfully applied to the grinding of various tungsten carbide components [9]. However, there was very little work reported concerning the grinding performance of tungsten carbides with plated diamond wheels as compared with the grinding of metallic materials with plated CBN wheels. The limited amount of literature suggested that the wheel speeds are around 2 m/s, and workspeeds from.2 to.4 mm/s in creep-feed grinding of tungsten carbide with plated diamond wheels [7-9]. Performing grinding operations in multi-tasking machining centres is one of the recent increasing trends for producing components with complex geometric features [1]. Plated wheels are more attractive in this regard since they do not need truing/dressing and can be programmed as milling cutters. Simulated grinding of nickel alloy turbine blade slots using small profiled plated wheels on machining centres showed very promising prospects [11]. INDUSTRIAL DIAMOND REVIEW 4/8 65

The present investigation was undertaken to evaluate grinding power, forces, surface roughness, and wheel wear in grinding tungsten carbide using small diameter plated diamond wheels. The objective is to explore the feasibility of, and process parameters for, grinding deep slots or grooves. Wheel and holder Washing nozzle Experimental work Plunge surface grinding experiments were conducted in up-mode on a Makino A88E high-speed (18, min -1 ) and high-power (5 kw) five-axis machining centre that had unique grinding capabilities. A picture is presented in Fig 1. showing the experimental setup. The grinding wheel was an 8- grit (nominal grain dimension d g = 181 µm) electroplated diamond mandrel of diameter d s = 19 mm. It was clamped onto an HSK1A tool holder. The wheel and holder assembly was mounted onto the machine spindle after being dynamically balanced to G2.5 at 2, min -1. The workpieces were rectangular tungsten carbide bars with a length of 27 mm along the grinding direction and a width of 5 mm corresponding to the grinding width. To minimise the negative effect associated with changing power and surface roughness in grinding with plated wheels, tests were conducted only after initial run-in of the new wheel at a fixed wheel speed v s = 18 m/s, but different workspeeds (v w =.2,.4,.8 mm/s) and wheel depths of cut (a = 2, 4 mm) in full combinations corresponding to a specific removal rate Q w ranging from.4 to 3.2 mm 2 /s. A 7% solution of soluble oil was applied at a flow rate of 21.3 litres/min through a nozzle with a rectangular orifice of 2 mm 2 (1 mm x 2 mm) to match the fluid shooting velocity with the wheel speed. The nozzle was oriented tangent to the grinding surface at about 7 mm away from the grinding zone. A separate wheel-washing nozzle was used to clean the wheel surface during grinding from a pump with a nominal flow rate of 34 litres/min at 7. MPa. Measurements included grinding power, forces, surface roughness, and radial wheel wear. A Hall-effect power sensor on the spindle motor was used to measure the total power of the motor. Vertical and horizontal forces were measured using a Kistler dynamometer. An example is given in Fig 2 showing the total components of power and forces, the non-grinding components recorded in a complete spark-out pass, and the net components obtained by subtracting the non-grinding components from the corresponding total power [12]. The average net horizontal force F h in this example was about 67 N, vertical force F v about 88 N, and average net power P about.6 kw. It should be noted that about 14.4 kw of the total power of 15 kw was consumed on non-grinding components due to the very high spindle rotational speed of 18, min -1. Surface roughness of the ground surface was measured across the grinding direction using a portable profilometer at the start, middle, and end locations along the grinding pass and the results were averaged. Radial wheel wear was measured by taking a shallow grinding pass on a specimen somewhat wider than the grinding wheel and about 5 mm thick along the grinding pass. The section of the wheel width not used for grinding provided a reference surface to measure the wear [13]. The wear depth was then obtained by making a stylus trace across the grinding track of the specimen using a Taylor-Hobson Form Talysurf profilometer. In addition to these measurements, wheel topography and ground surface texture were also scanned using the same Form Talysurf profilometer and observed using an optical microscope. Power P (kw) Vertical forces F v (N) Horizontal forces F h (N) Profile height (µm) Fig 1 Experimental setup 6 4 2-2 -4-6 -8 12 1 8 6 4 2 15 14 2 1 Total Spark-out Net 1 2 3 Time t (S) Grinding nozzle v s = 18 m/s v w =.8 mm/s a = 2 mm Fig 2 An example of measured grinding power and forces 8 6 4 2-2 -4-6 Wear depth Reference sections 4 5 62 63 64 65 66 67 68 69 7 71 72 Profile width (mm) Fig 3 Radial wheel wear at the end of initial wheel run-in 66 INDUSTRIAL DIAMOND REVIEW 4/8

