the cutting insert, or tool they are most familiar with, because they know – from practical experience – that it performs and wears in a progressive manner, rather than the unpredictability associated with an insert of ‘uncertain machining capability’ that might otherwise prematurely fail. Prior to discussing criteria for determining when a cutting insert is ‘worn-out’ , it is necessary to estab - lish in practice, what this actually means. For example, does ‘worn-out’ refer to when the: dimensional accu- racy becomes unpredictable: or if the surface nish has signicantly deteriorated; or perhaps the fact that its automatic chip-breaking behaviour has become inef- cient? In many situations it is by the user’s experience that one can judge how much ank wear can be toler- ated on the cutting edge before machining is discon- tinued. As a rule, ank wear is a dependable criterion for assessing when the cutting edge is eectively worn- out. Moreover, from the previous discussion, perhaps the degree of cratering may in certain machining cir- cumstances prove to be more signicant than the ank wear, in respect to the shortening tool life. Tool wear can be established by several techniques, but the usual method is to observe and then measure the actual wear as it progressively develops. e eec- tive cutting time, or tool life ‘T’ , is specied as time- elapsed prior to a predetermined degree of wear has been reached. A typical procedure for determining ank wear can be: to observe cutting edge(s) in-situ on the machine tool; then remove from the machine and visually inspect the tooling; followed by its respective wear rate can then be optically magnied in suitable equipment allowing accurate dimensional measure- ment – against the following criterion (i.e. see Fig. 174): • Extent of ank wear from original edge – if this wear is of relatively uniform nature, it may be dis- tributed across three zones, ‘A’ ,‘B’ and ‘C’. e mean ank wear ‘V B,C–A ’ is measured over the cut- ting region of the leading edge across these zones – it is oen just referred to as simply: ‘V B ’. If excessive wear develops at one position on the cutting edge, for instance where the wear-notch ‘V N ’ occurs, this zone is usually ignored when establishing the ‘mean wear’. Here and under these conditions, it is usual to quote the maximum ank wear as ‘V Bmax ’ , • Extent of cratering – this is usually specied by the maximum crater depth from the plane of the original rake face ‘K T ’ and in some cases, by its di- mensional size: ‘K B ’- width and ‘K M ’ – length (not shown). e above wear criteria, are normally utilised for esti- mating the extent of ank and crater wear. Over many years of experimental research into tribological wear mechanisms, it has been established that progressive ank wear develops according to a xed pattern, with three distinct stages to this wear regime, they are (Fig. 176): 1. Initial, or primary wear – if a new cutting edge is used to machine a workpiece, there is a rapid breakdown of the of the cutting edge. is early ank wear on the tooling is depicted in the graph of wear against time in Fig. 176a, indicated by its preliminary high wear-rate, is wear-rate is de- pendent upon the cutting conditions and type of workpiece material, plus any cutting uid applica- tion – if utilised. Flank wear increases in relation to an higher cutting speeds, 2. Progressive, or secondary wear – occurs aer the initial ank wear has taken place. During the fol- lowing time period, there is a steady and progres- sive stage to the cutting tool’s/insert’s wear, with a much less pronounced increase than that indicated at the initial wear stage, this is when the productive machining output occurs. Toward the end of this progressive wear stage, this being the case when the ank wear ‘V B ’ reaches approximately 0.8 mm in height, here, it is normal practice to replace this old tool with a ‘sister tool’ – to continue machining the component batch, or production run. Once ank wear has reached this arbitrary dimensional value, then to all practical purposes its productive life is ended, 3. Catastrophic, or tertiary wear – will normally only become apparent if the tool is taken toward, or up to, its complete failure. Such catastrophic failure is the result of a combination of several tool wear mechanisms: high ank wear; large crater forma- tion – reaching the point where the tool has been suciently weakened for the increased tool forces now operating to cause it to fracture. Inevitably, if such an immediate breakdown occurs during the nal pass over the workpiece’s surface, it is prob- able that the component has to be scrapped. If the workpiece