Machining of High Strength Steels With Emphasis on Surface Integrity by air force machinability data center_9 docx

neering work, then further rened 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 prole and its as- sociated inclination angle will vary depending upon the cutting speed selected, with individual cutting speeds having specic 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 specic amount of ank wear ‘V B ’ is established. 70 is ‘general case’ Taylor formula, has been expanded and developed which denes the machining characteristics with more mathematical rigour, including the eects 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 eect of cutting speed on tool life, since the exponent of ‘V’ is 1.0. Moreover, aer 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 diculties 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 dierent 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 modied 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 oen 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-eciency 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 signicant 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 ecient 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 supercial 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 certicated 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 deection 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 eects 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 eect 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 eect 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 eect’ , 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, oer 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 denite 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 innitely high resistance, together with ampliers that are perfectly free from any form of leakage and ‘non-operate cur- rents’ , with an amplication factor of innity! e dri in today’s charge ampliers 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 aer 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, oering 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 oer 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 aer, 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 specic 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 dierent dynamometer conguration, 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, inuencing 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 signicant 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 eective 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 denite 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 ampliers, 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 ampliers 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 dierent 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 congu- 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  Figure 180. A Rotating Cutting-force Dynamometer (RCD), utilising piezoelectric sensor systems. [Courtesy of Kistler Instrumente AG] . Machinability and Surface Integrity  . 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. 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. 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