233 µm Forces F h and F v (N) Power P (kw) 2 mm Alpha = 6 Beta = 15 Grain pullouts Grain fracture 2.6 mm Fig 4 Scanned wheel topography at the end of tests 1..8.6.4.2 14 12 1 8 6 4 2..2.4.6 a = 2 mm a = 4 mm a = 2 mm, F h a = 2 mm, F v a = 4 mm, F h a = 4 mm, F v Fig 5 Net grinding power, horizontal and vertical forces.8 1. Results As mentioned above, grinding tests were conducted only after the initial wheel run-in. The measured radial wheel wear profile at the end of wheel run-in is presented in Fig 3, which shows an average wear depth of about 4 µm. In addition to the wear depth, this profile also reveals the progressive smoothing trend of the wheel surface with continued grinding as indicated by the larger peak-to-valley value of 4 µm produced by the reference sections of the wheel width as compared to the smaller value of about 3 µm produced by the section used for grinding. Wheel wear measurements following an accumulated volumetric material removal per unit width of 48 mm 2 at the end of tests indicated that there was almost no measurable change the in wear depth and peak-to-valley values. Therefore, any change in grinding power, forces, and surface roughness as presented below can be attributed to the effects of grinding parameters. Optical microscope observations during grinding intervals and at the end of tests revealed that diamond grains were worn down by grain fracture and grain pullouts. There was virtually no wear flat contrary to what was found in the grinding of ceramics with plated diamond wheel [13]. This wheel wear mechanism was further confirmed by the wheel topography scanned on the used wheel at the end of tests as shown in Fig 4. Net power, horizontal and vertical forces for all the grinding conditions are presented in Fig 5 as plots versus workspeeds for each depth of cut. In every case, the power and forces tended to increase with workspeeds and depths of cut. However, power increased much faster with the larger depth of cut and slower workspeeds than with the corresponding smaller depth of cut and faster workspeeds of equal material removal rates. It can also be seen in Fig 5 that the vertical force was bigger than the horizontal force with the smaller depth of cut of 2 mm. This relationship was, however, reversed with the larger depth of cut of 4 mm. Measured surface roughness are summarised in Fig 6 including the arithmetic average value R a measured after each grinding pass and after three spark-out passes. It can be seen from this figure that surface roughness fluctuates around 3.5 µm within a relatively small range. There was no apparent trend on the effects of depths of cut, workspeeds, and spark-out passes on the R a value. Surface roughness R a (µm) 4. 3.5 a = 2 mm, before spark-out 3. a = 2 mm, after spark-out a = 4 mm, before spark-out a = 4 mm, after spark-out 2.5 8 grit plated diamond wheel d s = 19 mm v s = 18 m/s 2...2.4.6.8 Fig 6 Measured surface roughness 1. Discussion The magnitude of energy consumption in grinding is usually appreciated from the specific energy, which is the energy consumption per unit volume of material removal. In addition to indicating the combined effect of depth of cut and workspeed, specific energy u also reveals the mechanism of workpiece abrasive grain interaction since the latter must be able to account for its magnitude and its dependence on the process parameters. Specific energy u is calculated as [14]: (1) Specific energy u (J/mm 3 ) 4. 3.5 3. 2.5 2...2.4.6.8 Specific removal rate Q w (mm 2 /s) a v w (mm) (mm/s) 2.2 2.4 2.8 4.2 4.4 4.8 Fig 7 Specific energy versus specific material removal rate 1. where P is the grinding power, Q w the specific material removal rate equal to v w a, a the depth of cut, and b the grinding width. The specific energy corresponding to the power in Fig 5 is presented in Fig 7 versus the specific removal rate. The specific energy dropped sharply from the maximum value of 2 J/mm 3 with the lowest specific removal rate of.4 mm 2 /s down to 56 J/mm 3 at the highest specific removal rate of 3.2 mm 2 /s with the depth of cut 4 mm. This indicates that some of the total energy was not consumed for removing materials. But for equal material removal rates, larger specific energy values were obtained with the larger depth of cut of 4 mm as compared to smaller depth of cut of 2 mm. This observation is consistent with what was previously found with resin bond wheels in shallow pendulum grinding [2]. INDUSTRIAL DIAMOND REVIEW 4/8 67

