# Analytical models for high performance milling. Part I: Cutting forces, structural deformations and tolerance integrity

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2 ATICLE IN PESS E. Budak / International Journal of Machine Tools & Manufacture 46 (6) In this paper milling force structural deformation form error prediction and control models are presented. Experimental results are also given to demonstrate the applications of these models. The models can be used to check the constraints such as available machine power or allowable form errors and determine the high performance milling conditions.. Milling process geometry and force modeling Milling forces can be modeled for given cutter geometry cutting conditions and work material. Two different methods will be presented for the force analysis: Mechanistic and mechanics of cutting models which differ in the way the cutting force coefficients which relate the cut chip area to the fundamental milling force components shown in Fig. are determined... Mechanistic force model In mechanistic force model cutting force coefficients are calibrated for certain cutting conditions using experimental data. However the same milling force model can be used for both mechanistic and mechanics of cutting models. Consider the cross-sectional view of a milling process shown in Fig.. For a point on the (j th ) cutting tooth differential milling forces corresponding to an infinitesimal element thickness (dz) in the tangential df t radial df r and axial df a directions can be given as df tj ðf; zþ ¼K t h j ðf; zþ dz df rj ðf; zþ ¼K r df tj ðf; zþ df aj ðf; zþ ¼K a df tj ðf; zþ ðþ where f is the immersion angle measured from the positive y-axis as shown in Fig.. The axial force component F a is in the axial direction of the cutting tool which is perpendicular to the cross-section shown in Fig.. In Eq. () the edge forces are also included in the cutting force coefficient which is usually referred to as the exponential force model. They are separated from the cutting force coefficients in edge force or linear-edge force y x w φ j df aj df rj df tj Fig.. Cross-sectional view of an end mill showing differential forces. φ p model [4]: df tj ðf; zþ ¼½K te þ K tc h j ðf; zþš dz df rj ðf; zþ ¼½K re þ K rc h j ðf; zþš dz df aj ðf; zþ ¼½K ae þ K ac h j ðf; zþš dz where subscripts (e) and (c) represent edge force and cutting force coefficients respectively. The radial (w) and axial depth of cut (a) number of teeth (N) cutter radius () and helix angle (b) determine what portion of a tooth is in contact with the work piece for a given angular orientation of the cutter f ¼ Ot where t is the time O is the angular speed in (rad/s) or O ¼ pn=6 n being the (rpm) of the spindle. The chip thickness at a certain location on the cutting edge can be approximated as follows: h j ðf; zþ ¼f t sin f j ðzþ (3) where f t is the feed per tooth and f j (z) is the immersion angle for the flute (j) at axial position z. Due to the helical flute the immersion angle changes along the axial direction as follows: f j ðzþ ¼f þðj Þf p tan b z (4) where the pitch angle is defined as f p ¼ p=n. The tangential radial and axial forces given by Eqs. () and () can be resolved in the feed x normal y and the axial direction z and can be integrated within the immersed part of the tool to obtain the total milling forces applied on each tooth. For the exponential force model the following is obtained after the integration: F xj ðfþ ¼ K tf t 4 tan b ½ cos f j þ K r ðf j ðzþ sin f j ðzþþš z juðfþ z jl ðfþ F yj ðfþ ¼ K tf t 4 tan b ½ðf jðzþ sin f j ðzþþ þ K r cos f j ðzþš z juðfþ z jl ðfþ F zj ðfþ ¼ K ak t f t tan b ½cos f jðzþš z juðfþ z jl ðfþ ð5þ where z jl (f) and z ju (f) are the lower and upper axial engagement limits of the in cut portion of the flute j. The engagement limits depend on the cutting and the tool geometries f st ðzþ ¼p cos w ðdown millingþ f ex ðzþ ¼cos w ðup millingþ. ð6þ Note that f ex is always p in down milling and f st is always in up milling according to the convention used in Fig.. The helical cutting edges of the tool can intersect this area in different ways resulting in different integration limits which are given in [45]. The total milling forces ðþ

