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464 Chapter 9 . Figure 229. Thermal expansion tooling its operation and high-speed turning chuck details Machining and Monitoring Strategies gal loss of the jaws – when rotating at typically high turning speeds (i.e see Fig. 229b for a diagrammatic cutaway assembly of an HSM quick-change chuck). These quick-change chucks incorporate a traditional wedge-style and lever mechanism, that instead of dir­ ectly acting on the jaws, the radial force acts through the actuator and lever mechanism, prior to transferring the effort to the jaws. So, when the chuck rotates at high-speed, the actuators are thrown outward by centrifugal force, but are restrained from moving by the lever, which pivots about the central connected sphere. At the opposite end of the lever, the jaws are also thrown outward and act to move the pivot in the opposing direction (i.e. to that of the actuators) – effectively balancing each other. The performance of the counter-balanced chuck depends upon the accuracy of the balance achieved, as the actuator mass is constant and the top jaw mass being variable depending upon the particular top jaws in use, thus the state of the balance will also vary. As a consequence, the clamping force may fall with rotational speed, or actually increase with heavy and light jaws, respectively. With standard hardtop jaws, the clamping force remains almost constant across the range of the operating speed, making it unnecessary to calculate the clamping force losses. An additional feature is that the static clamping force can be much lower, since there is no centrifugal loss. This lower static clamping force application, has the benefit that when turning either thin-walled, or more delicate workpieces that may otherwise distort with higher clamping forces, such chucks are unlikely to affect these components, when an HSM turning strategy is utilised. Much more could be said concerning HSM turning operations, particularly relating to the calculations and working practices, but it was not the intention here, to give a comprehensive account of such tech­ nical aspects, simply a concise account of the anticipated problems and possible solutions when turning at high rotational speeds. In the following section, a discussion concerning toolholder coupling to the machine tool’s spindle will be briefly reviewed. 9.4.2 Toolholder Design and Spindle Taper Introduction In the past, the taper cone and its associated driving dogs and pull-stud, provided adequate location and 465 torque for the cutter assembly when mounted into the machine tool’s spindle. The tool’s cone taper angle was adequately manufactured so that it perfectly ‘wedged’ into its mating spindle taper and the problem of the single-contact mechanical interface was not really exposed as deficient, until very high rotational speeds were being utilised, coupled to much greater feedrates that the newly-developed tooling geometries and tool materials could now exploit. In recent years, both dual- and triple-contact tooling systems have been introduced, these designs will now be briefly reviewed. Dual-Contact Tool/Spindle Design One of the most significant developments in maintaining a complete mechanical interface between the toolholder and the machine’s spindle was the dual-contact 7/24 taper system24. The CAT Standard incorporates this 7/24 taper, but also allows simultaneous contact on both the toolholder’s flange and taper, when HSM machining is the requirement. By achieving this dualcontact, the CAT-shank toolholders minimise any form inherent imbalance at say, 2,000 rev min–1. However, if the cutter assembly is to be rotated at 10,000 rev min–1, the toolholder must cope with a × 25 increase in centrifugal force, which may compound any unbalance present in the tooling assembly. Further, if the rotational speed is increased still further, into the HSM range, then here, the centrifugal force is × 100 greater and the onset of considerable imbalance may create chattering conditions. At such high rotational speeds, if coolant is utilised in the machining process, the HSM conditions could develop a vortex around the cutting tool, that conventional flood coolant pressures cannot penetrate. In these circumstances, possibly the only realistic option is to utilise a through-the-spindle coolant delivery application at pressures of >690 kPa (i.e. 1,000 psi), coupled to perhaps, micro-filtration of the coolant with special pipes and couplings. The CAT system of dual-contact offers reasonable rotational control of the tooling assembly at moderate-to-high rotational speeds, as the mechanical interface system of face-and-cone provides a certain security against 24 ‘Dual-contact 7/24 taper system’ , refers to the taper being to the 7 inches of taper per 24 inches of length. This 7/24 system incorporates several Standards: CAT and BT 40- and 50-taper tooling. 