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Figure 242. HSM (milling) of three machinable disks in-situ on a stereometric artefact, on a vertical machining centre. [Source: Smith, Sims, Hope & Gull, 2001] . Machining and Monitoring Strategies cranked-probe – with its calibration obtained from the ‘reference measurement sphere’ being located on the CMM’s table, utilised to inspect the φ10 mm hole geometry and there respective co-ordinate positions. is probe arrangement was swapped for a conven- tional ‘touch-trigger probe assembly’ to measure the machinable disk diameters – while holding the same cartesian co-ordinate relationships as when it was originally UHM. Later – without ‘breaking-down’ , while still maintaining the same angular orientation, this stereometric artefact assembly was inspected on a roundness testing machine (Taylor Hobson: ‘Talyrond 265’) for individual disk parameters of roundness and cylindricity 37 assessment – for the ‘three-disk relation- ship’. e results of all of these ‘averaged’ roundness measurements and Ballbar polar plots are graphically depicted as histograms in Fig. 243c. When a comparison is made of these results from three individual and completely diering inspec- tion procedures, namely: Ballbar; CMM; and Taly- rond, they show some degree of measurement con- sistency individually, but less so when each disk data is grouped. For example, in the case of the Ballbar, it indicated a 1 µm variation (i.e. range) from the top-to- bottom disks, while having a mean value of 17.5 µm. e Talyrond polar plots (i.e. ‘Least Squares Reference Circle’ 38 : departures from roundness) also produced consistent roundness results, ranging from <4 µm, having a mean value of ≈14 µm. Conversely, the larg - est variability occurred with the CMM, producing a 37 ‘Cylindricity’ , can be dened as: ‘e minimum radial separation of 2 cylinders, coaxial with the tted reference axis, which totally enclose the measured data’. (Source: Taylor Hobson, 2003) NB A ‘working denition for cylindricity’ , might be: ‘If a perfectly at plate is inclined at a shallow angle and a parallel cylindrical component is rolled down this plate. If it is a truly round cylinder then as the component rolls, there should be no discernible radial/longitudinal motion apparent’. (Sources: Dagnall, 1996; Smith, 2002) 38 ‘Least Squares Reference Circle’ (LSC1), can be dened as: ‘A line, or gure tted to any data such that the sum of the squares of the departure of the data from that line, or gure is a mini- mum’. is is also the line that divides the prole into equal minimum areas NB is LSC1 is the most commonly used ‘Reference Circle’. e ‘out-of-roundness’ , or ‘departures-from-roundness’ as it is now known, is then expressed in terms of the maximum departure of the prole from the LSC1 (i.e. the highest peak to lowest valley – on the ‘polar plot’). (Source: Taylor Hobson, 2003) range of 15 µm, with a mean value of ≈23 µm. Prior to discussing why the CMM results signicantly varied from those obtained by both the Ballbar and Talyrond, it is worth visually looking at a comparison between the general proled shapes of typical ‘polar plots’ pro- duced by both these techniques. In Fig. 242a, a rep- resentative ‘polar plot’ from a Ballbar is shown and, likewise in Fig. 242c, one from a Talyrond is depicted. eir respective proled shape geometry in terms of harmonics, is remarkably alike, illustrating the same generally similar lobed-shape combined with its iden- tical angular orientation. Returning to the CMM results, only a few data points are utilised to obtain a measured diameter, while with the Ballbar and Talyrond alike, they liter- ally take thousands of data points to obtain the polar plotted proles. If, when the CMM touches each of the machinable disk’s prole with the ‘touch-trigger probe’ , this co-ordinate’s data could have been ob - tained at the extremes of the elliptical shape, namely, at its major and minor diameters, this may account for such a variation in both the range and discrepancies, when compared to the data obtained by the Ballbar and Talyrond. e four φ10 mm holes in each disk that were pro- duced by HSM utilising circular interpolation at their respective quadrant positions (Figs. 241 and 242b), are given in the form of tabulated data in Table 16 – in terms of their positional accuracy and radial change, from their theoretical centres. From