3 48 ATICLE IN PESS E. Budak / International Journal of Machine Tools & Manufacture 46 (6) can then be determined as F x ðfþ ¼ XN j¼ F y ðfþ ¼ XN j¼ F z ðfþ ¼ XN j¼ F xj ðfþ; F yj ðfþ F zj ðfþ. The cutting torque and power due to the tooth j can easily be determined from Eqs. () and (3) as follows: T j ðfþ ¼ K tf t tan b ½cos fšz juðfþ z jl ðfþ P j ðfþ ¼OT j ðfþ. ð8þ The total torque and power due to all cutting teeth can be determined similar to the summation for the forces given Eq. (7). Maximum value of the forces torque and power can be determined after one full revolution of the tool i.e. f: p is simulated. For the linear-edge force model the forces are obtained similarly by using Eq. () and integrating within the engagement limits as follows []: F xj ðfþ ¼ tan b K te sin f j ðzþ K re cos f j ðzþþ f t 4 F xj ðfþ ¼ ½K rc ðf j ðzþ sin f j ðzþþ K tc cos f j ðzþš tan b K re sin f j ðzþ K te cos f j ðzþþ f t 4 ½K tc ðf j ðzþ sin f j ðzþþ K rc cos f j ðzþš F xj ðfþ ¼ tan b ½K aef j ðzþ f t K ac cos f j ðzþš z ju z jl ð9þ The forces given by Eqs. (5) and (8) can be used to predict the cutting forces for a given milling process if the milling force coefficients are known. As mentioned in the beginning of this section in the mechanistic models the force coefficients are calibrated experimentally which is explained in the following section... Identification of milling force coefficients In mechanistic force model milling force coefficients K t K r and K a can be determined from the average force expressions [] as follows: K r ¼ P F y Q F x P F x þ Q F y F x K t ¼ f t ðp QK r Þ K a ¼ F z f t K t T zju z jl zju z jl ð7þ ðþ where P ¼ an p ½cos fšf ex f st Q ¼ an p ½f sin fšf ex f st T ¼ an p ½cos fšf ex f st. ðþ The average forces F x ; F y and F z can be obtained experimentally from milling tests. In exponential force model the chip thickness affects the force coefficients. Since the chip thickness varies continuously in milling the average chip thickness h a is used: h a ¼ f t cos f st cos f st f ex f st. () In calibration tests the usual practice is to conduct experiments at different radial depths and feed rates in order to cover a wide range of h a for a certain tool material pair. The force coefficients can then be expressed as following exponential functions: K t ¼ K T h p a K r ¼ K h q a K a ¼ K A h s a ð3þ where K T K K A pqand s are determined from the linear regressions performed on the logarithmic variations of K t K r K a with h a. In linear-edge force model the total cutting forces are separated into two parts: edge forces and cutting forces. The edge force represents the parasitic part of the forces which are not due to cutting and thus do not depend on the uncut chip thickness whereas cutting forces do. Then the average forces can be described as follows: F q ¼ F qe þ f t F qc ðq ¼ x; y; zþ (4) where the edge and cutting components of the average forces ð F qe ; F qc Þ are determined using the linear regression on the average measured milling forces. The milling force coefficients for the linear-edge force model can be obtained from the average forces similar to the exponential force model as follows: K tc ¼ 4 F xc P þ F yc Q P þ Q ; K rc ¼ K tcp 4 F xc ; K ac ¼ F zc Q T xe S þ F ye T K te ¼ F S þ T ; K re ¼ K tes þ F xe T K ae ¼ p F ze ð5þ an f ex f st where P Q and T are given by Eq. () and S ¼ an p ½sin fšf ex f st.