has a high residual raw-stock value, then aer machining, signicantly more added-value will have accrued. So, any initial savings made by using these tools into the tertiary ank wear stage, will be more than cancelled-out by scrapping this component! Machinability and Surface Integrity Tool-life Diagrams Machinability is a subject that has yet to be fully-de- ned and analysed, in particular the interactive mech- anisms that take place at the chip/tool interface, with the user’s own experience being a good start-point for any future machining operations. As has been men- tioned above, tool wear varieties can have several dif- fering causes and eects (i.e. see Appendix 11). With any machining batch, or production run, it is custom- ary practice to establish a ‘norm’ for both the tolerable ank wear dimension and the depth/size of crater for- mation. In particular, as the ank wear pattern usu- ally takes in-cut time to progressively develop and this predictable tool/wear relationship has been well estab- lished some years ago, initially by F.W. Taylor’s pio- Figure 176. Tool wear under steady-state conditions: (a) tool wear as a function of time, (b) if cutting speed is changed, then tool life is aected, (c) amalgamation of these ‘Taylor curves’ and derivation of the ‘general taylor curve’. [Courtesy of Sandvik Coromant] . Chapter neering work, then further rened over the years. As can be seen from the graph of ank wear (V B ) against time (T) 69 shown in Fig. 176a, tool wear does not usu- ally follow a straight-line relationship’. Invariably, the ‘V-T curve’ for ank wear initially develops quickly then settles to a moderate growth over a reasonable time-period, then has a rapid escalation – almost ex- ponentially – at an ‘end-point’ as it catastrophically fails. e actual plotted wear curve prole and its as- sociated inclination angle will vary depending upon the cutting speed selected, with individual cutting speeds having specic wear curves (i.e. see Fig. 176b). So from the graph in Fig. 176b, it can be visually-es- tablished that the higher the cutting speed utilised, the greater the ank wear. e composite graph depicted in Fig. 176c – le (i.e. taking ‘standardised’ ank wear @ times: T 1 to T 5 – from Graph 176b), shows that a direct relationship exists between logarithmic time (logT) and cutting speed (logV C ). e features that characterise this ‘straight-line’ are its position and gra- dient, these values can be expressed through the ‘gen- eral-case’ tool formula 70 developed by Taylor (1907), as follows: V T α = C 69 ‘Cutting time’ (T) here, is the tool-life of the cutting edge, be- fore a specic amount of ank wear ‘V B ’ is established. 70 is ‘general case’ Taylor formula, has been expanded and developed which denes the machining characteristics with more mathematical rigour, including the eects of: feed- rate; D OC ; as well as component hardness, as follows: V T α f m d p H q = K T ref α f ref m d ref p H ref q Where: ‘f’ = feedrate (mm rev –1 ), ‘d’ = D OC (mm), ‘H’ = Hardness (e.g. HR C ), with ‘m’ , ‘p’ and ‘q’ are exponents whose values are experimentally-established for the production operation, ‘K’ = a constant analogous to ‘C’ , while ‘T ref ’ , ‘f ref ’ , ‘d ref ’ , and ‘H ref ’ are the reference values for feedrate, D OC – when they are <1.0*. *is 1.0 numerical value, indicates the greater eect of cutting speed on tool life, since the exponent of ‘V’ is 1.0. Moreover, aer cutting speed, the feedrate is the next in importance, so ‘m’ has a value > ‘p’**. **e exponent for work hardness ‘q’ is also <1.0.In reality, there are diculties in the application of the above equation for practical machining operations, due to the vast amount of machining data that is necessary to determine the parameters of this equation – producing considerable statistical variance. In order to reduce the variability, while making the overall equation more manageable, the D OC and hardness parameters, reduce the equation to the following expression: V T α f m = K T ref α f ref m Where: ‘terms’ are the same, but the parameter ‘K’ will have a slightly dierent interpretation – see the available literature for a more rigrous mathematical treatment on the subject of tool life. (Groover et al., 2002) e two constants (α, C), can be established graphi- cally from the graph (Fig. 176c – right) of the ‘plot - ted’ sloping straight-line gradient. Hence, the value of constant ‘α’ , can be obtained graphically from the trigonometrically relationship of the respective values of the ‘X’ and ‘Y’ coordinates. Likewise, the other con- stant ‘C’ , may be found by extrapolating this sloping line down to the cutting speed