this φ10 mm hole data, then the radial change for each disk, from the top, middle and bottom disks, was: 46 µm: 45 µm: and 42 µm: respectively, giving a positional uncertainty across these disks of 4 µm. Conversely, if the dierence is considered for the three stacked disks with respect to their angular rela- tionships to each other, at: 0°; 90°; 180°; 270°; then their angular positional changes are: 46 µm; 26 µm; 32 µm; and 49 µm; giving a positional uncertainty of 23 µm. is positional uncertainty is still relatively small con- sidering that in this case, each hole’s position is on a dierent Z-axis plane – spanning 200mm in height. Although if one considers the ‘Grand mean’ for both cases then they have a positional uncertainty of just 1 µm, which for a machine tool that at this time was around three years old is quite exceptional – having by now, undertaken considerable industrial machinabil- ity trials for the automotive and aerospace industries, but admittedly, this vertical machining centre had pre- viously been both Laser- and Ballbar-diagnostically corrected – showing the ‘true’ relevance of calibration to resolve and reduce any ‘errors’! Chapter is initial calibration work using the stereomet- ric artefact, has shown that its overall positional un- certainty when utilised for HSM accuracy and preci- sion assessments in combination with machine tool diagnostics, then compared to other more-exacting inspection techniques and measurements, has been successful. Hence, a large artefact having replaceable machining components (disks), has proved to be a valid means to verify the actual ‘machine tool’s health’ – in a real sense, as it is undertaken in a loaded state, while partaking in high-speed cutting trials. Once, the part- program has been written, then the whole machining activity can be completed in a relatively short time- frame. Although admittedly, the additional metro- logical inspection and data analyses takes sometime to complete, but these inspection functions can be un- dertaken while the machine tool is still in operation – producing quality machined components. 9.7 HSM: Rotating Dynamometry High-speed rotating dynamometers are being utilised across diverse elds of industry, where: hard-part ma- chining such as die-sinking are necessary; free-form sculptured parts are required; expensive and delicate workpieces that need closely-monitored cutting con- ditions; together with applied and fundamental ma- chinability research programmes. HSM rotating cut- ting force dynamometry, allows one to insert the unit into the machine tool’s spindle, where it measures the forces acting on the tool. is measurement data are directly amplied within the dynamometer, then radio transmitted to a specially-congured receiver (Fig. 244). ere are many technical advantages why a rotating dynamometer is preferred to that of its spa- tially-stationary counterpart (i.e. platform-type: see Fig. 237), such as: • Measurement occurs at the rotating tool’s cutting edge – meaning that the cutting force near to the cutting process is measured during engagement with the workpiece, not a reaction force, • Dynamometer’s mass remains constant, maintain- ing a uniform level of natural frequency – while in contrast, a stationary version (e.g. a platfrom-type), has the workpiece’s mass (i.e normally placed onto the dynamometer) which continuously diminishes as the machining operation takes place, • Its small mass provides little inertial eects, re- maining constant throughout the machining op- eration – unlike that of a stationary platform dyna- mometer, • Spindle-mounted dynamometers can occupy/ori- entate themselves into any position, during ma- chining – whereas the stationary platforms have certain volumetric space-requirements, having just one orientation to the applied cutting forces. Table 16. The φ10 mm hole positional deviations for the truncated frustum – based upon the three-dimensional Isosceles tri- angle in the disks at four quadrants (from the theoretical), in terms of their radial change Position of holes: Top disk Middle disk Bottom disk Range Mean Grand Mean 0 degree 987 941 969 46 966 ↓ 90 degree 978 952 953 26 961 972 180 degree 1014 986 982 32 994 270 degree 968 946 995 49 965 Range 46 45 42 Mean 987 956 975 Grand mean → 973 NB Values in: µm [Source: Smith, Sims, Hope and Gull, 2001] . Machining and Monitoring Strategies Figure 243. Technique