4 ATICLE IN PESS E. Budak / International Journal of Machine Tools & Manufacture 46 (6) Mechanics of milling force model The mechanistic force models introduced in the previous section yield high accuracy force predictions for most applications. However since the cutting force coefficients must be calibrated for each tool material pair covering the conditions that are of interest this approach may sometimes be very time consuming. In this sense mechanics of milling approach is more general and may reduce the number of tests significantly. The basic idea in this approach is to use analytical cutting models relating the chip area to the cutting forces and to determine the parameters required in the model experimentally when necessary. In case of milling an oblique cutting model has to be employed due to helical flutes. In oblique cutting models there are several important planes which are used to measure tool angles and write down velocity and force equilibrium relations [6]. The normal plane which is perpendicular to the cutting edge is commonly used in the analysis. After several assumptions and velocity and the force equilibrium equations the following expressions are obtained for the cutting force coefficients in an oblique cutting process: K tc ¼ t cosðg n a n Þþtan Z c sin g n tan b sin f n c t sinðg K rc ¼ n a n Þ sin f n cos b c K ac ¼ t cosðg n a n Þ tan b tan Z c sin g n ð6þ sin f n c qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi where c ¼ cos ðf n þ g n a n Þþtan Z c sin g n. In Eq. (6) (t) is the shear stress in the shear plane f n is the shear angle in the normal plane b is the angle of obliquity or helix angle and Z c is the chip flow angle measured on the rake face. The chip flow angle can be solved iteratively based on the equations obtained from force and velocity relations [94]. However for simplicity Stable s rule [7] may also be used which states that Z c Eb. g n and a n are the friction angle and the rake angle in the normal plane respectively and are given by [6] tan g n ¼ tan g cos Z c ; tan a n ¼ tan a r cos b (7) where a r is the rake angle measured in the velocity plane which is normal to the tool axis and g is the friction angle on the rake face. The procedure proposed by Armarego and Whitfield [9] and later by Budak et al. [] for the prediction of milling force coefficients will be briefly described here. First of all the required data is obtained from the orthogonal cutting tests in order to reduce the number of variables thus the number of tests and to generate a more general database which can be used for other processes as well. The shear angle shear stress and friction coefficient can be obtained from orthogonal cutting tests as follows [9]: tan f ¼ r cos a r sin a t ¼ ðf p cos f F q sin fþ sin f bt tan g ¼ F q þ F p tan a F p F q tan a ð8þ where r is the cutting ratio or the ratio of the uncut chip thickness to the chip thickness a is the rake angle F p and F q are the cutting forces in the cutting speed and the feed direction respectively. If the linear-edge force model is to be used then the edge cutting force components must be subtracted from the cutting forces measured in each direction using linear regression [4]. The edge force coefficients are identified from the edge cutting forces. After the orthogonal cutting tests are repeated for a range of cutting speed rake angle and uncut chip thickness an orthogonal cutting database is generated for a certain tool and work material pair. These data can then be used to determine the milling cutting force coefficients using the oblique model given by Eq. (6). The force coefficients and the milling forces predicted using this approach have been demonstrated to be very close to the milling experiment results [4]..4. Example application As a demonstration of the force models presented here a titanium (Ti6Al4V) milling example is considered. First of all an orthogonal cutting database is generated using carbide tools with different rake angles at different speeds and feedrates [4]: t ¼ 63 MPa; b ¼ 9: þ :9a r ¼ r h a ; r ¼ :755 :8a; a ¼ :33 :8a K te ¼ 4 N=mm; K re ¼ 43 N=mm: In addition many test were conducted to calibrate the milling force coefficients directly from the milling tests. Budak et al. [] and Budak [4] showed that there is in general a good agreement between the predicted and the identified cutting force coefficients using this approach. As an example the predictions for one of the cases are shown in Fig. together with the measured milling forces for one full rotation of the cutter. This is a half-immersion up milling test performed using a 3 helix 9.5 mm diameter and four-fluted end mill with rake angle. The axial depth of cut is 5 mm and.5 mm/tooth feed was used at 3 m/min cutting speed. The measured and the predicted cutting forces using the force coefficients identified from the milling tests and calculated using the oblique model in all three directions are shown in the figure. As it can be seen from this figure the predictions are very close to the measured forces. The models were tested for many cases and good results were obtained [].