466 Chapter 9 the onset of imbalance. Typical applications for these HSM dual-contact systems include: aerospace part production; precision die and mould making; automotive component production; as well as medical component manufacturing. It is worth digressing somewhat, to explain the situation of why the single-cone mechanical interface is simply not effective for HSM production applications. When rotational speeds begin to approach 20,000 rev min–1, it is not an unusual occurrence for the singlecontact conventional, or standard CAT V-flange tooling assembly to be effectively sucked into the spindle (i.e. as there is no mechanical contact at the flange), this being the result of a combination of the pull-stud pressure and the machine’s spindle ‘taper swelling’ – due to the very high centrifugal force acting at such high rotational speeds. In fact, this minute amount of ‘taper swelling’ can cause the tool holder to separate from the spindle’s surface and as a result cause considerable damage to both the cone’s male and female surfaces. In order to alleviate this HSM problem and run the tooling assemblies at even faster rotational speeds, the HSK dual-contact toolholders were developed, which will now be briefly mentioned. Hsk Dual-Contact Tooling There are a number of toolholder designs that are alternatives to the conventional steep-taper spindle connection. Probably the most popular version for HSM is the HSK-designed tooling connection (i.e see pertinent HSK tooling details in Fig. 126c). HSK toolholder connections offer simultaneous fitment on both the taper and face, at the front of the spindle. The reason for their acknowledged popularity amongst the HSM machining companies, is because the increased rigidity of the joint, coupled with their inherent reduction in dimensions, compared to the equivalent conventional steep-taper connection. In Fig. 126c, the HSK 8° (included angle) short taper with its gauge face contact and simultaneous taper interference can be seen, which was designed in Germany to Standard: DIN 69893, being introduced in 1993. HSK is a German acronym that translates into English as: ‘Hollow short taper’. Thus, the HSK connection provides: • both high static and dynamic stiffness, • offering great axial and radial repeatable accuracy, • with low mass and stroke, • having inner clamping. Therefore, with all these proven design advantages over conventional spindle connections, it allows the HSK tooling assemblies to utilise the increased rotational speeds necessary for an HSM strategy. Triple-Contact Tool/Spindle Design The triple-contact connection is being offered by a few toolholder manufacturers (i.e. shown in Fig. 230). The triple-contact design relies on an inner expanding sleeve which maintains uniform contact between the machine tool spindle and the: toolholder’s top taper; bottom taper; and flange; this being regardless of the spindle speed employed. Of particular note is the ­inner expanding sleeve which functions particularly well at high spindle speeds. So, as the centrifugal forces increase – with higher rotational speeds, it causes the spindle to grow (i.e. ‘swell’), the toolholder’s spring mechanism forces the split-cone sleeve to proportionally-expand with the spindle. Further, the expanding sleeve also acts as a vibration-dampening device. The expanding sleeve extends the tool’s life on average by between 300 to 500%, by virtually eliminating vibration. As a result of this ‘vibration-free interface’ between the tool and workpiece, it provides smoother machining of: tool steels; aluminium alloys; plus other metallic alloys. This triple-contact connection system, also performs efficiently with extra-long tools (i.e see Fig. 231), notably when utilised on horizontal machining centres. The main reason for the enhanced triplecontact tool’s cutting performance with extended tooling assemblies, is the result of the ‘floating’ inner sleeve (Fig. 230) which acts to minimise any potential Z-axis deflection, thus maintaining its rotational concentricity. Such triple-contact tooling is not inexpensive to purchase, but these toolholders really do amortise their cost, by significantly extending cutter life, while improving part production rates. Further, it is claimed by the tooling manufacturer that the toolholder is ‘maintenance-free’ , while its spring-mechanism in ‘life-testing’ has achieved upward of one million tool changes. With the advent of either