Keynote Papers

Keynote Papers Presented at the ScientificTechnical Committees

Paper Discussion Sessions

Cutting of Hardened Steel

H. K.Tonshoff ( l ) , C. Arendt, R. Ben Amor Institute for Production Engineering and MachineTools, University of Hannover, Germany

Abstract Cutting of hardened steel is a topic of high interest for toda 's industrial production and scientific research. Machine parts consisting of hardened steel are high peiormance components which are often loaded near their physical limits. The functional behavior of machined parts is decisively influenced by the fine finishing process which represents the last step in the process chain and can as well be undertaken by cutting as grinding. An overview of the mechanisms of chip removal in hard cutting and the thermo- mechanical influence of the work area is presented. Furthermore, several models of chip removal in hard turning are introduced and discussed summarizing the metallurgical fundamentals and giving an overview on stress and temperature distributions in the work area. Boundary conditions for hard cutting as e.g. machine tools, cutting materials and others are subject to discussion to determine the achievable workpiece quality and economic efficiency of hard cutting processes in comparison with grinding. Keywords: Hard cutting, chip removal, machine, cutting tool, quality of work

1 INTRODUCTION Cutting of hardened steel is a topic of high interest for today's industrial production and scientific research. Tests and first introductory steps have been made in automotive-, gear-, bearing-, tool and die making industry. In research institutes and universities, basic investigations experimentally as well as theoretically have been made since more than 25 years [l]. In the ClRP community, the subject has been worked out by several researchers. For this review many colleagues contributed b discussion and publications. The authors thank the f&owing contributors for their valuable assistance: - M. Balazinski, Montrkal, Kanada - E. Brinksmeier, Bremen, Germany - G. Byrne, Dublin, Ireland - T. H. C. Childs, Leeds, UK - M. A. Davies, Gaithersburg, USA - M. A. Elbestawi, Hamilton, Canada - C. J. Evans, Gaithersburg, USA - F. Klocke, Aachen, Germany - W. Koenig, Aachen, Germany - R. Komanduri, Stillwater, USA - C. R. Liu, W. Lafayette, USA - A. Moisan, Aix-en-Provence, France - C. A. van Luttervelt, Delft, Holland - K. Nakayama, Tokyo, Japan t - M. C. Shaw, Tempe, USA - H. Schulz, Darmstadt, Germany - K. Weinert, Dortmund, Germany - R. Wertheim, Tefen, Israel

Machine parts consisting of hardened steel are high performance components which are often loaded near their physical limits. The functional behavior of machined parts is decisively influenced by the fine finishing process which represents the last step in the process chain and can as well be undertaken by cutting as grinding. For this reason, fine finishing is defined as an important process and its results have to satisfy high quality requirements. The product specific issues and demands also meet general trends in cutting such as flexibility, ecology, cost effectiveness, time to market and process agility.

Developments in machine tools as well as in process technology focus on cutting hardened steel and rapidly lead to a highly raised industrial relevance of hard cutting. In fact, hard cutting can seriously be regarded as an alternative for grinding operations under certain circumstances. A comparison between hard cutting and grinding is, due to variability of workpiece geometry and required quality, difficult. Koch specified economy, flexibility, ecology and quality as important criteria for process evaluation [2]. High flexibility and the ability to manufacture complex workpiece geometry in one set-up represent the main advantages of hard cutting in comparison to grinding. Furthermore, the substitution of grinding processes with cutting processes enables to avoid coolants and there- fore can actually be regarded as interestin alternative

comparison of the processes hard cutting and grinding is given.

even from the ecological point of view. In a gure 1.1, a

kosnd. 0 uuilN0-

l ? i z 3 m m n

5WZ474tsO Vw

Figure 1.1 : Comparison of hard cutting with grinding In the following, several aspects of the cutting of hardened steel are examined and discussed. Metallur- gical fundamentals of hardened ferrous materials serve to characterize the workpiece material concerning required process quality as well as quality of work. Furthermore, an overview of the mechanisms of chip removal in hard cutting and the thermo-mechanical

Annals of the ClRP Vol. 49/2/2000 547

influence of the work area is presented. Moreover, several models of chip removal in hard turning are introduced and discussed summarizing the metallurgical fundamentals and giving an overview on stress and temperature distributions in the work area, too.

Boundary conditions for hard cutting as e.g. machine tools, cutting materials and others are subject to discussion to determine the achievable workpiece quality and economic efficiency of hard cutting processes in comparison with grinding.

2 M ETALLU RGlCAL FUN DAM ENTALS Mechanical properties of hardened or case hardened steel can be adjusted in a wide range and consequently influence the cutting process. Depending on alloys and the heat treatment, the hardness of hardened ferrous materials can be varied between 50 and 70HRC (figure 2.1). The relatively high hardness of ferrous work materials can be reached by martensitic transformation and/or carbide precipitation.

Figure 2.1 : Hardness and microstructure of hardened ferrous work materials

2.1 Physico-chemlcal fundamentals In contrast to annealing in which an equilibrium state of texture is adjusted, hardening serves to initiate an imbalance status. Depending on cooling rate and consequently on carbon diffusion deceleration, pearlite, bainite and/or martensite is formed. The time- temperature-transformation curve (TIT curve) indicates the composition of the texture. Figure 2.2 shows an example of the carbonized case hardening steel 16MnCr5 (comparable to AlSl5115) which has been achieved by continuous cooling.