5 48 ATICLE IN PESS E. Budak / International Journal of Machine Tools & Manufacture 46 (6) Force (N) Experiment Fy Simulation (Calibrated Fz Simulation (Predicted y p (xz) surface generation points - -4 Fx otation Angle (deg) Fig.. Predicted and the measured milling forces for the example in Section.4. z y x δ y (z) Fig. 3. Surface generation in peripheral milling. 3. Deflections and form errors 3.. Surface generation In peripheral milling the work piece surface is generated as the cutting teeth intersect the finish surface. These points are called the surface generation points as shown in Fig. 3. As the cutter rotates these points move along the axial direction due to the helical flutes completing the surface profile at a certain feed position along the x-axis. The surface generation points z cj corresponding to a certain angular orientation of the cutter f can be determined from the following relation: ( f j ðz cj Þ¼f þ jf p tan b ¼ for up milling; (9) p for down milling: The surface generation points can then be resolved from above equation as follows: z cj ðfþ ¼ ðf þ jf pþ for up milling; tan b z cj ðfþ ¼ ðf þ jf p pþ for down milling: tan b ðþ As the surface is generated point by point by different teeth resulting in helix marks on the surface in helical end milling the surface finish is not as good as the finish that would be obtained by a zero-helix tool. In case of nonhelical end milling the whole surface profile at a certain feed location is generated by a single tooth as the immersion angle does not vary along the axial direction. Thus the helix marks do not exits with zero-helix end mills and a better surface finish is obtained. However helical flutes result in much smaller force fluctuations lower peak forces and thus smoother cutting action with reduced impacts. In addition helical flutes improve chip evacuation. The deflections of the tool and the work piece in the normal direction to the finish surface are imprinted on the surface resulting in form errors which are analyzed the next. 3.. Form errors in peripheral milling The form error can be defined as the deviation of a surface from its intended or nominal position. In case of peripheral milling the deflections of the tool and the part in the direction normal to the finished surface cause the form errors as shown in Fig. 3. Then the total form error at a certain position on the surface e(xz) can be written as follows: eðx; zþ ¼d y ðzþ y p ðx; zþ () where d y (z) is the tool deflection at an axial position z and y p (xz) is the work deflection at the position (xz) Structural model of the tool Several modeling approached can be used to determine the end mill deflections which will be summarized here Cantilever beam model. The end mill can be modeled as a beam with clamping stiffness as shown in Fig. 4. k x and k y represent the linear and torsional clamping stiffness at the holder tool interface. They can be identified experimentally for a certain tool holder pair [88]. The cutting tool is divided into n elements along the axial direction. The normal force in the mth element f ym can be written as f ym ðfþ ¼ K tf t X N h ðf 4 tan b j ðzþ sin f j ðzþþ j¼ þk r cos f j ðzþ i zm z m ðþ where z m represents the axis boundary of the cutter in element m shown in Fig. 4. The elemental cutting forces are equally split by the nodes m and (m ) bounding the tool element (m ). The deflection at a node k caused by the

6 ATICLE IN PESS E. Budak / International Journal of Machine Tools & Manufacture 46 (6) force applied at the node m is given by the cantilever beam formulation as [84]: d y ðk; mþ ¼ f ymz m 6EI ð3u m u k Þ þ f ym k x þ f ymu m u k k y for ou k ou m d y ðk; mþ ¼ f ymu m 6EI ð3u k u m Þþ f ym k x þ f ymu m u k k y for u m ou k ð3þ where E is the Young s modulus I is the area moment of inertia of the tool u k ¼ L z k L being the gauge length of the cutter. The total static defection at the nodal station k can be calculated by the superposition of the deflections produced by all (n+) nodal forces: d y ðkþ ¼ Xnþ d y ðk; mþ. (4) m¼ k x k θ m Fig. 4. Cross-sectional view of an end mill showing differential forces. The tool deflections at the surface generation points can be determined from Eq. (6) and substituted into Eq. (3) to determine the form errors Segmented beam model. The area moment of inertia must take the affect of the flutes into account. Use of an equivalent tool radius e ¼ s where s ¼.8 for common end mill geometries was demonstrated to yield reasonably accurate predictions by Kops and Vo [9]. An improved method of tool compliance modeling is given by Kivanc and Budak [8] where end mill deflections were approximated using a segmented beam model. For such a case if a load is applied at the tip of the tool the maximum deflection is given by [] y max ¼ FL3 3EI þ FLðL LÞðL þ LÞ 6 EI þ FLðL LÞðL þ LÞ ð5þ 6 EI where D is the mill diameter D is the shank diameter L is the flute length L is the overall length F is the point 3 load Iand I are the moment of inertias of the fluted and unfluted parts respectively. In case of distributed forces and existence of clamping stiffness a formulation similar to Eq. (3) can be derived. Due the complexity of the cutter cross-section along its axis the inertia calculation is the most difficult aspect of the static analysis. The crosssections of some end mills are as shown in Fig. 5. In order to determine the inertia of the whole crosssection inertia of region is first derived and inertia of the other regions are obtained by transformation []. The total inertia of the cross-section is then obtained by summing the inertia of all regions. The inertia of region is derived by computing equivalent radius eq in terms of the radius r of the arc and