axis (logV C ), for its nu- merical value. Tool Costs e tool cost ‘CT’ will normally consist of the sum of the purchase cost, grinding costs – where applicable, as well as tool-changing cost for each machined com- ponent. In Table 12 (below), it tabulates how tooling- Table 12. Calculating the cutting-tool cost per cutting edge Costs for: cutting tool, edges and regrinding: Sum of costs: Initial cost of tool = (A) Number of cutting edges per tool = (B) Tool cost per cutting edge = (A/B) = (C) Cost of insert = (D) Number of cutting edges per insert = (E) Insert cost per cutting edge = (D/E) = (F) Machine charges per hour = (G) Tool changing time (minutes) = (H) Cost per tool change = (G x H/60) = (I) Regrinding charges per hour = (J) Regrinding time (minutes) = (K) Regrinding cost = (J x K/60) = (L) Tool cost per edge = (C + F + I + L) = (M) Number of components per cutting edge = (N) Tooling cost/component (i.e Cutting tool cost per edge) = (M/N) = (C T ) . Machinability and Surface Integrity costs can be calculated, for a simple turning operation: the method can be modied for machining centres and for most other machining operations. As can be seen from Table 12, the cutting s[peed has a major impact on the value of ‘CT’ , because an increase in cutting speed normally results in faster tool wear rates, with the tool charge per component increasing as a result. Tool costs today, now account for only a small proportion of the total costs of pro- duction, owing to the fact that the latest tooling can operate at higher feeds and speeds than their earlier counterparts. When the tool costs actually rise – due to greater wear rates as a result of increased cutting speeds, it naturally follows that associated tool perfor- mance will also increase. For any machining operation there exists an ‘eco- nomical tool-life’ (T e ), which can be calculated from the following formula: T e = ( α − )( C T C m + t C ) Economical tool life Where: T e = Economical tool life (minutes), α = Slope of the V-T curve (i.e. measured from graph), C T = Cutting tool cost per edge (i.e. obtained as de- scribed above), C m = Machine tool, labour, and related overhead costs – charged per minute, t C = Tool-changing time per minute for operation in question*. *is tool-changing time will vary depending upon whether the chosen cutters are of the ‘conventional‘, or ‘modular quick-change’ tooling varieties. In Fig. 177a, the tool-life at ‘maximum production rate’ (T q ) is shown, which is a variation of the calculation given above for ‘economical tool life’ (T e ), where the variables are identical, but in the former case a higher cutting speed is employed, resulting in shorter tool- life. Even though the lowest possible machining cost per component can be calculated with the most eco- nomical cutting speed, it is oen desirable to utilise a faster machining strategy. is increased speed, will involve supplementary costs, although it can only be warranted if higher production output results. If the number of components per hour (P r ) is plotted (Fig. 177d) in relation to the cutting speed (V c ), a repre- sentative curve will result. is curve is redrawn and shown in Fig. 177e, this now being a ‘composite’ of the sum of the: machine/labour/overhead costs (Cm); with tooling cost (CT). e zenith of this curve (Fig. 177e), represents the highest production rate (P r max). While the cutting speed (V q ) is associated with the peak of the curve, which is greater than the most economical rate (V e ), with the values between these two points, representing the ‘high-eciency range’ for a particular operation. In Fig. 177e, the additional vertical axis depicted, represents the production rate (P r ) – this being the number of components machined per hour, it can be calculated in the following manner: P r = 60 (1 – t C/ T)/tp Where: t C = Tool-changing time per minute for operation, T = Tool life, tp = Total time per component (i.e. including: ma- chining, handling and down-tome). e relationships mentioned above represent theoreti- cal associations. So, some caution should be applied when using these factors and they need to be treated as a ‘starting-point’ only for both the values and trends represented here. Moreover, they are subject to vari- ability, due to the complex relationships and interac- tions found during machining operations. .. Return on the Investment (ROI) As an alternative approach to the above mentioned cutting tool costs and production output interactions, is to relate any productivity improvements to both the actual machinery cost and the total invested capital – to achieve this level of manufacturing yield. In the