C S I M P S A l C r C u h b N ~ 1 t ,mlon j1 ,2 lo0~0aMl /oo ,s /oso~012~O0210 .22J - - _

lo00

‘C K m!xedcryr(.l 800 700

5oo E 400 3

300 200 100 0

M n m i m s i b

g! $ 600

0.1 1 10 10’ 10J 10. 10” s ld time t 3171148552 0 IFW uurrrARou

Figure 2.2: Time-temperature-transformation curve of case hardening steel 16MnCr5 (AISI 51 15)

The composition of the produced workpiece material is a result of a controlled cooling speed. Nucleation and c stal growth occur in the pearlitic structured area by dzusion of carbon and iron. Ferrite and cementite lamellas appear in the material structure in case of transformation of austenite. The time for diffusion

processes is low applying an increased cooling rate and leads to a reduced displacement of carbon atoms. If the critical cooling rate is exceeded, a tetragonal distorted martensite will be produced at the martensite- starting-temperature Ms. During this extremely fast trans- formation the carbon remains in a constrained soluble state in the a-lattice. A comparison between tetragonal and cubic structured lattices shows that the carbon atoms are arranged on interstitial in the lattice as far as the production of martensite is concerned (figure 2.3). All three edges of the elementary cell of the body centered cubic a-iron are of the same length. By elongation of one and compression of the other two axes, a tetragonal lattice is formed. In the tetragonal lattice C-atoms are not obligatory arranged in every interstitial. The dissolved C-atom makes slip dislocations almost impossible and effectively hinders plastic deformation causing high hardness of martensite [3]. The specific volume of martensite is higher than the specific volume of austenite [4] [5]. The dilatometric curve of a steel with 0,9% carbon content (figure2.3) shows the change of length and therefore of volume during the transformation process. A characteristic sign is the increase in volume at the beginning of the generation of martensite at temperatures of Ms = 210°C. The dilato- metric curve does not develop linearly in case of rising the temperature up to starting temperature. The increase of martensite volume results in compressive stress in the austenite and obstructs further martensitic trans- formation. Thermal stress basically caused by temperature differences between workpiece surface and core is superposed to these transformation stresses [6] [A [8]. Both types of stresses suppress further progress in the martensite transformation [5] [9].

work(pIece material: steel with 0,9%C

martensite I X J Fsatom

-: H. sdllmnm

0.6 mm

4 0.4

6 0.3 g 0.2

2 0.1 0

4.1

o ~ o o m SM) c 800

V s w . msrtsns~s = 0.1298 d l g Vrw. = 0.1265 d g

temperature

317/1485EC 0 IFW

Figure 2.3: Change of structure during martensite

The carbon content influences martensitic hardening and should not be less than approximately 0,25%. Otherwise, the cooling rate necessary for martensitic transformation is difficult to reach. Case hardening is adapted to improve mechanical properties of workpiece materials with low carbon content. Carbonizing consists of enriching the subsurface with carbon in order to enable the subsequent hardening process. Other chemical elements such as boron or nitrogen are also used in the heat treatment of ferrous materials. Boronizing is usually applied to increase the hardness and improve the resistance against corrosion. The precipitation hardness of the material of the subsurface improves the stability of components exposed to static and especially to oscilla- ting loads. Hereby, the disperse nitride precipitation ameliorates the local fatigue strength and obstructs the movement of dislocations. The attainable fatigue strength depends on the amount of absorbed nitride, the depth of nitride precipitation and the strength inside the workpiece.

formation

548

Keynote Papers

2.2 Properties and demands In order to design mechanical products and machines with optimal economical and technical properties, possible origins of failures have to be identified. In machines with moving parts surfaces as well as volumes are loaded. Volume loads such as mechanical and thermal loads lead to deformation and fracture while surface loads cause corrosion and wear. Often both types of loads are superimposing (figure 2.4). Machine components being in contact with moving parts are simultaneously subject to wear and corrosion. This effect is called wear corrosion or frictional corrosion. It presents the main origin of friction fatigue fracture and has to be taken into account in machine part design [lo].

endangering functionality of machine parts

is. volume loads: surface loads: -mechanical

Figure 2.4: Main origins for failure of machine parts

Steel components as e.g. camshafts, gearwheels, bearings and cutting tools being exposed to high mechanical loads are often heat treated. Hardening and tempering or case hardening of ferrous materials lead to an improvement of mechanical properties of workpieces. Besides the improvement of strength and hardness, the fatigue strength is considerably increased [I I]. Figure 2.5 shows the influence of the workpiece hardness on rolling contact fatigue having tested gearwheels of different alloyed steels. It can be indicated that the rolling contact fatigue is increased with increasing hardness. The working life is also remarkably improved with increased workpiece hardness and so reaches about 10 time the initial state.