of manufacture of the aerospace aluminium disks and the ‘averaged’ tabulated inspection data. [Source: Smith, Sims, Hope & Gull, 2001] . Chapter e HSM rotating cutting force dynamometer depicted in Fig. 244, is of compact dimensions having an overall protrusion length of 106 mm, having a standard collet chuck adaptor. With its short overall length, this HSM dynamometer has the sensor screwed onto the sensor bearer, giving a high moment of bending resistance, which is reected in the small amount of ‘cross-talk’ values of radial forces on the F Z -force component (i.e. F Radial → F Z ). e rotating HSM dynamometer was designed to have a ‘ceiling-speed’ in accordance with ISO Stan- dard: 15641: 1998, which species a test speed of 40,000 rev min –1 . Although the manufacturer originally tested the antenna casing up to speeds of 56,000 rev min –1 , corresponding to a centrifugal acceleration of: a Zcentrifugal = 130,000 g. Any HSM tooling assembly, such as a rotating dy- namometer must be balanced with particular care, in this case, there is a pre-balance quality of 6.3 for the antenna casing and sensor, with the nal balance be- ing 2.5 g-mm, which ensures that the cutting force measurements are not adversely aected. e overall compact construction minimising ac- celeration masses (i.e. weighing just 5.3 kg) for this HSM rotating dynamometer, with its low-level of natural frequency, having both a very low response threshold coupled to high resolution precision, pro- vides ‘sound’ cutting force and torque information making it an ideal ‘tool’ for either qualitative, or quan- titative machinability research work. Hence, the signal quality achieved by this instrument, makes it possible to utilise this HSM rotating cutting force dynamom- Figure 244. A high-speed rotating cutting force dynamometer, employed during drilling and multi-axis milling operations. [Courtesy of Kistler Instru- mente AG] . Machining and Monitoring Strategies eter, to monitor minute process interference for the control of critical machining processes, such as: in the development of new cutting tools; or during the ‘ramping-up phase’ from initial workpiece prototyping to that of small batches, or even up to component mass production. 9.8 Complex Machining: of Sculptured Surfaces Introduction In today’s world of rapidly developing CAD/CAM sys- tems with components expected to have a short lead- times coupled with a ‘streamlined appearance’ , many of these parts and assemblies have components that exhibit complex double-curvatures to there shapes and proles. ese so-called ‘sculptured surfaces’ are so-called, because they appear to have been carved, rather than machined requiring complex tool paths to initially machine these convex and concave blended proles. ese sculptured surfaces oen necessitate at least a ball-ended cutter and an 3-axis milling op- eration, or the more likely scenario being by 5-, or 6- axis milling, using a cylindrical cutter’s periph- ery and tip to follow the desired contour. So, with a 5-axis machining operation, the direction of a mill- ing cutter’s axis must be adjusted to the geometry of the workpiece’s surface curvature, by a tilting the tool head (Fig. 245b) to produce a uniform cutting opera- tion, regardless of the complex contour being milled. As we will see shortly in the following section, any free-form surfaces can only be approximated by mill- ing, as a tool path’s groove produced (i.e. its resulting ‘cusp’) is mainly determined by the ‘pick-feed’ 39 (i.e see Fig. 245a). So, free-form surfaces can only be approxi- mated by milling (i.e see Figs. 246a, 247a and 248), as a grooved (i.e. cusped) prole is generated, which is 39 ‘Pick-feed’ , or as it is sometimes represented by the term: ‘a e (p)’ is illustrated in Fig. 247b, which refers to the ‘step-over distance’ for the adjacent tool path. us, the ‘pick-feed’ can vary depending upon the required chip-thickness for the operation currently being undertaken, which inuences the milling tool’s cutter-loading. is cutting force will be modi- ed, as the contoured prole’s geometry also changes during the sculptured surface’s milling operation. strongly inuenced by the width of the pick-feed. is milled contour on the convex/concave, or ‘parabolic curved surfaces’ 40 will necessitate a hand-nishing op- eration aerward (i.e grinding and perhaps, polishing for high-quality die-cast, or injection-moulded ‘end-, or nal-product’ nishes). However, due to the high feedrates that can be utilised when adopting an HSM milling strategy, cutter paths can be much more closely spaced (i.e the pick-feed widths can be reduced). is HSM scenario will considerably reduce the resultant milled cusp height (Fig. 245a), enabling faster, or in some cases no workpiece nishing-o processes being necessary. In the die and mould industry, it is the norm to em- ploy HSM by milling, to produce the cavities and in- tricate features required for the nal cast, or moulded part. A range of cutters are generally utilised to gen- erate and form the workpiece’s free-form proles, ranging from say, tooling such as, large ball-ended slot-drills, down to very small diameter ball-nosed endmills (i.e see Fig. 245c). .. Utilising the Correct Tool for Profiling: Roughing and Finishing In the rst instance and, to ensure that the correct choice of highly-productive cutting tools are employed in the manufacture of for example, a die and mould set, the following logical sequence could be adopted: • e actual geometry of the die, or mould requiring machining operations, should be carefully studied, to identify applicable tools and techniques for metal removal, 40 ‘Parabolic curved surfaces’ , can be mathematically-dened as simply: y = x 2 , although if we want to express this equation in words, we can say that a parabola is: A plane curve generated by a point moving so that its distance from a xed second point is equal to its distance from a xed line. [or] A parabola changes both its radial length with its associated angle and has a unique focal point. NB ‘Parabolic interpolation’ , can be achieved by controlling the milling cutter’s path by interpolation between three xed points, by assuming the intermediate points are on a Parabola. Such machining action, normally requires sophisticated CAM soware to achieve this complex kinematic motional control by the machine tool. (Sources: Smith, 1993, Stroud et al., 2001) Chapter Figure 245. The complex machining of either a sculptured, or die and mould surfaces, will usually necessitate both multifarious and sophisticated programming techniques . Machining and Monitoring Strategies • Dening the minimum radii requirements and the maximum depth of the cavity, needs consideration, ensuring selected tooling can cope with these part geometries, • Approximately estimate the amount of excess stock material from the die, or mould that needs to be removed by milling operations 41 , NB Establishing what is roughing-out and semi- nishing operations, will for a large die-set oen mean that roughing-out is both more ecient and productive on conventional-speed machining cen- tres, with any semi-nishing undertaken by HSM. • In preparation and prior to milling, ensure that the workpiece xturing is both accurate and precise as well as very robust and rigid, otherwise this latter factor in particular, is a classic source for any resul- tant vibrations and will signicantly inuence the tool’s life together with degradation of the die and mould surfaces 42 , NB HSM requires a totally rigid xturing, if vibra- tional tendencies are to be minimised, as it proves disastrous for any long length-to-diameter tool ratios, that are oen utilised for high-speed milling operations. • For the machining processes, they should ide- ally be divided into at least three types of milling 41 ‘Material removal rate’ for HSM milling is generally consider- ably smaller than in conventional machining (i.e except when aluminium and non-ferrous machining occurs). Formula for material removal rate Q = a p � a e � v f ( cm min − ) Where: a p = axial D OC (mm); a e = radial D OC (mm); v f = feed per minute (mm min –1 ). 42 ‘Die and mould milled surface texture’ , by HSM milling op- erations dramatically reduces the manual polishing require- ment – by reducing the resultant milled surface ‘cusp-heights’. Oen conventional milling operations produce relatively large ‘cusps’ (i.e see Fig. 245a – resulting from the large width of the ‘pick-feed’). For example, when a large automobile bonnet (i.e. ‘hood’ – in the USA) die-set has been produced by conven- tional milling practices, any manual polishing activities range between: 350–400 man-hours! NB is order of manual polishing will aect the geometrical accuracy of the die-set. (Source: Sandvik Coromant, 2000) operations, namely: roughing-out; semi-nishing; nishing. NB ‘Restmilling operations’ 43 are normally under- taken during any semi-nishing, or nishing op- erations. .. Die-Cavity Machining – Retained Stock Whenever a rough-milling operation is undertaken with a square-shouldered cutter, this creates the well- known ‘stair-case prole’ 44 (i.e see Figs. 246 a and b) of remaining stock that must now be removed by a semi- nishing milling operation. e die-cavity’s cross-sec- tional prole will signicantly inuence the amount of stock remaining against the cavity wall, which will create a variation in the cutting forces and have an in- uence on tool deection. e consequence of this un- even stock will be that when semi-nishing the prole, it could aect the geometrical accuracy and precision of the die, or mould. Clearly, in the schematic diagram shown in Fig. 246a – le, the large chamfered die fea - ture when being roughed-out for a given D OC , will leave signicant material here for subsequent semi-nish- ing. Likewise, in the cavity of the convex-to-concave prole illustrated in Fig. 246 – right, it has signicant stock material remaining at the lower regions of the concave feature, obviously necessitating a following machining removal operation (i.e. semi-nishing). When a square-shouldered cutter is utilised with a triangular geometry insert, it will have relatively weak corner cross-sections (i.e. by way of illustrating this eect of insert shape strength, see Fig. 155 – bottom), creating a somewhat unpredictable machining behav- iour. Triangular, or rhombic insert geometries, will also create large radial cutting forces and as a result of number of cutting edges, they are unexpectedly, less economical than some other counterparts for such op- 43 ‘Restmilling operations’ , are those milling operations where any Ball-nosed: Slot-drills; Endmills; or in some cases, toroi- dal-geometry inserted cutters; are employed. 44 ‘Stair-case prole’ , is so-called, because it resembles an actual stair-case when taken in cross-section (i.e. see Fig. 246). e height and width of the remaining stock for each step, is de- pendent upon a proportion of the actual ‘step-size’ (pick-feed) and the D OC previously selected. Obviously requiring a semi- nishing operation at the very least, to remove this unwanted material. Chapter erations. On the contrary, round cutting inserts that allow milling paths to be undertaken in any direction, are oen specied because they provide a smooth transition between successive tool passes, while also leaving behind the twin benets of less and more even stock, for later removal in semi-nishing. is residual Figure 246. Die-sinking sculptured proles with a 90° square-shouldered milling cutter, introduces a ‘staircase eect’ on the machined prole. [Courtesy of Sandvik Coromant] . Machining and Monitoring Strategies eect of less additional stock produced by round in- sert’s on the workpiece prole, is shown schematically in both Figs. 247ai and aii and, should be compared to Fig. 246a – this latter eect being the result of utilising square-shouldered cutting inserts, in terms of stock to be removed later in semi-nishing operations. Amongst the notable benets of using round inserts, are that they produce a variable chip thickness, which allows for higher feedrates if compared to other insert- shaped geometries. Round cutting inserts provide a very smooth cutting action (i.e see Fig. 246 – bottom right: inset), because the entering angle changes from almost zero – in the case of very shallow D OC ’s, to that of 90° – under certain conditions with the larger D OC ’s. us, at the maximum D OC , the entering angle is 45° and when copying with the periphery, the angle is 90°. is D OC variability using round inserts, also goes some way in explaining why these inserts are so strong in comparison to other insert shapes. Namely, round in- serts with their actual ‘work-loading’ – at the cut’s ini- tial progression – is successively built-up, rather than almost immediately with inserts having greater enter- ing angles, usually provided by their less-than-robust geometry counterparts. Consequently, round inserts should always be regarded as the primary choice in cutter selection when either roughing, or for medium- roughing operations. When 5-axis machining, the use of round cutting inserts can be usefully exploited, as they have virtually no limitations when machining sculptured surfaces. erefore, with