肥料顆粒機(jī)設(shè)計(jì)【非擠壓式圓盤造粒機(jī)的總體設(shè)計(jì)】【圓盤式肥料制粒機(jī)】【說(shuō)明書(shū)+CAD+SOLIDWORKS】

肥料顆粒機(jī)設(shè)計(jì)【非擠壓式圓盤造粒機(jī)的總體設(shè)計(jì)】【圓盤式肥料制粒機(jī)】【說(shuō)明書(shū)+CAD+SOLIDWORKS】,非擠壓式圓盤造粒機(jī)的總體設(shè)計(jì),圓盤式肥料制粒機(jī),說(shuō)明書(shū)+CAD+SOLIDWORKS,肥料顆粒機(jī)設(shè)計(jì)【非擠壓式圓盤造粒機(jī)的總體設(shè)計(jì)】【圓盤式肥料制粒機(jī)】【說(shuō)明書(shū)+CAD+SOLIDWORKS】,肥料,顆粒機(jī) Journal of Materials Processing Technology 83 (1998) 151158 Machining of Al:SiC particulate metal-matrix composites Part I: Tool performance M. El-Gallab a , M. Sklad b, * a Pratt and Whitney Canada, Mississauga, Ontario, L5T 1J3, Canada b Department of Mechanical Engineering, McMaster Uni6ersity, Hamilton, Ontario, L8S 4L7, Canada Received 20 March 1997 Abstract Despite the superior mechanical and thermal properties of particulate metal-matrix composites, their poor machinability has been the main deterrent to their substitution for metal parts. The hard abrasive reinforcement phase causes rapid tool wear during machining and, consequently, high machining costs. A series of dry high-speed turning tests were performed to select the optimum tool material, tool geometry and cutting parameters for the turning of 20%SiC:Al metal-matrix composites. The results indicate that polycrystalline diamond tools (PCD) provide satisfactory tool life compared to alumina and coated-carbide tools, where the latter tools suffered from excessive edge chipping and crater wear during the machining of the metal-matrix composite under study. Furthermore, the cost of PCD tools could be justified by using dry cutting at feed rates as high as 0.45 mm rev 1 , cutting speeds of 894 m min 1 and a depth of cut of 1.5 mm. With these cutting parameters, the relatively small built-up edge formed on the tool protects it from further wear by abrasion and micro-cutting. Polycrystalline tools with zero rake angle and large tool nose radii are recommended for the roughing operations. 1998 Elsevier Science S.A. All rights reserved. Keywords: Metal-matrix composites; Tool wear 1. Introduction Metal-matrix composites (MMCs) form one group of the new engineered materials that have received consid- erable research since the trials by Toyota in the early 1980s 1. The most popular reinforcements are silicon carbide and alumina. Aluminium, titanium and magne- sium alloys are commonly used as the matrix phase. The density of most MMCs is approximately one third that of steel, resulting in high specific strength and stiffness 2. Due to these potentially attractive proper- ties coupled with the inability to operate at high tem- peratures, MMCs compete with super-alloys, ceramics, plastics and re-designed steel parts in several aerospace and automotive applications. The latter materials, how- ever, may not have much further capacity for the inevitable future increases in service loads 3. Particulate metal-matrix composites (PMMCs) are of particular interest, since they exhibit higher ductility and lower anisotropy than fiber reinforced MMCs 2. Moreover, PMMCs offer superior wear resistance 3. While many engineering components made from PMMCs are produced by the near net shape forming and casting processes, they frequently require machin- ing to achieve the desired dimensions and surface finish. The machining of PMMCs presents a significant chal- lenge, since a number of reinforcement materials are significantly harder than the commonly used high-speed steel (HSS) and carbide tools 4. The reinforcement phase causes rapid abrasive tool wear and therefore the widespread usage of PMMCs is significantly impeded by their poor machinability and high machining costs. 