100

40

20

10

4

2

1 54 56 5 0 6 0 62 6 4 H R C 6 8

hardness :KH.b!&q

0 Y

J s 1 8 F H .-

loo0 3wo 5M)o Wmm2 goo0 tooth flank hardness HV

52W2@36c 0 IFW

Figure 2.5: Influence of heat treatment on rolling

Heat treatments which aim to increase the hardness of components or of work surfaces lead to the improvement of wear resistance. The subsurface hardness adjusted after machining or after being loaded in the required function is of high importance. A higher workpiece subsurface hardness generally increases the resistance against wear. Further factors like structure constituents additionally influence the resistance against wear.

contact fatigue

3 MECHANISMS OF CHIP REMOVAL DURING HARD CUTTING

Applying hard cutting as a finishing process requires the generation of machined surface by pure plastic defor-

mation. Therefore the stress, strain and temperature distribution the cutting zone is of interest. The proper understanding of the material removal mechanisms taking place during hard cutting is essential for process evaluation. The analysis of the work area is necessary to describe the chip generation in hardened materials. Depending on cutting parameters and workpiece material properties, cutting may either lead to continuous or discontinuous chip formation. Komanduri and Brown have compiled a detailed classification of chips produced by non homogeneous cutting, which are called wavy segmented and discontinuous chips 1121 (figure 3.1).

Figure 3.1: Classification of chip The wavy chip is generated by a cyclic variation of the chip thickness due to an oscillation of the shear angle. The discontinuous chip tends to form in case brittle materials are cut at low speeds. The segments of the discontinuous chip are completely separated by fractured surfaces. Meanwhile, a distinction between the segmented chip and the shear chip is difficult to draw and normal1 they are both classified as segmented chip U21 [I 31 [I 4f The behavior of materials causing elastic deformation, plastic deformation, cracks or localized shearing, also called adiabatic shearing, is decisively influenced by temperature and stress distribution in interaction with strain and strain rate [15].

3.1 Plastomechanics The stress-strain curve of hardened steel, gained in a tension test, is almost linear till fracture. There is practically no plastic deformation. Therefore, a main question is how and why smooth surfaces can be achieved with obviously required large plastic deformation. There are two theories to explain plastic deformation which are 0 the thermodynamic theory and 0 the hydrostatic theory. The thermod namic theory explain the formability of the

applied mechanical energy is almost exclusively trans- formed to thermal energy which heats up the material in front of the cutting edge. Under elevated temperature steel softens and get a high formability [ I6 [I7 [I81 [I91

machining". According to this theory there is a decisive influence of the heat conductivity of the tool material. It should not be too high to keep the heat in the chip formation zone and to generate a sufficient level of temperature. To examine this theory, Brand made cutting experiments under constant conditions varying only the heat conductivi of the tool material [3]. He could not find a significant in 3; uence of the heat conductivity on the cutting force (figure 3.2). This indicates that the thermodynamic theory can not fully explain the hard cutting phenomenon. This is underlined by force measurement if varying the cutting speed. Here, if self induced heating of the chip formation zone would be the dominant effect, the cutting force should decrease with increasing the cutting speed. Basically, the heat conduction is a time dependent effect. If there is less

work materia Y by heating up the chip formation zone. The

[20] [21] [22]. The effect is called "se I/ induced hot

549

time the thermal energy is more concentrated near the heat source. Thus, the temperature in the plastically deformed zone is higher and consequently the specific cutting ene y is lower. Investigations show that there is only a s m 5 influence of the cutting speed on force components [3]. Therefore, it is reasoned that the thermodynamic theory can not explain the plastic deformation of hardened material alone.

tool lnawlai

i i i n m ~ c o m ma

Figure 3.2: Influence of heat conductivity on the

The hydrostatic theory is based on the fact that the formability of a material is strongly dependent on of the stress state. This was first shown by v. Karmann who undertook his famous experiments on brittle marble. Under high hydrostatic pressure marble is plastically deformable. The model based on this effect is given in (231. Both low chip thickness and a strongly negatrve chi angle induce high compressive stresses and hig hydrostatic pressure.

cutting force

R

, I I ,

&da1:2 -w- &da2:1-absoopn(llon I drQ 3 : 2 . .d. r-

- D h m s a * W I r17nsm 0 m

Figure 3.3: Chip formation and Mohr's circle in hard

Appl ing Mohr's circle different, stress and strain limits can explained [24] (figure 3.3). Under biaxial tension (cirde 1) the material is fractured when the normal stress reaches the yield strength. The shear fracture limit is relevant under biaxial and compressive stress (circle 2). Final the material even brittle is deformed lastically if

with considerable high hydrostatic pressure (center of the circle) stands for stress state in front of the cutting edge with a little undeformed chip thickness and a highly negative rake angle. The chip thickness in this region in which the workpiece surface is generated is extremely small in comparison with the cutting edge radius and measure some micrometers. This geometrical condition leads to an effective rake angle of -60' to a. A high pressure is generated to remove the material in this part of the cutting edge. Flow chi s and segmented chips can be observed in cutting Lrdened steel. Generally, flow chips are generated in case the chip thickness is smaller than hc20pm. If the chip thickness exceeds 20pm a segmented chip will be formed [2] [25] [XI. For a chip thickness above 20pm the mentioned condition fails near the free workpiece subsurface. Hardened workpiece