previously described machinability tests, the work seldom considers rates of production, or relates the ndings to actual increases in the total economics of production. A signicant scal argument is that any gures obtained from such testing, should highlight the ‘im- proved’ ROI, which can be obtained by any manu- facturing company utilising the latest tooling, in conjunction with the application of ecient cutting conditions. e following simple formula can be utilised to cal- culate the ROI for a particular: production operation; machine tool; machining cell; etc.: Chapter ROI = (T S )(M C )/(MT I ) Where: T S = Time savings per year (i.e. in hours), M C = Machine tool charge (per hour), MT I = Machine tool investment. In this section, only a supercial treatment has been given to the economic argument relating associated capital equipment costs and their overheads, to out- put productivity. More intricate and sophisticated eco- nomic models can be obtained in the literature. 7.8 Cutting Force Dynamometry Introduction During machining operations, plastic deformation, friction between the tool and workpiece, together with micro-fractures and -ssures occur. ese mechanical phenomena produce measurable cutting and forming forces with very high-frequency acoustic emissions (AE). e application of AE in association with other Figure 177. The correlation of typical manufacturing cost factors: machining costs, together with their resultant productivity. [Courtesy of Sandvik Coromant] . Machinability and Surface Integrity sensors, such as: force transducers, accelerometers can be coupled to neural networks to give a psuedo-form of articial intelligence (AI) – more will said relating to cutting tool monitoring and analysis in a succeeding chapter. Many of the early attempts at cutting force monitoring were by using several strategically-placed resistance-type of strain-gauged mechanical elements. ese strain-elements were designed so that at a par- ticlur portion of their geometry they could either minutely: buckle, bulge, or twist – well within their elastic limit. At the positions of greatest sensitivity on these mechanical elements, strain-gauges were se- curely placed and wired into a ‘Wheatstone bridge re- sistance circuit’ 71 , which as the gauges distorted they changed their micro-resistance, which could then be fed through suitable instrumentation. ese strain- gauged elements could be calibrated against ‘known’ mechanical devices (e.g. ‘proving-rings’ 72 , or similar), 71 ‘Wheatstone bridge resistance circuits’ – invented by Sir Charles Wheatstone: circa 19 th Century, for accurately measuring resistance in an electrical circuit. Simply, a ‘bridge circuit’ consists of: four resistances; a galvanometer; with a d.c. power supply. In essence, in these highly sensitive resistance ‘bridges’ they are used to detect minute changes in strain gauge resistance. A typical ‘full-bridge’ consists of the four resistors: ‘R 1 ’; ‘R 2 ’; ‘R 3 ’; ‘R 4 ’; suitably coupled to the galvanometer and d.c. supply. Typically, the most simple strain-gauged circuit would consist of: a resistance ‘R 1 ’ which here for argument, is the gauge used for strain measurement. Resistance ‘R 4 ’ is a second strain gauge which here, could remain at constant resistance. e other ‘half of the bridge’ , resistances ‘R 2 ’ and ‘R 3 ’ are variable resistors which by adjustment, are utilised to ‘balance’ and ‘rebalance’ the bridge (i.e employed to reduce the current across the galvanometer arm to zero). erefore, when the ‘bridge’ is ‘balanced’ , the ratio of the gauges and the variable resistances are equal, thus: R 1/ R 4 = R 2 /R 3 ∴ R 1 = R 4 × R 2 /R 3 (Collet and Hope et al., 1974) 72 ‘Proving-rings’ , are laboratory calibrated and certicated me- chanical device, normally consisting of: a steel ring; dial gauge and loading pads. It is usually employed in the calibration of force-measuring systems – only within its maximum permis- sible load. Such ‘proving-rings’ can be manufactured for either high sensitivity – for strain-gauge applications, or for more robustness – when calibrating tensile testing machines. In practice, the steel ring if compressed, allows the diameter to minutely contract in direct proportion to the applied force, with its deection accurately measured by a dial gauge located across the centre of the internal portion of the proving-ring’s diameter. Changes in the dial gauge readings, can be converted to force measurement by means of a suitable calibration graph, or more simply, by multiplying the gradient of the graph – as the graph produced has a straight-line relationship. (Ramsey et al., 1981) allowing the resolved