position of the center of the arc (a) []: qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi eq4-fluteðyþ ¼a sinðyþþ ðr a Þþa sin ðyþ; oypp= eq3-fluteðyþ ¼a cos y þ p 3 rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi þ ðr a Þþa cos ðy þ p 3 Þ ; oypp=3 p eq-fluteðyþ ¼ a cosðyþþ ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ðr a Þþa cos ðyþ; oypp: ð6þ The moment of inertia of region of a four-flute end mill about x- and y-axis can be written as " Z p= Z # eq4-fluteðyþ I xx4-flute ¼ r 3 sin ðyþ dr dy 4 fd 3 Fig. 5. Cross-sections of 4-3- and -flute end mills. " 8 p fd 4 þ pðfd=þ r þ a fd # " Z p= Z # " eq4-fluteðyþ I yy4-flute ¼ r 3 cos ðyþ dr dy 8 p fd # 4 ð7þ where orp eq (y). The same formulation can be written for region of the 3- and -flute tool. After transforming the inertia of region the total inertias are found as follows: I xx4-flute;tot ¼ I yy4-flute;tot ¼ ði xx4-flute þ I yy4-fluteþ I xx3-flute;tot ¼ I yy3-flute;tot ¼ :5ðI xx4-flute þ I yy4-fluteþ I xx-flute;tot ¼ ði yy-fluteþ; I yy-flute;tot ¼ ði yy-fluteþ ð8þ Finite elements modeling. In order to verify and improve the accuracy of analytical model predictions finite elements analysis FEA is also used for tool deflections. fd fd

7 484 ATICLE IN PESS E. Budak / International Journal of Machine Tools & Manufacture 46 (6) Approximately sixty tools with different configurations were simulated. Although FEA can be very accurate it can also be very time consuming for each tool configuration in a virtual machining environment. Therefore simplified equations were also derived to predict deflections of tools for given geometric parameters and density: deflection max ¼ C F L 3 E D 4 þ ðl3 L 3 N Þ D 4 (9) where F is the applied force and E is the modulus of elasticity (MPa) of the tool material. The geometric properties of the end mill are in millimeters. The constant C is and 7.93 and N is and.974 for 4-3- and -Flute cutters respectively Clamping stiffness After the tool is modeled accurately the clamping stiffness must also be known for the total tool deflection. Depending on the tool and clamping conditions the contribution of the clamping flexibility to the total deflection of the tool can be significant. There is not a model available in the literature for modeling of the tool clamping stiffness. However a model given by ivin [] for the stiffness of cylindrical connections can be utilized. According to this model the initial interference-fit pressures in the connection create a pre-loaded system which is generally shaped by the applied clamping torque. Since the contact area is a function of tool diameter and contact length the magnitudes of the interference displacements can be related to them. The elastic displacement in the connection can be determined as [] d ¼ cq pd (3) where c is the contact compliance coefficient q ¼ F=L is the force per contact length and d is the tool diameter. Therefore the clamping stiffness can be expressed as k ¼ F m =n d ¼ pld c (3) where n is a constant used to compensate the effect of the small lengths of contact. If Lo3 then this constant should be equal to 3 otherwise n ¼. The effect of using different materials for tools is represented by m. For carbide it is taken as and for HSS tools it is.9. Coefficient c should be experimentally determined for a connection. In other words c is the ratio of the deflection representation to the force representation and it is constant for a type of tool holder: c ¼ d rep (3) F rep where d rep ¼ dpd F rep ¼ F m nl. ð33þ After many static deflection tests with different clamping conditions c was determined to be approximately.7 for holders without collets such as power chucks and shrink fit holders whereas for collet type holders c changes substantially (.5.5) depending on the type of the collet Structural model of the work piece Work pieces deflect under cutting forces contributing to form errors. In general the finite-element method (FEM) can be used to determine the structural deformations of the work piece. The elemental cutting forces in the normal y- direction given by Eq. () are to be used as the force vector. For the cases where the part is very thin such as a turbine blade or a thin plate the change in the structural properties of the work piece due to removed material can be very important for accurate prediction of the deflections [84]. In addition the tool work contact and thus the force application points change as the tool moves along the feed direction. Therefore the form error due to work piece deflections in milling require that the FE solutions be repeated many times in order to consider these special effects i.e. varying part thickness and force location. Fig. 6 shows the work piece model which is used by Budak and Altintas [8] and Budak [4] for deflection calculations. The part thickness is reduced from t u to t c at the cutting zone where the cutter enters the part at point B and exits at point A in down milling mode. The nodal forces on the tool are applied in the opposite direction on the corresponding nodes in the cutting zone. For a down milling case the cutting teeth on an end mill with a positive helix angle enters the cut at the bottom of the part where it is the most rigid. As the tool rotates the contact points move along the axial direction where tool deflections are much smaller. For a plate-like structure however these are the most flexible sections of the part resulting in high work piece deflections. Therefore depending on the application both part and tool deflections can be very significant and must be included in the calculations. The form error t u cutter entry B B b n K A A t c cutter exit Fig. 6. Work piece model used in deflection and form error analysis.