Horizontal and vertical force components were measured. However, it is the components tangent and normal to the wheel surface are of prime importance since they are directly related to grinding power and the contact pressure between the wheel and workpiece in the grinding zone, respectively. In addition, these two force components can be easily applied for predicting the force acting on individual grains, given the active grain density. The relationship between horizontal force F h, vertical force F v, tangential force F t, normal force F n, and resultant force is illustrated in Fig 8. In this figure, α is angle from the bottom of contact arc to the resultant force location with respect to the wheel centre, and ß is the included angle of contact arc. Detailed calculations for converting F h and F v into F t, and F n can be found in [15]. The measured horizontal and vertical forces in Fig 5 are converted into tangential and normal forces and presented in Fig 9. It can now be seen that that the normal force is about three times higher than the corresponding tangential force for both depths of cut. The ratio of α and ß, which represents the normalised location of resultant force measured from the button of the contact arc, is presented in Fig 1. It can be seen that the resultant force is near the middle of contact arc with the depth of cut of 2 mm. A higher ratio of about.6 was obtained with the larger depth of cut of 4 mm, which implies that the resultant force location moves away from the bottom of the contact arc. Measurements presented in Fig 6 showed no apparent effect of grinding parameters and spark-out passes on the ground surface roughness within the tested range. To understand this fact, it is helpful to consider the characteristics of ground surface texture. With an ideal wheel having cutting points equally spaced and protruding the same height, the ideal peak-to-valley roughness along the grinding pass can be expressed as [14]: (2) Forces F t and F n (N) F v Ft F h F Fig 8 Relationship between F h, F v, F t, and F n 2 15 1 5 a = 2 mm, F t a = 2 mm, F n a = 4 mm, F t a = 4 mm, F n F n d s ß α v w.2.4.6 Fig 9 Normal and tangential forces versus workspeed v s.8 a 1. where L is the distance between cutting points around the wheel periphery. For wheels with grains protruding at different heights, the R t value is much bigger than expressed in Eq (2). In the extreme situation, if the highest point at each section of the wheel width is the only active grain in removing the material, the corresponding R t will reach its maximum value or upper bound of that section which can still be calculated using Eq (2) by considering the wheel is ideal with a grain distance L = πd s in that section. Combining L = πd s and Eq (2) leads to the upper bound value of R t : (3) α/ß.8.6.4.2 v w =.2 mm/s v w =.4 mm/s v w =.8 mm/s v w =.2 mm/s v w =.4 mm/s 1 2 3 4 Depth of cut a (mm) Fig 1 Normalised resultant force location v w =.8 mm/s 5 where n s is the wheel rotational speed. Substituting the grinding parameters into Eq (3) leads to the R t value ranging from 6x1-6 µm for the lowest workspeed of.2 mm/s to 9x1-5 µm for the highest workspeed of.8 mm/s. It should be noted that Eq (3) is for the upper bound value of R t developed from the kinematics of grinding. Actual values are much higher than predicted due to vibrations and other factors [14]. The extremely small R t value, however, does indicate that the workpiece surface corresponding to each section of the wheel width was generated only by the highest cutting point or points of equal protruding heights on that section. The effect of the very high wheel rotational speed n s and very slow workspeed v w can be appreciated from the ratio of v w /n s, which corresponds to the infeed distance per wheel revolution of the workpiece along the grinding pass. In the present work, the infeed per wheel revolution ranged from.7 to 2.7 µm, which was about 1 to 2 times smaller than obtained in conventional creep-feed grinding. These very small values imply that clean cutting dominated, whereby the cutting edges removed all materials that they encounter in their paths. Any material plowed aside by a cutting edge was removed immediately by the cutting edges on its adjacent section of wheel width within one wheel revolution. 68 INDUSTRIAL DIAMOND REVIEW 4/8