cutting forces during subsequent machining to be data-logged, for suitable in-depth analysis by the user. Strain gauge dynamometers based upon the Shaw and Cook (1954) model, normally re- quire several design criteria to be addressed, if they are to perform satisfactorily, these factors are: 1. at the dynamometer should have a sensitivity of 1% of its mean designed force, 2. Such a dynamometer requires a natural frequency of at least 4 times the ‘forcing frequency’ , 3. e strain-gauged circuit elements should produce the minimum of cross-coupling (i.e. ‘cross-talk’ is <2%) – when calibrated. NB is latter point, can be assessed by a range of calibrated ‘proving rings’ , or ‘torque arms’ – if required to measure torque eects in the circuit, thegraphical calibration should indicate: both plot- ted linearity and also be coupled to minimal hys- teresis 73 . Today, most multi-axes cutting force dynamometers utilise sensing elements, based upon the piezoelectric eect 74 and these ‘active sensors’ , will now be more fully discussed. Piezoelectric Dynamometers ese high-rigidity force transducers provide an elec- trical output signal under the eect of direct element deformation. Hence, element deformation can be kept several degrees of magnitude smaller than that of the ‘passive systems’ – such as those utilising strain-gauged elements. With most ‘dynamic systems’ such as those employing quatrz-based elements, their inherent de- gree of rigidity and a broad measuring frequency range creates smaller measurement interference and 73 ‘Hysteresis loop’ is an area bound between the loading and unloading paths (i.e. typically found in a stress-strain curve), indicating energy dissipation, or damping. 74 ‘Piezoelectric eect’ , was discovered by Pierre and Jacques Curie in 1880. A piezoelectric material (e.g. quartz, or Ro- chelle salt) is a special kind of insulator which, if compressed along one of its axis, acquires an electrostatic charge on the material’s opposite faces. Hence, when such material is ac- curately and precisely cut to the desired shape, it acts as a piezoelectric transducer, hence its input is force and its output is charge. ese piezoelectric elements, can be suitably posi- tioned and arranged and thus, used in dynamometers for dy- namic cutting force measurement. Chapter as a result, oer extremely fast process response, in comparison to those of the of the ‘non-rigid type’ – having long measurement paths (i.e. found in con- ventional strain-gauged elements). Unlike ‘passive sys- tems’ utilising strain-gauges, it is virtually impossible to perform static measurements by using piezoelectric transducers, even though an electric charge delivered under static load can be registered, it cannot be stored for any realistic time period 75 . e design of most of today’s multi-component dynamometers use piezoelectric elements which are quintessentially comprised of a stack of quartz discs, or plates with accompanying electrodes being installed into a stainless steel housing (Fig. 178ai and aii). Every disk, or plate has been precisely cut in a denite crys- tal axis, with their sensing orientation coinciding with that of the axes of the force components to be measured (Fig. 178). In practice, the electrodes can collect the ‘charge’ on their respective quartz disk’s surfaces, these being suitably ‘hard-wired’ to their appropriate and corresponding plug connectors. As shear forces can only be transferred by frictional contact, a certain minimum of friction is essential between the quartz disks, electrodes and housing. Depending upon the shear force magnitude (i.e. measured) a more-or-less high pre-loading of the system must be generated by a pre-loaded bolt. is act of preloading the system is absolutely essential and is usually undertaken when the force transducer is initially installed into the dyna- momoter by the manufacturer. A typical three-component dynamometer (Fig. 178ai), consists of sensors with two shear quartz pairs – namely for ‘F X ’ and ‘F Y ’ , plus one pressure quartz pair – for ‘F Z ’ , assembled in a suitable housing. Each quartz pair has two identical plates stacked with a common 75 ‘Piezoelectric storage’ – for static loading, this cannot be achieved because an insulating material would have to have innitely high resistance, together with ampliers that are perfectly free from any form of leakage and ‘non-operate cur- rents’ , with an amplication factor of innity! e dri in today’s charge ampliers is below ±0.03 pC s –1 . In static mea- surements performed with ‘load washers’ , this means that in practice, the zero-shi is limited to within ±10 mN s –1 . For example, if a static