8 ATICLE IN PESS E. Budak / International Journal of Machine Tools & Manufacture 46 (6) calculations at a certain location of the tool result in the surface profile at that position. After repeating this at many locations along the feed direction the complete form error map of the surface is obtained Structure process interaction Part and tool deflections affect the cutting process in two ways. The deflections in the chip thickness direction will change the actual value of the chip thickness. However due to the continuous feed motion the chip thickness will approach to its intended value in a few tooth periods [3]. For most of the finishing operations the radial depth of cut is very small in order to reduce forces deflections and thus the form errors. The deflections will change the radial depth of cut and thus the start and exit immersions angles f st and f ex as follows: f st ðzþ ¼p cos w fðzþ down milling; f ex ðzþ ¼cos w fðzþ up milling: ð34þ The effective start and exit angles vary along the axial z- direction as the actual value of the desired width of cut w changes due to deflections: w f ðzþ ¼w þ d y ðzþ y p ðx; zþ. (35) The effectives start and exit angles given by Eq. (34) must be used in the force and form error predictions for accurate results when the deflections are comparable to the radial depth of cut. The force model which includes these effects was named as flexible force model by Budak and Altintas [8] and Budak [4]. The solution is an iterative one as the deflections and forces depend on each other Example application Peripheral milling of a cantilever plate made out of titanium (Ti6Al4) is considered [84]. The plate is very flexible with dimensions of mm. Its flexibility is further increased by reducing the thickness by.65 mm down to.8 mm in a single milling pass i.e. the axial depth of cut is 34 mm. A 9 mm carbide end mill with 3 helix and single flute was used in order to eliminate run out effects. The tool gauge length is 55.6 mm and the linear clamping stiffness was measured to be 9.8 kn/mm where the torsional clamping stiffness was negligible. In order to constrain deflections and eliminate work tool separation a very small feed rate of f t ¼.8 mm/tooth was used. The cutting force coefficients were identified from the milling tests using the exponential force model: K T ¼ 7ðMPaÞ; p ¼ :67; K ¼ :39; q ¼ :43. The Young s Modulus of the tool and work materials are 6 and GPa respectively. Experimentally measured and simulated form errors on the plate are shown in Fig. 7. Only the flexible force model predictions are shown as the deflections are very high compared to the radial depth. The form error at the cantilevered edge of the plate is due to the tool deflection and it is approximately 3 mm. The maximum form error occurs close to the free end of the plate where it is most flexible. The error increases from the start of the milling at the left side to the end due to the reduced thickness and increased flexibility of the plate. The rigid model where the deflection process interaction is neglected overestimates the form errors significantly about 5 mm at the maximum error location [8]. The flexible model predictions agree with the measurements as shown in the figure. 4. Control and minimization of form errors 4.. Feedrate scheduling In machining operations the usual practice is to use a constant feed rate for a machining cycle. In cases where there are tolerance violations due to excessive deflections the feed may have to be reduced in the whole cycle even the maximum form error occurs only at a specific location. Surface Error (mic) Axial Depth -z (mm) Food Direction (mm) Surface Error (mic) Axial Depth -z (mm) Food Direction (mm) Fig. 7. Measured and simulated form errors in the peripheral milling of a cantilever plate.