2. Literature review From the available literature on PMMCs, it is clear * Corresponding author. Fax: 1 905 5727944; e-mail: mech- macmcmaster.ca 0924-0136:98:$19.00 1998 Elsevier Science S.A. All rights reserved. PII S0924-0136(98)00054-5 M. El-Gallab, M. Sklad : Journal of Materials Processing Technology 83 (1998) 151158152 Table 1 Typical physical properties of Duralcan F3S.20S 20 Property Density (g cm 3 ) 2.77 Thermal conductivity (cal cm 1 s 1 K 1 ) at 22C 0.47 Specific heat (cal g 1 K) 0.218100C 0.239200C 0.259300C Average coefficient of thermal expansion 17.550100C 50300C 21.1 21.450500C Ultimate strength (MPa) 262 Yield strength (MPa) 21.4 1.9Elongation (%) 98.6Elastic modulus (GPa) 6791.5Rockwell hardness (B) Table 2 Tool geometry Tool Geometry PCD (tipped-80rhomboid) Rake angle, a: 5, 0, 5 Clearance angle: 7 Approach angle: 5 Tool nose radius (mm), r: 0.8, 1.6 TiN coated carbide Rake angle, a:0 (triangular) Clearance angle: 7 Approach angle: 5 Tool nose radius (mm), r: 1.6 Rake angle, a:0Al 2 O 3 :TiC ceramic (triangular) Clearance angle: 7 Approach angle: 5 Tool nose radius (mm), r: 1.6 crystalline diamond (PCD) tools are the only tool mate- rial that is capable of providing a useful tool life during the machining of SiC:Al PMMCs. PCD is harder than Al 2 O 3 and SiC and does not have a chemical tendency to react with the workpiece material. Tomac et al. 5 compared the performance of chemical vapor deposi- tion (CVD) inserts to that of TiN, Ti(CN) and Al 2 O 3 coated tools. CVD tools offered better overall perfor- mance than that of the other tools. Lane et al. 18 studied the performance of different CVD tools with thin and thick films. According to their observations, CVD tools with thin films failed catastrophically during the end milling of 20%SiC:Al PMMC. This tool failure was attributed to coating spalling and consequent dam- age to the relatively soft carbide substrate. Further- more, PCD tools with a grain size of 25 mm better withstand abrasion wear by micro-cutting than tools with a grain size of 10 mm 8,14. Further increases in PCD grain size do not benefit the tool life, but rather cause significant deterioration in the surface finish. This that the morphology, distribution and volume fraction of the reinforcement phase, as well as the matrix prop- erties, are all factors that affect the overall cutting process 2,4, but as yet relatively few works related to the optimization of the productivity process have been pulished. For example, Monaghan 2 studied the wear mechanism of carbide tools during the machining of 25% SiC:Al PMMC at speeds below 20 m min 1 . Tomac et al. 5 developed a tool life relationship for carbide tools during the machining of SiC:Al PMMCs at speeds lower than 100 m min 1 . However, the authors of Ref. 5 recommended further research on the built-up edge phenomenon that is observed in all tools during the machining of SiC:Al PMMCs. OReilly et al. 6 ranked various tool materials with respect to tool wear, however, their cutting parameters did not exceed 125 m min 1 and 1.0 mm depth of cut, which was achieved using cubic boron nitride tools. Similar test results were reported by Brun et al. 7 who related the tool wear rate, mainly due to abrasion, to the tool hardness. Winert 8 attributed the wear of the carbide tools to abrading Al 2 O 3 particles that form on the surface and rub the tool in the direction of the chip flow. However, pulled SiC particles could also lead to the same effect, SiC particles also being harder than the WC. Tomac et al. 5 suggested that coatings with less hardness than that of Al 2 O 3 and SiC offer little to no advantage during the machining of SiC:Al PMMCs, Brun et al. 7 suggested using lower cutting speeds to reduce the cutting temperature, which accelerate diffu- sion and adhesion wear and thermally weaken the tool. Since aluminium tends to seize on the tool face and since grain boundaries are the sites of seizure, the authors recommended using cemented carbide tools with a large grain size. Several researchers 725 have indicated that poly- Fig. 1. SEM figure showing wear on Al 2 O 3 tool (6 488 m min 1 , f 0.2 mm rev 1 , doc 0.5 mm, r 1.6 mm, a 0). M. El-Gallab, M. Sklad : Journal of Materials Processing Technology 83 (1998) 151158 153 Table 3 Effect of tool material on cutting forces and temperatures (r 1.6 mm; a 0) Measured cut-Tool material Measured cutting temperature (C)ting force (N) PCD (6 894 m min 1 ; 97.00 440 f 0.45 mm rev 1 ; doc 2.5 mm) 98.10PCD (6 670 m min 1 ; 410 f 0.25 mm rev 1 ; doc 1.5 mm) 520Al 2 O 3 (6 248 m min 1 ; 183.85 f 0.2 mm rev 1 ; doc 0.5 mm) 143.52 500TiN (6 248 m min 1 ; f 0.2 mm rev 1 ; doc 0.5 mm) Fig. 3. Effect of depth of cut on the cutting forces (PCD tool: r 1.6 mm, a 0; square points: 6 894 m min 1 , doc 1.5 mm; round points: 6 894 m min 1 , doc 2.5 mm). of cut that are as aggressive as possible during the roughing operations. Finally, with regard to the coolant application, researchers at Duralcan USA 1123, rec- ommend investigating the possibility of dry-rough ma- chining in order to take advantage of the protective built-up edge phenomenon. In summary, the literature review carried out showed that the effect of more aggressive cutting parameters (speed, feed and depth of cut) still needs further re- search in order to improve the economics of the cutting process. Also, several important parameters have been overlooked by previous researchers, among which are the tool geometry and coolant application. 3. Test material and cutting tools 3.1. Workpiece material The machining investigations were carried out using Duralcan F3S.20S Al:SiC metal-matrix composite. The SiC particles had an average diameter of 12 mm. Table 1 shows some of the physical and mechanical properties of A356-20%SiC PMMC. Prior to carrying out the cutting experiments, the test material was fully heat- treated to the T71 condition. The test material was in the form of bars of 177.8 mm diameter and 305 mm length. 3.2. Cutting tools Various tool materials (coated carbide, Al 2 O 3 :TiC and PCD) and geometries were employed in the study. Oblique turning tests were carried out and different cutting parameters were employed for each tool mate- rial. However, for the purpose of comparing tool wear, all cutting tests were carried out at a fixed volume of metal removed (300 mm 3 ). Table 2 summarizes the tool data. is because PCD grains with size 25 mm are easily pulled out of the cutting edge. Regarding the effect of the cutting parameters on the tool life, Lane et al. 1115,1719 attributed the in- crease in the wear of PCD tools (by abrasion) to increase in kinetic energy gained by abrading SiC parti- cles. On the other hand, Brun et al. 7 attributed the increase in tool wear in the thermal degradation of the tool material. Tool wear was found to be inversely proportional to the feed rate 9. Tomac et al. 5 attributed the increase in tool life at higher feed rates to the thermal softening of the composite. The authors suggest that the workpiece material becomes softer and the SiC particles become pressed into the workpiece, causing less abrasion on the tool itself. However, Finn et al. 24 and Morin et al. 25 attributed the reduction in tool wear with greater feed rates to the reduced contact between the cutting edge and the abrasive SiC particles. Despite the controversy in explaining the mechanism behind the tool wear at different feed rates, all researchers recommend using feed rates and depths Fig. 2. Effect of cutting speed on the cutting forces (PCD tool: r 1.6 mm, a 0; square points: 6 670 m min 1 , doc 1.5 mm; round points: 6 894 m min 1 , doc 1.5 mm). M. El-Gallab, M. Sklad : Journal of Materials Processing Technology 83 (1998) 151158154 Dry turning tests were carried out on a 10 HP Standard Modern NC lathe. The cutting force compo- nents (Fx, Fy, Fz) and tool wear were measured contin- uously for each combination of cutting parameters. The tool forces were measured using a Kistler 3-component dynamometer and the cutting conditions were selected carefully for each tool material. In some of the cutting tests, the tool temperature was measured using K-type thermocouples that were glued onto the tool rake face, 1 mm away from the cutting edge. The reliability of the measurement techniques was checked constantly by repeating the experiments and the results of each set of experiments were accepted if they exhibited a variance of less than 5%. At the end of each cutting test, the tool wear was examined using a scanning electron micro- scope and the X-ray dispersion technique. The tool Fig. 5. (a) Built-up edge on PCD tools (6 670 m min 1 , f 0.45 mm rev 1 ; doc 2.5 mm, r 1.6 mm); (b) as for Fig. 5(a), but for 6 894 m min 1 . Fig. 4. (a) Built-up edge on PCD tools (6 670 m min 1 , f 0.25 mm rev 1 , doc 2.5 mm, r 1.6 mm, a 0); (b) X-ray dispersion of built-up edge shown in Fig. 4(a). flank wear (VB) was measured using a toolmakers microscope. 