load of 10 kN is measured for one min- ute, then aer this time, the result of the measurement can only be invalidated by a maximum ±0.6 N, that is by ±0.006%. Hence, it is a simple task to piezoelectrically measure large forces for minutes, or hours, but small forces can only be mea- sured ‘statically’ for very short time-periods. us, piezoelec- tric transducers are normally referred to as being: ‘quasistatic measurement elements’. electrode between them, oering twice the sensitivity. is three-component dynamometer is constructed with four of these three-component sensors mounted in parallel between the base and top plate, being as- sembled with a high preload. Given that the outputs from the four sensors are in the form of an electrical charge, they are able to be interconnected within the dynamometer body. With this particular sensor ar- rangement, it is possible to obtain up to eight charge outputs from the dynamometer. e four-component dynamometer shown in Fig. 178aii, has several shear quartz plates arranged in a cir- cle (i.e. top circular element), with their sensitive axes being tangential, allowing an element to be formed responding to a moment ‘M Z ’. In this dynamometer, the four-component sensor is obtained by assembling this element (i.e. ‘M Z ’) in a housing, together with two shear quartz pairs – for ‘F X ’ and ‘F Y ’ , plus one pressure quartz element pair for ‘F Z ’. So, by mounting this sen- sor assembly under a high preload between the base and the top plate, it results in a four-component dy- namometer, capable of simultaneous measurement of: ‘F X ’ , ‘F Y ’ , ‘F Z ’ and ‘M Z ’. For both accurate and precise machinability and data-gathering assessment, these invaluable piezoelec- tric dynamometers oer the following advantages and typical properties: • High rigidity (‘c x ’ , ‘c y ’: >1 kN µm –1 , ‘c z ’: >2 kN µm –1 ), hence producing a high natural frequency (‘f o ’: ≈3.5 kHz), • Wide measuring range (-5 to 10 kN), • Extreme linear sensitivity (‘F X ’ , ‘F Y ’: ≈-7.5 pC N –1 , ‘F Z ’: ≈–3.7 pC N –1 ) and virtually free from hyster- esis (≤0.5%FSO), • Minimal cross-talk (≤±2%), • Environmentally-protected (IP67), typically sealed against of both cutting uids and debris ingress. NB Such high-quality apparatus is not cheap to purchase, therefore it should be carefully main- tained and looked aer, to ensure an extremely long life and fail-safe operation. Piezoelectric dynamometers can have their quartz sensing elements arranged to t into platforms (Fig. 178b), for tment onto a turning machine tools turret with the cutting tool suitably arranged for turning op- erations. ese type of dynamometer platforms can be located on the machine tool’s bed, with the workpiece clamped onto the dynamometer for milling, drilling, or grinding operations to be used for specic types of machinability investigation. In the former case, it is Machinability and Surface Integrity Figure 178. Multi-axis non-rotating dynamometers, used for: milling, drilling and turning experimental data-gathering and analysis. [Courtesy of Kistler Instrumente AG] . Chapter not possible to index the turning centre’s turret, due to the nature of the electrical couplings to the plat- form, but this problem can be overcome by mounting a dierent dynamometer conguration, situating the sensing equipment within the turret – as depicted in Fig. 179a. Here, the installation of an acoustic emission sensor (AE) 76 behind the turret in combination with a multi-component force sensor mounted in the tur- ret’s pocket – this force equipment having previously required the necessary of preloading which was pro- vided by a suitable ‘preload wedge’. In Fig. 179b, are exhibited the measuring results from the sensor installation decribed in Fig. 179a – for a longitudinal (external) turning operation, with the cutting force components and the AE signals being simultaneously recorded. e resulting graphs pro- duced in both Figs. 179bi and bii, show the AE rms 77 and force signals in the case of a tool breakage. e tool breakage can be readily seen in both signal traces. Figs. 179biii and biv, show the AE and force sensor sig- 76 Acoustic emission sensors (AE), in metal machining applica- tions usually capture frequencies in the range of 50 kHz to >1 MHz in range, this being a usual aid for any form of in- process monitoring operations. By using a combination of AE and force monitoring, this has been shown to be a means of condition monitoring of the cutting tool’s state – more will be said on this topic later. In metal cutting operations AE oc- curs due to plasto-mechanical processes of crack formation and chip removal, in combination