optimum CNC programming, either round inserts, or toroid-shaped milling cutters can normally be substituted for ball- nosed end mills (Fig. 79b), as they can oer: superior cutting performance; improved chip-breaking e- ciencies; as well as better chip evacuation; this latter point is important when deep cavities might otherwise retain work-hardened swarf. Typically, the increases in productivity range between 5-to-10 times better, if compared to that of previously utilising ball-nosed end mills. Round insert tooling is very rigid so as a re- sult, they only produce a small amount of run-out and, when combined with ground, positive and light cut- ting geometries, may be used for semi-nishing and occasionally some nishing operations (Fig. 246 – bot - tom right: inset). Some of the main questions to be answered re- garding the correct application of technology is con- cerned with optimising: the cutting data; likely insert grades; together with their geometries; in relation to the: specic workpiece material to be machined; actual machining operations to be undertaken; anticipated productivity requirements; and the likely workhold- ing restaint/security issues. Die and mould work in- variably involves complex sculptured male and female surfaces, with any calculations of the eective cutting speed being based upon either the ‘true’ , or eective diameter in-cut (‘D e ’ – see Fig. 247b). So, if the D OC is very shallow – as is the case when semi-nishing op- erations are being carried out, then the ‘true’ cutting speed will be much lower (Fig. 247b). If the original cutter diameter was chosen for the cutting data calcu- lations, then for a shallow cut – due to ‘D e ’ being the eective diameter, this drastic reduction in actual cut- ting speed will not have been anticipated, causing the feedrate utilised to be severely compromised, as it is dependent on the calculated cutter’s rotational speed. is will not only severely impede component produc- tivity, but will increase the tool’s potential wear-rate signicantly, this being the case for all round insert cutters, ball-nosed end mills, plus end mills having large corner radii. Due to the adverse and miscalcu- lated cutting data, there is a likelihood for premature cutting edge frittering and chipping – created by too low a cutting speed and localised heat in the cutting zone. When undertaking either nishing, or super-n- ishing of the die and mould sculptured surfaces (Fig. 246biii) on hardened tool steel, it is vitally important to choose tool materials and coatings with ‘hot hard- ness’ capabilities 45 . A major factor to consider when milling for either nishing, or super-nishing hardened steel sculptured surfaces by HSM, is to take shallow cuts. Notably, the D OC should not exceed 0.2/0.2 mm (a e /a p – Fig. 247b). is strategic machining decision should be made, so that excessive deection of the cutting tool assembly is avoided, enabling a high tolerance level and geometric accuracy to be held on the die, or mould. Accordingly, very sti tool assemblies are essential, usually utilising solid cemented carbide: due to its inherent stiness; coupled with the maximum core diameter possible; that the die, or mould part features will allow. 45 ‘Tool materials for: hardened steel milling’ , they are usually coated cemented carbide, with the micro-grain structural matrix (i.e. typical grain size being <1 µm), providing good wear resistance and transverse rupture strength (i.e. this be- ing ‘related’ to its toughness). Coatings can include: titanium aluminium nitride (TiAlN); titanium carbonitride (TiCN); having multiple coatings of between 2 to 12 µm thick, applied by Plasma Vapour deposition (PVD). Diamond-like coatings (DLC) are also utilised. (Source: Dewes and Aspinwall, 1996) Chapter . Mean Grand Mean 0 degree 9 87 941 969 46 966 ↓ 90 degree 978 952 953 26 961 972 180 degree 1014 986 982 32 994 270 degree 968 946 995 49 965 Range 46 45 42 Mean 9 87 956 975 Grand mean → 973 NB. The complex machining of either a sculptured, or die and mould surfaces, will usually necessitate both multifarious and sophisticated programming techniques . Machining and Monitoring Strategies. Sandvik Coromant] . Machining and Monitoring Strategies eect of less additional stock produced by round in- sert’s on the workpiece prole, is shown schematically in both Figs. 247ai and
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