cutting

the s '1: ear flow limit is achieved (arde3). fhe circle3

material is not plastically deformed in this region and instead of fbw chips, segmented chips are formed. The transition from continuous chip to segmented chiplshear localized chip also takes place with increasing cutting speed [27l. The material behavior in the work zone and so the thermo-mechanical mechanisms strongly depend on cutting parameters and especially on chip thickness h and cuttins speed vc Figure 3.4 shows the influence of the chip thickness on the chip formation process [26]. In 1964, Recht developed a classical model for describing catastrophic shear instability in machining [28]. According to Recht, catastrophic shear will occur in plastically deformed regions in a material if the slope of the true stress true strain curve becomes zero. Semiatin and Rao developed another model of shear localization which incorporates a heat transfer ana sis and material

temperature dependence of the flow stress and the strain rate sensitivity of the flow stress, to establish the tendencv towards localized flow 1291.

properties, such as the strain har r ening rate, the

-: H.Q -, vw IIlaTRoFw

Figure 3.4: Influence of chip thickness on chip

In 1985, Recht introduced the adiabatic shear theory to characterize the chi segmentation process during hard cutting operations. lierefore, themeplastic instability is where a decrease in flow stress due to thermal softening associated with an increase in strain more than offsets the associated strain hardening. Proponents of this theory refer to this to explain the chip segmentation. R is assumed that for some reasons a themeplastic insta- bility occurs along a line, extending from the tool tip and curving upwards to the free surface of the workpiice [30].

formation

Figure 3.5: Crack initiation in the formation of saw

In addition to themplastic instability which leads to shear localization other mechanisms can reduce the shear stre th within the shear band without a thermal softening e k . According to Nakayama, the segmented chip will generate, if the shear strain on the free surface of the workpiece attains an ultimate value . In a simplified model the crack initiates at a point Q o k e free workpiece surface. The free surface at Q must be a principle stress direction. Hence the shear plane includes an angle of d 4 with the free surface. If W, is the inclination angle and @ the shear angle, it is @ = d 4 - W,.

toothed chip

550

Keynote Papers

___- i.44;; ~ ;e , ;I;; - ~ crack initiation ! r e m a r k s

deformed region I

S h surface layer energy

I Semiatin and Rao 1983 dnldb drldr doid, ladlized flow

R&t 1985 doid- > doldr llW adiabatic Shear

I Nakayama 1988 yYln = yC l freesurface Y', independent

I of wtbng paramelei

I ' E I M E ~ w I /9 I..-.. ~mgulanties hoe surface

From figure 3.5 it can be seen that this relation is quite different from the theory based on the model of continuous chip formation.

In cutting experiments of a bearing steel of 760 HV (- 62 HRC), the rake angle is varied between y = -10" and y = -50". In spite of this wide variation of the rake angle, crack inclination 'Y, is almost constant and amounts to 'Y, = 30 [31].

Elbestawi investigated the chip formation during machining of hardened steel working out criteria for crack initiation and propagation [32]. Surfaces which have to be machined are not perfectly smooth but rough and composed of microscopic ridges, cracks, voids etc. Machining case hardened materials, high compressive stress creates subsurface material flow but particularly leads to the formation of crack in the free surface depending on the brittleness of the workpiece material. The irregularities in the free surface play an important role in crack initiation.

In 1974, Sih tried to incorporate the surface effects into a continuum by considering a layer thickness d to separate surface and interior regions. Sih suggested a surface layer energy failure criterion. According to this criterion, the crack will initiate at a location along the boundary layer of the workpiece where the local energy caused by loading exceeds some experimentally determined material constant. This layer is assumed to be a continuum such that its gross properties may be treated as an isotropic and homogeneous medium [33].

Cutting experiments show that the chip formation in hard turning starts with crack initiation near the free surface. Cracks propagate and end up in a plastically deformed region close to the tip of the cutting edge (figure 3.6).

Figure 3.6: Microstructure and chip formation

According to the hypothesis mentioned above, the crack initiates from the free surface to a point where transition from brittleness to ductility takes place due to higher hydrostatic pressure near the tool tip area. On the basis of this hypothesis Elbestawi concluded that the crack in the free surface will initiate at the critical angle cpcr against the cutting direction when the surface layer energy reaches its maximum value at the minimal applied cutting pressure.

This angle of crack initiation will remain constant and will not be affected by changing cutting conditions provided that the cutting material is very hard, brittle and homogeneous. Elbestawi analyzed the crack angle introducing an energetic theory. In accordance to Nakayama's theory he showed that the crack initiates at the free surface of the workpiece with an angle of cpcr = 33" [32].

In 1999, Shaw compared the theory of segmented chip to the adiabatic shear theory for large undeformed chip. He suggests that a thermally initiated process should be expected to have its origin where the temperature has its maximum which is at the tool tip. This corresponds to Recht's model but does not fit with observations made by [31] [34] who showed that the crack starts from the free surface to the tool tip along a relatively straight shear

plane. A shear crack should be expected to initiate near a point of maximum shear where the compressive stress is a minimum. Finite element analysis of stress along the shear plane shows a substantial increase in normal stress in processing along the shear plane from the free surface to the tool tip [35]. A higher normal stress will be encountered if a shear crack progresses from the free surface to the tool tip. A continuous gross crack may gradually be converted into a discontinuous micro- cracked region [34]. Figure 3.7 summarize the different models for chip formation in hard cutting.