the double- and triple-contact systems, enabling contact between the machine tool’s spindle and the toolholder’s mechanical interface: top-taper; bottom-taper; plus flange; while ‘eliminating vibration’; this has been achieved under the unique conditions that arise with today’s HSM and high-accuracy and precision manufacturing needs. Machining and Monitoring Strategies 9.5 Dynamic Balance of Toolholding Assemblies Introduction Balancing tools that are intended for HSM applications is vitally important and there are quite a few In- 467 dustrial/Manufacturing engineers and users who do not really understand the concept of how to achieve balanced tooling, or why it is really necessary. Either very long extended tooling required for say, for deeppocketing (Fig. 231), or tooling that is out-of-balance, will more than likely produce: chattering effects; gouging of a step, or face; loss of workpiece accuracy and precision; not to mention uneven and premature cutter wear. Whenever a new tooling assembly is destined . Figure 230. Triple-contact tool connection system is ideal for any potential HSM ope­rations. [Courtesy of Heartech Precision Inc. (HPI)] 468 Chapter 9 . Figure 231. Tool runout (≥10 µm) should be of prime importance when machining deep pockets. [Courtesy of Sandvik Coromant] Machining and Monitoring Strategies for HSM applications on a workpiece, a balancing oper­ation needs to be undertaken, this statement is also true for many sub-HSM applications, particularly when extended tooling is used for whatever reason (Fig. 231). In fact, every rotating object (i.e. chuck, or tooling assembly, etc.), will generate vibration. As has been explained in the previous section, this vibration results from a number of sources, but principally here, from centrifugal forces produced by the rotation of an unbalanced mass. There are several types of unbalance that could arise, but here, we are mainly concerned with what is termed dynamic unbalance, which increases by the square of the rotational vel­ ocity. For example, any vibration produced by a tooling assembly at 3,000 rev min–1, is × 100 greater than an identical tooling configuration that is rotating at 300 rev min–1. Moreover, what is often either misunderstood, or indeed overlooked, is that any change to the tooling assembly – no matter how small it might seem, requires re-balancing! These tooling modifications include any occasion when a cutting tool is adjusted, or changed, or similarly if the toolholder is also either adjusted, or changed. Such changes to the ‘status quo’ of the tooling, will directly affect its ensuing balance, even minutely when just a ‘few microns’! So that, these miniscule changes to the tooling’s dynamic condition, causes a degree of tooling oscillation, hence an out-of-balance condition – with the likely problems that this creates. With the wide variety of tooling that is held in: tool storage carousels; magazines; turrets; etc.; they must all be ‘balanceable’ by some means. A range of balancing techniques can be employed here for either single-, or dual-plane balancing – more will be said concerning these effects will be made in the following section. The techniques utilised in achieving tool balance could include: • ‘Hard-balancing’ (i.e. see Fig. 234b) – when the complete assembly either has to have material removed, or added at a certain part of its assembly. NB The major problem associated with ‘hard-balancing’ is that if the tooling setup changes, so will the likely rotating mass change, which will mean modifying the amount of material to be either added, or subtracted from this newly-distributed mass, • ‘Adjustable balancing rings’ (i.e. see Fig. 232) – by rotating the twin lower and higher balance rings 469 either clockwise, or anti-clockwise they minutely modify the balance-condition, allowing singleplane balance to be achieved. NB These matched pair of balance rings are in a symmetrical state of unbalance (i.e. they are both ‘unbalanced’ to the same degree). Letting the user adjust the pair to counter any unbalance in the cutting tool/toolholder assembly and locking them into place – usually achieved on commerciallyavailable balancing machines (i.e. see Fig. 234a). The state of unbalance is not merely a subject to the ‘caprice’ of the machine tool operator, a tool assembly’s balance is given by various quality Standards, such as ISO 1940/1, or ANSI S2.19 – being basically exact reflections of each other. In the following related sections, they deal with how and in what manner rotating cutter assembly balance is achieved, utilising such HSM balance calculations and associated graphical details as necessary, from these Standards. 