10 ATICLE IN PESS E. Budak / International Journal of Machine Tools & Manufacture 46 (6) st (mm/rev-tooth) Feed Direction -x (mm) 6 Surface Error (mic) Axial Depth -z (mm) Food Direction -x (mm) Fig. 8. Form errors measured on the plate after it is machined with the scheduled feeds shown. Max. Surface Error (mic) affecting the form errors which further increases the M significantly. 5. Conclusions Simulation ft=. mm/tooth ft=.5 ft=. Experiment adial Depth of Cut (mm) Fig. 9. Measured and predicted maximum form errors for different radial depths and feed rates. In this paper analytical milling force part and tool deflection and form error models are presented and their application in improving the performance of the process is demonstrated. The milling process is considered due to its complex geometry and mechanics however similar modeling methodology can be applied to other machining processes such as turning. On the other hand the models can be extended to more complex milling processes such as ball end and five-axis milling. These models provide general information about the relations between the process performance and the process parameters. In addition they can be used to simulate real cases and the best set of parameters to improve the process performance can be selected. It should be noted that the process optimization can only be done on a stable process as demonstrated in [4]. Thus the chatter suppression methods are presented in the second part of the paper [5] must be applied first. For very practical and fast implementation of these models in industry however interfaces between CAD/CAM systems and/or CNC codes have to be developed first. eferences [] M.E. Merchant Basic mechanics of the metal cutting process ASME Journal of Applied Mechanics 66 (944) [] M.E. Martelotti An analysis of the milling process Transaction of the ASME 63 (94) [3] F. Koenigsberger A.J.P. Sabberwal An investigation into the cutting force pulsations during milling operations International Journal of Machine Tool Design and esearch (96) [4] W.A. Kline.E. DeVor I.A. Shareef The prediction of surface accuracy in end milling Transactions of the ASME Journal of Engineering for Industry 4 (3) (98) [5] W.A. Kline.E. DeVor I.A. Shareef The effect of run out on cutting geometry and forces in end milling International Journal of Machine Tool Design and esearch 3 (983) 3 4. [6] J.W. Sutherland.E. DeVor An improved method for cutting force and surface error prediction in flexible end milling systems Transactions of the ASME Journal of Engineering for Industry 8 (986) [7] E. Budak Y. Altintas Peripheral milling conditions for improved dimensional accuracy International Journal of Machine Tool Design and esearch 34/7 (994) [8] E. Budak Y. Altintas Modeling and avoidance of static deformations in peripheral milling of plates International Journal of Machine Tool Design and esearch 35 (3) (995) [9] E.J.A. Armarego.C. Whitfield Computer based modeling of popular machining operations for force and power predictions Annals of the CIP 34 (985) [] E. Budak Y. Altintas E.J.A. Armarego Prediction of milling force coefficients from orthogonal cutting data Transactions of the ASME Journal of Manufacturing Science and Engineering 8 (996) 6 4. [] Y. Altintas P. Lee A general mechanics and dynamics model for helical end mills Annals of the CIP 45 (996) [] Y. Altintas S. Engin Generalized modeling of mechanics and dynamics of milling cutters Annals of the CIP 5 () 5 3. [3] E. Ozturk E. Budak Modeling of 5-axis milling forces in: Proceedings of the Eighth CIP International Workshop on Modeling of Machining Operations Chemnitz May 5 pp

11 488 ATICLE IN PESS E. Budak / International Journal of Machine Tools & Manufacture 46 (6) [4] E. Budak The mechanics and dynamics of milling thin-walled structures Ph.D. Dissertation University of British Columbia 994. [5] Y. Altintas Manufacturing Automation Cambridge University Press Cambridge. [6] E.J.A. Armarego.H. Brown The Machining of Metals Prentice- Hall Englewood Cliffs NJ 969. [7] G.V. Stabler Fundamental geometry of cutting tools Proceedings of the Institution of Mechanical Engineers (95) 4 6. [8] E. Kivanc E. Budak Structural modeling of end mills for form error and stability analysis International Journal of Machine Tools and Manufacture 44 () (4) 5 6. [9] L. Kops D.T. Vo Determination of the equivalent diameter of an end mill based on its compliance Annals of the CIP 39 (99) [] F. Beer E. Johnston Mechanics of Materials McGraw-Hill U.K 99. [] J.A. Nemes S. Asamoah-Attiah E. Budak Cutting load capacity of end mills with complex geometry Annals of the CIP 5 () [] E. ivin Stiffness and Damping in Mechanical Design Marcel Dekker New York 999. [3] E. Budak Y. Altintas Flexible milling force model for improved surface error predictions in: Proceedings of the ASME 99 European Joint Conference on Engineering Systems Design and Analysis Istanbul Turkey ASME PD pp [4] E. Budak Improvement of productivity and part quality in milling of titanium based impellers by chatter suppression and force control Annals of the CIP 49 () () [5] E. Budak Analytical models for high performance milling Part II: process dynamics and stability International Journal of Machine Tools and Manufacture in press.

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