with surface friction. Any form of tool wear alters the contact surfaces between the tool and workpiece, inuencing and increasing the AE signal in- tensity. Hence, advanced warning of potential tool breakage sometimes results in the appearance of micro-ssures in the tool, which cause an escalation of the AE signals – allowing a basic form of tool and process monitoring to be achieved. AE generation in metallic machining operations, can extend over frequencies of several MHz, although the signal intensity is normally very low and diminishes with increasing distance from its source. Any form of machine vibrations and inter- ferences from the local environment introduce signals from a low frequency range, meaning that any form of signicant analysis is normally only possible above 50 kHz. Machine tool interference sources are usually the result of either electrical, or hydraulic main and feed drives, as well as from bearing noise, spindles and gears. ese unwanted interferences can be suppressed by utilising suitable high-pass lters, or alterna- tively a well-designed AE sensor(s), with inherent high-pass frequency characteristics. 77 ‘Root mean square’ (rms), is a measure of the eective mean current of an alternating current. Its actual rms value is de- rived from the power dissipation by an ac current. nals respectively, on the ‘over-turning’ of transversal holes present in the external turning of the workpiece. Hence, the interrupted cut can clearly be seen peri- odically in the resultant force traces. In Fig. 179biii, the AE rms signal shows this interference, albeit not very well pronounced, unlike that of the force trace produced in Fig. 179biv, where a denite noise spike can be seen. is combination of two complementary sensing elements and their sensor signals, allows the reliable detection of a process fault, such as tool break- age detection. Until approximately the mid-1990’s, commercial versions of cutting force monitoring equipment for the measurement of a rotating cutting tool, or an edge was not readily available for: drilling, reaming, tapping and milling applications. A major advantage of these rotating cutting force dynamometers, is that they can be used for multi-axes contour milling applications, or simply for an investigation of a discrete tool’s cut- ting edge geometry and its anticipated machining per- formance. An early version of such a rotating cutting force dynamometer, is depicted in Fig. 180a. In Fig. 180b, graphs have been produced showing the cutting force and torque results respectively, pro- duced by the rotating cutting force dynamometer. In- terest frequently centres on the forces and moments acting on the rotating tool. A rotating cutting force dynamometer (Fig. 180a), allows measurement of three orthogonal forces: ‘F X ’ , ‘F Y ’ , ‘F Z ’ , together with the moment ‘M Z ’. e data measured by the rotating dynamometer occurs via miniature charge ampliers, which are then transferred by telemetry to an appro- priately positioned stationary antenna. e telemetry involves a bi-directional transmission, with measured data being transmitted to the ‘stationary side’ of the monitoring system and any control commands for the integral charge ampliers transmitted to the appropri- ate section of the rotating dynamometer. e power supply to the electronics in the rotor, occurs by the same antenna, but having a dierent carrier frequency to that of the data transfer. Typical resultant signals produced by the rotating dynamometer are shown in Fig. 180b and have been ‘zoomed’ for the investigation of a single drill’s cutting edge. Cutting force dynamometers of various congu- rations, are invaluable tools for any form of in-depth machinability study, as they indicate the precise condi- tions at the cutting tool’s edge(s), in a truly dynamic situation. All dynamometers that are purchased from the manufacturer must come with an appropriate cali- Machinability and Surface Integrity Figure 179. A Rotating Cutting-force Dynamometer (RCD), utilising piezoelectric sensor systems. [Courtesy of Kistler Instrumente AG] . Chapter . scrapping this component! Machinability and Surface Integrity Tool-life Diagrams Machinability is a subject that has yet to be fully-de- ned and analysed, in particular the interactive. types of machinability investigation. In the former case, it is Machinability and Surface Integrity Figure 1 78. Multi-axis non-rotating dynamometers, used for: milling, drilling and turning. machining costs, together with their resultant productivity. [Courtesy of Sandvik Coromant] . Machinability and Surface Integrity sensors, such as: force transducers, accelerometers can be