3.2 Forces and stress distribution In cutting hardened steel, resultant forces and stress distributions in the contact area are substantially influenced by the hardness of the workpiece material as observed by Matsumoto. Figure 3.8 shows the relationship between components of machining force and the workpiece material hardness. It has to be mentioned that in this case an extremely high feed was applied. The forces occurring in cutting soft steels are relatively high and decrease if hardness increases. With the hardness exceeding 50 HRC the cutting force increases suddenly. The passive force is known to alter the maximum strain value in the chip and the chip type [36].

cumng speed vC = 90 mlmin prwess outer diameter turning wok malenal AISI 4030 deplh ofcul ap = 0 15 mm cumng material MC (/u?OyTC) feed f = 0 9 m m

t m l gwmelw DNGN 190804 F

6

UN

li-

tLu 4

0 2 3

m

2

20 M 40 HRC 60 20 30 40 50 HRC 70 mater ial hardness m a t e n a l hardness

sourre Y MatsUmOfO 3z8n8au e IW

Figure 3.8: Relationship between cutting force and

Depending on the hardness of the workpiece material two types of cutting mechanisms can be observed. With

steel hardness

551

I \ tool

waiwece material 16 MnCr51AIS15115

tmI rnatenal CBN 60-62 HRC

Y)YM 0 Brandl IFW

< N

dry culling 317114372c 0 IFW 9769

Figure 3.9: Cutting geometry and forces in hard turning

Chip formation takes place mainly in the region of the corner radius and the chamfer of the cutting tool. Due to small values of depth of cut, the effective tool cutting edge angle is determined by the tool holder [41]. The passive force F is the largest cutting force component. With an increasfng cutting length leading to an increasing width of flank wear, a significant rise of the passive force F can be observed. Consequently, in contrast to c8nventional cutting the force ratio FJF, is reversed [42]. Larger negative rake angle increases the cutting force F, only to a minor extent while increasing the passive force F remarkably. Nakayama explained these phenomena d t h the formation of saw-toothed chip. In machining of annealed steels, both force components increase with the increase of the negative rake angle [31] (figure 3.10). As shown above, cutting of hardened steel often leads to the formation of segmented chip. This phenomenon is directly related to a fluctuation of cutting force and stress distribution in the work area and influences the temperature distribution and so the process result.

Dynamic forces that fluctuate at a frequency over 10 kHz are difficult to measure. A conventional piezoelectric dynamometer limits the frequency response to about 1.5 kHz. However, an estimation of the relative change in the force components and the frequency of force fluctuation may be obtained by using wire resistance strain gages [34] or accelerometers [43]. The force to time relationship can also be related to the cutting process through high speed motion pictures [13].

400 workpece hardness * A HV760

\ ' U m

m,= 78' rake angle Ihear angle * - 46' m = w m = 2 7

source K Nakayama 328178839 C IFW

Figure 3.1 0: Effect of negative rake angle on cutting

The behavior of the cutting force to time dependence is shown in figure 3.1 1. The fluctuation in percent can be given applying the strain gage and measure 6 % of the cutting force. Lemaitre qualitatively analyses the behavior of the cutting force depending on the chip formation. The results are presented in figure 3.11. Just prior to the maximum force, the chip begins to slide up the rake face, and deformation in the chip is confined to a narrow zone extending from the tip of the cutting tool to the free surface of the workpiece.

forces

800 Ibr

400 , r=318mrn 0

workpiece rnalenal Fe 18 5 NI 0 52C time t

rake angle "srled I rake angle I = 7'

Cuning speed low I workpiece malenal Ck45lAIS11045 depth of wl v m e d I cutting speed vC = 103 mlmin

source J C Lemaitre ' IOY~UJ M C Shaw 378128&(1 Q IFW

Figure 3.1 1 : Force fluctuation in hard cutting

As stated above the rake angle in hard cutting is strongly negative because of a more or less large cutting edge roundness rB. a small depth of cut ap and a large corner radius r,. The rake angle is of significant influence on the stress distribution in the chip forming zone. Schmidt made interesting experiments with varying edge radii r or effective rake angle yefi respectively (figure 3.12) [44f In a depth of cut range of 2,5 yrn to 50 pm and an edge radius range of 35 pm to 115 pm the effective rake angle varied between -30" and -75". With the depth of cut a,, and the material removal rate Q, the cutting and passive force increase. Larger edge radii result in larger forces. Comparable results show the investigation of Nakayama [31] (see also figure 3.10). Schmidt demonstrate the strong influence of the effective rake angle on the chip formation. Smaller rake angle between -62" and -64" leads to chip segmentation.

The steep increase of the passive force with smaller rake angle is due to two causes: On one hand there is a pure geometric effect. With decrease of the rake angle a larger position of the thrust force (vector addition of feed and passive force) is transmitted in the passive direction. On the other hand the higher cutting force, a conse-

552

Keynote Papers

twwcolor digital micro-

quence of higher shearing, increase the passive force by friction on the rake face.