9.5.1 HSM – Problem of Tool Balance Unbalance of a rotating body (i.e. here we are concerned with a complete tooling assembly), can be defined as: ‘The condition existing when the principal mass – axis of inertia – does not coincide with its rotational axis’ (i.e. shown schematically in Fig. 232). For example, such an undesirable state of affairs can be comprehended by considering the following situation: if a φ50 mm face mill assembly is rotated at 15,000 rev min–1, it will produce a peripheral speed >240 km hr–1, which may prove to be disastrous if it is unbalanced! Basically there exists, three types of unbalance conditions for rotating assemblies – such as tooling, these are: 1. ‘Static unbalance’ – single-plane. This type of unbalance occurs when the mass does not coincide with the rotational axis, but is parallel to it and the force created by such unbalancing, is equal to the magnitude at both ends of the rotating body. Thus, if some relief – metal removal (i.e see Fig. 234b) – on the toolholder body equal to the out-of-balance mass that occurs, then a nominal static unbalance is achieved, 2. ‘Couple unbalance’ – Under these circumstances, the cutter assembly – mass axis – does not coincide 470 Chapter 9 . Figure 232. The taper fitment against runout/eccentricity for a milling cutter and its associated balanced toolholder Machining and Monitoring Strategies 471 with the rotational axis, but intersects it at the centre of gravity of the ‘assembly’s body’. Under such conditions the force vectors equalise, but are 180° apart. 3. ‘Dynamic unbalance’ – dual-plane. Such a condition of the toolholder assembly arises when the axis does not coincide with the rotational axis and is not either parallel to, nor intersecting this axis (i.e. see Fig. 232). For any rotating tooling assembly, estimating the cutter unbalance is possible using the following variables: M = cutter/holder mass, S = mass centre, e = displacement of mass centre, r = distance from centre of tooling, to the centre of gravity of mass (m), ω = angular velocity, m = mass unbalance, U = cutter unbalance, 9549 = a constant. Determining the relative unbalance (U) of a rotating tooling assembly, can be found by the following expression(s): U = M × e or, alternatively: U = m × r (i). It is usual to express unbalance in terms of the product of the mass times distance, typically using the units: ‘g-mm’. Finding the magnitude of centrifugal force produced by the rotating tooling assembly with a given unbalance, can be established as follows: F = U × ω2 (ii). Where: ‘ω’ is the angular velocity in units of radians sec–1. The formula to find ‘ω’ is expressed by: ω=  � π � r pm  (iii). Therefore, by combining formulae: (i) and (iii), in (ii), we can obtain the magnitude of centrifugal force ‘F’ , as follows: F = m × r × ( 2 × π × rpm/60)2 (iv). As established in equation (iv), the centrifugal force caused by tooling unbalance will increase by the ‘square of the speed’ , in a similar manner to the spindle nose taper swelling (i.e. growth) previously mentioned. Nonetheless, assuming that this specific toolholder initially has a low unbalance, this will become a problem if the rotational speeds are increased beyond 10,000 rev min–1. For example, with most toolholders exhibiting single-plane unbalance25, research experimentation has shown that the initial unbalance of a typical tooling assembly will be of the order: 250 gmm. When such tooling is rotated at 15,000 rev min–1, this 250 g-mm of out-of-balance develops a continuous radial force of 642.6 N. Unbalanced tooling can introduce considerable detrimental effects on not only the machine tool – this high centrifugal force causing internal bearing stresses leading to premature spindle failure, but affects cutter life and degrades workpiece surface texture. Much of the principal tooling unbalance problems can be traced-back to several sources, such as: • Toolholders of the V-flange type, which might have different depth of drive/slots, these toolholder features being part of the inherent design, • Toolholders for some end mills and slot-drills, having set screws for locking the cutter securely in place, so due to necessary clearance and the radial application of the set screw, this creates minute cutter eccentricity – causing unbalance, • Out-of-balance caused by an unground V-flange base, • Collet and its collet nut tend to be recurring sources of unbalance in HSM tool holders. NB Most of these tool holding-related issues can be eliminated by simply modifying the tooling design. As can be seen from Fig. 232, the marginally eccentric adjustable balance rings can be rotated to adjust the degree of single-plane balance, with several of the tooling manufacturers offering differing adjustment methods for HSM toolholders. Finally, consideration needs to be given to the level of balance-quality required and in HSM applications for example, a milling cutter is expected to withstand 25 ‘Single-plane unbalance’ , relates to the type of unbalance that occurs in either one of two planes. Namely, the tooling assembly’s single-plane unbalance will be in either its axial, or radial directional plane. 