Clean cutting by only the highest cutting point on each wheel section with highly wear resistant diamond grains is expected to produce ground surface with parallel grooves and ridges along the grinding pass and extending to the entire pass length, instead of cross overlapping each other as reported under shallow pendulum grinding conditions [16]. This was actually found in optical microscope observation of the ground surface. From the above discussion, it can be concluded that the distribution of grooves and ridges across grinding direction is dependent on the radial distribution of the highest cutting point at each section of wheel width and not on the grinding parameters within the present tested range. Spark-out passes should remove some material left from the previous grinding pass due to system deflection. But they should have no effect on the ground surface texture. Therefore, it is apparent that the surface roughness measured across the grinding direction or groves and ridges is independent of grinding parameters and spark-out passes. The grinding wheel topography is the controlling factor on surface roughness. Conclusions Diamond grains in the plated wheel used for this investigation were worn down by grain fracture and grain pullouts. There was no measurable wheel wear for an accumulated material removal per unit width of 48 mm 2 conducted after the initial wheel run-in, which revealed that smaller plated diamond wheels can last long enough for tungsten carbide grinding applications. Depths of cut as high as 4 mm corresponding to a specific material removal rate of 3.2 mm 2 /s were achieved. Surface roughness is independent of grinding parameters in creep-feed grinding with smaller wheels running at high wheel rotational speeds, but is nearly solely controlled by wheel topography. To evaluate the wheel life, additional work is needed to carry out tests all the way down to the end of wheel life. References [1] O. Zelwer, and S. Malkin. Grinding of WC-Co Cemented Carbides Part I. Industrial Diamond Review 198 April, pp 133-139. [2] O. Zelwer, and S. Malkin. Grinding of WC-Co Cemented Carbides Part II. Industrial Diamond Review 198 May, pp 173-176. [3] E. Ratterman. An Analysis of the Low Speed Grinding of Tungsten Carbide. The 3rd DWMI Technical Symposium, Cleveland, Ohio 1975, pp 1-8. [4] J. L. Metzger. The Effect of Grit Size and Carbide Grade in Plunge Grinding. Industrial Diamond Review 1982, Vol. 3, pp 145-149. [5] J. L. Metzger. Wheel Wear versus Workpiece/Wheel Rim Impact in Carbide Grinding. Industrial Diamond Review 1981, Vol. 4, pp 192-195. [6] H. O. Juchem. Creep Feed Grinding A Review. Industrial Diamond Review 1984, Vol. 3, pp 17-114. [7] C. Andrew, and T. Howes. Creep Feed Grinding. Industrial Press Inc., New York 1985. [8] B. W. Bruce. Reduced Grinding Costs with Plated Diamond Wheels. Cutting Tool Engineering 1979, Vol. 31, pp 71-74. [9] F. Curn. Diamond Profiling Wheels for Grinding of Sintered Carbide. Industrial Diamond Information Bureau. London 1967, pp 291-37. [1] S. Salmon. Creep-Feed Grinding is Surprisingly Versatile. Manufacturing Engineering 24 November, pp 59-64. [11] D. K. Aspinwall. The Use of Diamond and Cubic Boron Nitride Grinding Points for the Machining of Nickel-based Superalloys. InterTech, an International Technical Conference on Diamond, Cubic Boron Nitride and their Applications 2 July, pp 1-13. [12] M. Ganesan, C. Guo, and S. Malkin. Measurements of Hydrodynamic Force in Grinding. Transactions of NAMRI/SME 1995, Vol. 23, pp 13-17. [13] T. W. Hwang, and C. J. Evans. High Speed Grinding of Silicon Nitride with Electroplated Diamond Wheels, Parts I & II. ASME Journal of Manufacturing Science and Engineering. 2, Vol. 122, pp 32-5. [14] S. Malkin. Grinding Technology: Theory and Applications of Machining with Abrasives, Ellis Horwood Ltd, Chichester, and John Wiley & Sons, New York 1989. [15] K. Li, T. W. Liao. Wear of Diamond Wheels in Creep-feed Grinding of Ceramic Materials Effects on Process Responses and Strength. Wear 1997, Vol. 211, pp 14-112. [16] X. Zhou, F. Xi. Modeling and Predicating Surface Roughness of the Grinding Process. International Journal of Machine Tools & Manufacture 22, Vol. 42, pp 969-977. Authors All authors work for the National Research Council of Canada. Z. Shi is a Research Officer, H. Attia is a Principal Research Officer and Program Manager, D. Chellan is a Senior Technical Officer and T. Wang is Technical Officer. Contact address is the Aerospace Manufacturing Technology centre, Institute for Aerospace Research, National Research Council Canada, 5145 Avenue Decelles, Montreal, Quebec, H3T 2B2, Canada. Acknowledgment This article is based on a paper presented at the 2nd International Industrial Diamond Conference held in Rome, Italy on April 19-2 27 and is printed with kind permission of Diamond At Work Ltd. INDUSTRIAL DIAMOND REVIEW 4/8 69