120 I , 120 , I l l I

Clearance angle a = 6'

I I I u

E P En

30

J a 1W 200 300 mm~1min500 -0 100 2w 3w m m J i m i n ~ ~

matenal removal rate Q, material removal rate Q,

312,2531DCQ IW

Figure 3.12: Influence of cutting edge radius on cutting

Concerning tool load and chip formation, the stress distribution in the work region is of decisive importance.

Fallbohmer simulated both, temperature and stress in tool and workpiece (figure 3.13). He was interested in the free surface of the chip-work interface. The simulation is carried out for orthogonal cutting process of 35CrMo4 (AIS14130) with PCBN cutting tools varying the depth of cut between 10 and 60 pm. As soon as the depth of cut exceeds 18 pm, c7 drops quickly. From a = 10 pm to a = 30 pm the daximum shear stress Jecreases to 88 % of its original value whereas the principle stress in the y-direction decreases to 71 %.

Additionally, local temperatures decrease because the distance to the tool tip increases with a depth of cut preventing from material softening which is necessary for plastic deformation. It can be concluded that the highest shear stress can be found at the free end of the chip workpiece interface. The above-mentioned shear stresses in combination with decreasing principle stress in the y-direction as well as decreasing local temperature lead to crack initiation with increasing depth of cut. Thus, a segmented chip is generated [25].

and passive force

Figure 3.1 3: Simulated stresses and temperatures

3.3 Energy and temperature During machining processes, the cutting energy is almost completely converted into heat [45]. The mechanisms of material deformation, friction and material removal lead to the initiation of heat sources in the work area. Regarding the surface integrity of machined components, the friction between rake face and the new generated workpiece surface is the most important mechanism.

In 1999, Ueda et al. investigated the influence of cutting parameters and workpiece material on the temperature of the cutting edge Tte (figure 3.14). A two-color pyro- meter was used to measure the temperature over the thermal radiation of the tool conducted through a hole of an internal tumed tube. Different kinds of steels in different hardness level have been investigated. It can be

I

' 3-comp0~)nt dynamanetw

- J

0 I=

Figure 3.14: Influence of cutting speed and workpiece material on the temperature

The investigation of the influence of the workpiece material hardness on cutting temperatures confirm the results of the examination of the machining force (figure 3.8). In a higher range of hardness leads an increase of the material hardness to an increase of the cutting force. With increasing the cutting force cutting energy is becoming higher and results in elevated temperatures.

With respect to the surface integrity the temperature distribution in the workpiece subsurface is the focus of the investigations carried out by Schmidt. During the machining operation with a prepared tool holder, the temperature of the workpiece surface is directly measured below the zone of material separation with an infra-red camera. The measurements show that two areas with different temperatures are formed below the cutting edge. The maximum temperature on the component surface develops at the beginning of the contact area near the material separation. A second maximum is produced if the minimum cutting depth is exceeded and only plastic and elastic deformation as well as friction take place. The temperature of the workpiece surface is about 350°C. These measured temperatures enable to determine thermal load in the workpiece subsurface [44].

Figure 3.1 5: Contactless temperature measurement

The thermographical measurement of the workpiece surface temperature as shown in figure 3.15 does not allow the measurement of temperature in the contact area tooVworkpiece. Schmidt adapted a simulation program to determine this maximum value of tempera- ture in the new generated surface [44]. Furthermore, he determined thermal power and thermal energy load of the workpiece subsurface in the contact area.

Thermal power, thermal energy and maximum temperature for varied cutting parameters such as depth of cut and feed rate are determined with temperature simulation. The thermal power per unit length Pa' is defined as a physical parameter which appoints the maximum temperature in the workpiece subsurface and predicts the appearance of white layers.

553

The prediction of temperature maximum has also been tried by Wobker introducing the contact surface related power Pa".

pa =P,- p . W . v , lk VB, I,'VB,

With increasing cutting time, the contact surface related power Pa" decreases. This is contradictory to measure- ments showing higher temperature with increasing flank wear. Subsequently, the correlation between contact surface related power Pa" and the subsurface temperaure is not possible (figure 3.16).

I 3.5 5M) w e a r and forces 200 350 ,bwer 1-

0' ' 0 15 30 mm 60 0 15 30 mm 60

cuning time tc cuttlng time tC

wohpiece malenal 16MnCrS5iAIS15115 6062 HRC

YIU~C(I D Brandl IFW

tool N20f l tC vc = 145 mimin I = 0 1 rnrn SNGN 120416 ap = 0.2 rnm dry cuning

312/1957OC 0 IN(

Figure 3.16: Relationship between force contact area

The relationship between thermal power per length unit and maximum temperature in the subsurface of workpiece is verified taking further results of Brandt and Winands into consideration. Higher temperature generally results in higher residual stress and in the appearance of white layers. Figure 3.1 7 shows that residual stress measured using x-ray diffraction increases with growth of thermal power per length unit. If the thermal load per length unit exceeds Pa' = 150 W/m white layers appear in the subsurface of the workpieces.

and thermal power

Figure 3.1 7: Subsurface quality and friction power per unit length

4 CUTTING MATERIALS AND TOOLS

4.1 Demands and process interference Reflecting upon the surface roughness, surface integrity and workpiece accuracy, demands on hard cutting operations are usually comparable to those on grinding processes. Therefore, certain boundary characteristics and requirements of applicable cutting tools can be defined with regard to the accuracy of the machined surface.