472 Chapter 9 both high rotational speeds and associated cutting forces, thus here it can be considered as a ‘rigid rotating body’. This assumption allows one to use the ANSI S2.19-1989 Standard, for achieving balance – see Fig. 233, which defines the permissible residual unbalance of a rotating body relative to its maximum speed. This Standard and its equivalents (e.g. ISO: 1940:1; ISO: 1290 G), assigns different balance-quality grades termed: ‘G-numbers’ , related to the grouping of rotating bodies (i.e. not shown), these groupings being based upon the experienced gained with a variety of: sizes; speeds; and types. Thus, the balance-quality grade ‘G’ , equals the specific unbalance ‘e’ times the rotational speed ‘ω’ , as follows: Balance-quality G = e × ω (mm sec–1). Furthermore, the equation was described earlier, thus: e= U M (i) ∴ solving for ‘U’ , we obtain: U=  � M � G r pm (v). From the Standard, the balance-quality for machine tool drives is given as: G2.5, although in many instances the value utilised should ideally approach that of G1.0 – this being the specification for grinding machine tool drives, as today in HSM applications they are compatible. However, if for the purposes of clarification of the unbalance tooling condition the value of G2.5 is utilised, then the following worked example illustrates the balance-quality necessary using a toolholder weighing 3 kg, rotating at 25,000 rev min–1: U (higher) =  �  � . (g-mm) ,  ∴ U (higher) = 2.85 g-mm. As alluded to previously, this unbalance condition is the ‘worst case’ and the tooling should ideally approach G1.0, this balance-quality value, gives: U (lower) =  �  � . (g-mm) ,  ∴ U (lower) = 1.14 g-mm. This then follows that the balance is between 1.14 and 2.85 g-mm, which is toward the ‘upper-end’ for the maximum residual specific unbalance for the G2.5, while approaching this level for the G1.0 (i.e shown by the graph in Fig. 233). Even when the tooling assembly has been dynamically balanced in both planes (i.e. see Fig. 234a – more to be said on this topic shortly in Section 9.5.2), problems still exist, particularly in the fit of the spindle taper connection (Fig. 232). This is a result of the taper rate accuracy requirements between both the shank and taper socket. In fact, the situation is quite a confusing one, due to the relative cone ‘Angle Tolerance’ grades: AT-1 to AT-6, that are employed using the conventional fitment of: 7:24 taper. Not only do different countries often have their own connection Standards, but previously, even individual machine tool manufacturers within each country had adopted differing Stand­ards! Today, many machine tool companies tend to utilise taper spindle connections that are compatible to an appropriate Standard and complement those of the tooling manufacturers. 9.5.2 HSM – Dynamic Balancing Machine Application It has been discussed in the previous sections that cutting tool assemblies when combined with an HSM strategy, can be a large contributor to dynamic unbalance. For instance, in the production and manufacture of say, the geometry of a face-mill, the tooling stock material is: externally/internally turned on one side; unclamped; flipped-over and rechecked; then turned on the other side; then located onto a milling machine tool for operations on the individual insert pockets that must be milled; and indexed26 – as appropriate for the number of cutting edges; this necessary clamping/reclamping workpiece (i.e. face-mill) procedure, will create a tool that is marginally-unbalanced. With HSM, the otherwise unnoticeable unbalance at conventional rotational speeds, becomes intolerable in these high-speed ranges. Often, the most economical technique for achieving balanced tooling for tooling 26 ‘Insert pockets’ , are sometimes ‘differentially-pitched’ which means they have unequal spacing of teeth around the cutter’s periphery. This pitching technique for cutting insert pockets, is quite effective as a means of reducing machining vibrational effects often encountered with coarse-pitched face-mills. Machining and Monitoring Strategies . Figure 233. A graph to determine high-speed cutter unbalance ‘U’ (ANSI S2.19–1989) 473