4. Results and discussion 4.1. Effect of tool material A series of preliminary tests was conducted to asses the effect of tool material on the tool wear, cutting forces and cutting temperature during the rough turn- ing of 20%SiC:Al PMMC. Fig. 1 shows that Al 2 O 3 :TiC tools suffered excessive wear in the form of edge chip- ping. Al 2 O 3 particles are pulled out by the abrading workpiece particles, which have a greater Vickers hard- ness number (VHN) than the Al 2 O 3 particles (VHN for Al 2 O 3 TiC 2500 kg f mm 2 ; VHN for SiC 3000 kg f mm 2 ). Crater wear was also observed, which is due to the widening of grooves that were caused by abrasion. Due to severe edge chipping, the cutting forces for the M. El-Gallab, M. Sklad : Journal of Materials Processing Technology 83 (1998) 151158 155 Al 2 O 3 :TiC tool were much higher than those experi- enced by TiN coated tools (Table 3). The TiN coating provided some protection against the abrasive effects of the SiC particles. The superior performance of poly- crystalline diamond tools, compared to both Al 2 O 3 :TiC and TiN coated carbide tools, is attributed to their high abrasion resistance and high thermal conductivity, which led to lower cutting temperatures, as shown in Table 3. Therefore, all machinability studies carried out thereafter were concerned with the optimization of the cutting process using PCD tools. 4.2. Effect of cutting parameters Figs. 2 and 3 show that as the cutting speed and:or the depth of cut increase, the cutting forces decrease. This could be attributed to thermal softening of the workpiece material. Another possible reason is due to the changes introduced into the tool geometry upon the formation of built-up edge. Fig. 4(b) shows the X-ray dispersion of the built-up material shown in Fig. 4(a). Fig. 7. (a) SEM image illustrating the wear on the PCD tool rake face after dissolving the BUE with NaOH (6 670 m min 1 , f 0.15 mm rev 1 , doc 1.5 mm, r 1.6 mm, a 0); (b) higher magnification of rake face of the tool shown in Fig. 7(a). Fig. 6. (a) Built-up edge on PCD tools (6 670 m min 1 , f 0.35 mm rev 1 , doc 1.5 mm, r 1.6 mm, a 0); (b) as for Fig. 6(a), but for doc 2.5 mm. Built-up edge was observed in all tools under all cutting conditions. This is because particulate SiC:Al MMCs have all of the characteristics of materials that form Fig. 8. Effect of cutting speed on the tool flank wear (PCD tool: r 1.6 mm, a 0; square points: 6 670 m min 1 , doc 1.5 mm; round points: 6 894 m min 1 , doc 1.5 mm). M. El-Gallab, M. Sklad : Journal of Materials Processing Technology 83 (1998) 151158156 Fig. 9. Effect of depth of cut on the tool flank wear (PCD tool: r 1.6 mm, a 0, 6 894 m min 1 ; square points: doc 1.5 mm; round points: doc 2.5 mm). illustrate this point, in the case of a greater depth of cut a larger surface area of the tool flank face is exposed to abrasion. Increasing the feed rate had a beneficial effect. As shown in Figs. 8 and 9, as the feed rate increases, the tool wear decreases. In the case of higher feed rates, for a fixed volume of metal removal, the tool surfaces will have less contact with the abrasive PMMC. Another Fig. 10. (a) Effect of PCD tool rake angle on tool flank wear (6 894 m min 1 , doc 2.5 mm, r 1.6 mm; square points: 0; round points: 5; star points, 5); (b) effect of PCD tool rake angle on the cutting forces (6 894 m min 1 , doc 2.5 mm, r 1.6 mm; square points: 0; round points: 5; star points, 5); (c) SEM image illustrating the PCD tool wear by pitting (6 670 m min 1 , doc 1.5 mm, f 0.25 mm rev 1 , r 1.6 mm, a 5). BUE (i.e. strain-hardened two-phase material under high temperature and pressure). At high cutting speeds (6 894 m min 1 (Fig. 5(b), a smaller BUE is formed, compared to the BUE formed at 6 670 m min 1 (Fig. 5(a); the height of the BUE was measured perpendicular to the rake face) In con- trast, by increasing the depth of cut from 1.5 to 2.4 mm, a large BUE is formed (Fig. 6(a,b), which could break off the tool causing tool chipping and consequent adverse effects on the workpiece surface roughness and dimensional accuracy. Topographies of the tool indicate that the main wear mechanism of PCD is abrasion (manifested as grooves parallel to the chip flow direction). These grooves could be attributed to three factors. The first is that Al 2 O 3 is formed at the tool edge, which is hard enough to produce grooving wear in the PCD. The second expla- nation for the PCD grooving is aluminium seizure and the pull-out process of the PCD grain, as shown in Fig. 7(a,b). The third possible reason behind PCD grooving is that SiC particles abrade the tools. Thus, PCD tools with PCD grains larger than the grain size of the SiC particles could better withstand the abrasion and mi- cro-cutting by the SiC particles. However, one should note that as the size of the PCD grains increases, the fracture properties of the PCD tool deteriorates, due to an increased number of flaws in the material. The grooves that were formed on the tool face were filled with the workpiece material. This adhering layer somewhat protected the tools rake face against further abrasion. Nonetheless, the tool flank face continued to be subjected to abrasion. Hence, flank wear (Vb) was taken as the tool life criterion with Vb 1im 0.18 mm. Fig. 8 shows that as the cutting speed increases, the flank wear increases. This could be attributed to the increase in the kinetic energy of the abrading particles, as previously hypothesized by Lane 17. Increasing the depth of cut leads to an increase in the flank wear (Fig. 9). This is attributed to enhanced abrasion by micro-cutting at the tool flank face. To M. El-Gallab, M. Sklad : Journal of Materials Processing Technology 83 (1998) 151158 157 Fig. 11. (a) SEM image illustrating the PCD tool wear by chipping (6 894 m min 1 , doc 1.5 mm, f 0.35 mm rev 1 , r 0.8 mm, a 0); (b) effect of tool nose radius on the tool flank wear (6 894 min, doc 2.5 mm, a 0; tool FW: square points, r 1.6 mm; round pitch points, r 0.8 mm). wear in the case of negative rake angle, is the greater cutting forces encountered with such a rake angle (Fig. 10(b). Moreover, the chips produced became caught between the tool and the workpiece, causing damage to the tool surface. Tools with positive rake angle showed irregular flank wear and excessive pitting in the cutting edge zone, as shown in Fig. 10(c). The tool nose radius plays a key role in determining the wear mode of the tool. As the tool nose radius was decreased from 1.6 to 0.8 mm, the tool was found to suffer from excessive chipping and crater wear, as shown in Fig. 11(a). This tool chipping leads to an increase in cutting forces and flank wear, as shown in Fig. 11(b). Tools with small nose radii are thus recom- mended for finishing operations where light cutting parameters are used. Small nose radii are also expected to yield better geometrical accuracy. 5. Conclusion The results of the machinability studies carried out on 20%SiC:Al particulate metal-matrix composites in- dicated the following. (1) The main tool wear mechanism is abrasion and micro-cutting of tool material grains, manifested as grooves on the tool face parallel to the chip flow direction. All of the tools tested also suffered from flank wear due to abrasion. There was no evidence of chemical wear (e.g. by diffusion). (2) PCD tools sustained the least tool wear com- pared to TiN coated carbide tools and Al 2 O 3 :TiC tools. This is undoubtedly due to PCDs superior hardness and wear resistance, as well as low coefficient of fric- tion, together with high thermal conductivity. This led to lower cutting temperatures when PCD tools were employed. On the other hand, the TiN coated carbide tools and Al 2 O 3 :TiC tools s