High indentation hardness of the cutting tool is required. Usually, it has to be higher than three times the workpiece hardness [31]. In hard cutting operations, this is important in order to prevent

deformations of the tool tip in the contact area of tool and workpiece. Therefore, a high resistance of the contact area against strong impact and stress is necessary.

A high hardness to Young's modulus ratio is required in order to minimize the quantity of local plastic deformation after the cutting tool has passed over. Hence follows that the accuracy deviation becomes essential constituent for the finishing of hard materials [31].

Due to the heat generated in the cutting process, the thermal conductivity of the cutting tool material influences the expansion of tool and workpiece. Materials with high thermal conductivity reduce the probability that deviations in the geometric accuracy of the workpiece occur [3] [47].

The high specific forces cause stress on the contact area between tool and workpiece. For this reason, cutting tool materials must have high resistance against mechanical stress combined with a high wear resistance.

High stability against abrasive particles in the microstructure of the workpiece material is required to prevent that grooves in the cutting edge develop as well as to promote that even in case of progressing wear the tool tip keeps its original geometrical shape [23].

High thermal stability of the cutting tool material has to be guaranteed reliably because the energy resulting from high specific cutting forces is almost completely transformed into heat so that extremely high process temperatures are produced in the area of the contact zone.

4.2 Applicable cutting materials and their properties As to hard cutting processes, the cutting tool materials require especially high resistant properties against specific cutting forces in connection with process temperatures which are caused by the high deformation resistance of the workpiece material [26] [48]. In figure 4.1, mechanical and thermal properties of different cutting tool materials are given as an overview.

Cemented carbides are available in variations with different characteristic properties. The most interesting properties are the high tensile strength and the high fracture toughness. Due to their high hardness, only fine and ultra-fine grained cemented carbides are of interest for hard cutting operations. These cutting tool materials are applied in kinematicly more difficult processes demanding complex tool geometry. These processes are hard drilling and milling operations especially for dies and mold manufacturing [26] [49] [50] [51]. Apart from this, the hardness of cemented carbides is insufficient for hard turning or face milling. The application of cemented carbide tools is also limited by their comparatively low temperature stability.

The most often applied cutting tool materials for hard turning and face milling operations are AIzOdliC- ceramics and PCBN [3] [26]. Their high hardness combined with a high temperature stability enables these materials to resist the thermal and mechanical loads in the hard cutting process. A most important difference between ceramics and PCBN is the value of fracture toughness. Compared to ceramics, PCBN-tools are favorable in interrupted cutting operations. Considering the effects on dimensional and form accuracy of the workpiece, a high thermal conductivity and a low thermal expansion coefficient is of importance. Both characteristics favor PCBN as the more adapted tool material for hard cutting processes [20] [52] [53].

The hardness of PCBN is surpassed by PCD. Even at low temperatures, the diffusion ability of carbon in ferrous materials is too high so that it cannot be applied in hard cutting of steel [23].

554

Keynote Papers

- y) density [glcrnl] us? __ C C hardness HV 30 r" g __

z p p ! kn fracture lauohnerr lMPal

Youngs modulus [GPa] -___

I PCBN I PCD I I cementedcarbide I Ceram,C K 10

1 4 0 1 5 0 3 8 5 0 3 4 4 3 35-42

1500 1700 1800 2500 30004500 4000 5000

590-630 300400 580480 680-810

1108 2 0-30 3 7-6 3 6 8-88

- __ __ __

.. _ ____

= 5 4 i

75-80 i 3 6 4 9 I 4 2 4 9 I source Sum!lomo. De Been Herlel Sandvik Kwpp Widia 317110775cO IFW 9609

Figure 4.1 : Mechanical and thermal properties of

4.3 Wear mechanisms The cutting processes of hardened steels are characterized by high mechanical and thermal loads that have to be endured by the cutting tool material. As to hard turning processes, hard and ultra-hard cutting materials are therefore required.

Usually, pure ceramic cutting tool materials applied in hard cutting do not reveal special advantages [ I ] [I61 [54]. AlzO3-ceramic possesses high chemical stability and hardness but the resistance to fracture is insufficient. Therefore, the main criterion for the end of tool life are micro-chippings on the cutting edge [55]. The higher toughness of SisN4-ceramics can not be utilized in hard turning. Definitely, the hardness is lower compared to aluminum oxide ceramics but the high chemical reactivity of the binder phases lead to dissolution of tool material. For this reason, usually mixed ceramic tools are applied as cutting tool materials for turning hardened steels.

Figure4.2 shows tool life in turning of hardened steel, using different ceramic and PCBN inserts.

different cutting tool materials [23]

100

min

c

h(~irce o 0ranat IFW 3171iV112cQIFW B E 4

Figure

Sic-whisker reinforced ceramics show potential for the cutting of hardened steel, but compared to Tic-whisker reinforced ceramics, the achieved tool life is low. Here, the micro-resistance of the matrix shows the same wear behavior as pure AlzO3-ceramics. Although the tough- ness of Sic-whisker reinforced ceramics is higher, this advantage cannot be used in hard machining. The explanation is given facing the chemical resistance of Sic which shows solubility against the machined steel material, so that the effects of the reinforcement are poor. As to the tool wear behavior of SisN4-ceramics, this is similar.

Concerning the application of ceramic based tool material, the highest potential for the cutting of hardened steel is shown by Tic-reinforced Alz03-ceramics. Here, a comparatively low rate of flank wear is combined with a high resistance against cutting edge chippings. The comparison between cutting tool materials of different producers reveals just small deviations [3] [26].

4.2: Efficiency of different cutting tool materials

PCBN is the cutting tool material with the longest possible tool life. But the composition of the tool material exerts an influence on the wear mechanisms. The material properties of PCBN can be influenced by the PCBN content, the grain size and distribution as well as the composition of the binder phase which can be ceramic or metallic. For hard turning operations usually ceramic binders are preferred. The range of PCBN- content in the different tool materials in figure 4.2 is varied in a percentage of 50% to 90%. Even if a high content of PCBN increases the hardness of the cutting material, the tool wear by using these materials is comparable to the A120dTiC-ceramics. The mechanical and thermal loads in the cutting process require additional properties of the tool material, especially at low values on uncut chip thickness. Here, the tool material with a lower content of PCBN shows advantages due to lower thermal conductivity and higher toughness. Furthermore, additional properties of the cutting tool material are of importance, This is based upon possible deviations of tool life in relation to different cutting tool manufacturers. Finally, the highest potential for turning hardened steel is shown by PCBN-tools with a low content of PCBN of small grain size [20] [53] [56] [57].

Figure4.3 exemplifies wear pattern of a PCBN tool characterized by fine scourings. Chips formation exclusively takes place on the chamfer and in the cutting edge radius. The rake face as well as the flank wear land does not show any deep grooves from the cutting action. Especially the cutting edge itself does not show excavations. Such wear pattern indicates that it is possible to achieve high surface quality and size accuracy of the workpiece.

Figure 4.3: Wear pattern of a CBN-tool after hard

Tool geometry determines the contact conditions between tool and workpiece in the shear zone. This exerts an influence on the wear mechanisms because different tool geometry causes specific mechanical and also thermal loads in the contact area.

The interaction of the tool wear progression and the cutting forces using tools with different cutting edge radii is shown in figure 4.4 [44] [58].

Generally, the tool wear progression divides into two areas of different gradients. The gradient of the width of flank wear decreases in dependence from increasing cutting time, so that a linear relation does not exist. Tools with lower cutting edge radius show higher wear resistance. These tools allow a 1.4 times higher cutting time until reaching the width of flank wear of VBc = 100 pm, compared to the tool with the higher cutting edge radius. Independent from the tool wear, the cutting forces are on a higher level if using a tool with higher cutting edge radius.

Taking the cutting time into account, the tool with the lower cutting edge radius shows smaller flank wear. This advantage is limited by the possibility that cutting edge chippings may occur. A higher cutting edge radius stabilizes the cutting wedge.

turning

555

wear I

‘0 25 50 mm 100 cutting time

workpiece matenal 16MnCr5lAIS15115 60 62 HRC cutting matenal

source J Schmidt IFW PCBN CNMA i2Moa

cornponenls of resultant force

150 - - 5 . ; up 100

8 ; (o LL 75

E 8 ? i 50 g g 25

3

U

30pm tWpm 30pm 100pm flank face wear VB,

cutting parameters vc = 150 mlmm f = 0 05 mm ap = 0 05 mm dry Culling

312125Mx1c Q I F n

Figure 4.4: Influence of tool wear on components of

The SEM pictures in figure 4.5 show the wear pattern of two tools with the cutting edge radii rp = 35 pm and r - 130 pm at the flank wear of VBc = 100 pm reached at Jierent cutting times.

resultant force

outer diameter turning loo1 geometry CNMA 120408 cunlng parameters workpiece 16MnCr5. AIS15115

cutting t cd PCBN BN 4 dry cutting

vc = 150 mlmin f =O05mm -1 ap = 005 mm

60-62 HRC

s o ~ r u t J Schmidt IFW 31Z2560tc 0 IFW

Figure 4.5: Views of cutting edges with same wear

In both cases, the wear is very regular and characterized by abrasive and fine scouring areas on the flank. This indicates, that the mechanical and thermal loads are in a suitable range for reliable process conditions. Both tools show crater wear. Concerning the tool with the higher cutting edge radius, the crater wear is localized completely in the area of the cutting edge radius. The tool with the smaller cutting edge radius predominately shows the crater wear in the rake face. This weakens the cutting edge, because the possibility of chippings increases in further use of this tool.

The left-sided SEM picture shows a detail of the worn cutting edge with build up layers. These build up layers are confirmed by EDX-investigations and indicate the presence of a flow layer of workpiece material in the area of the cutting edge during the cutting process. The flow layer is generated by the high pressure due to the high mechanical and thermal load in the small contact area between the tool and the workpiece [44].

A model developed for grinding operations confirms that a highly viscous boundary layer between the strongly negative rake angle of the grinding grains is formed out which influences the chip formation [59]. It can be assumed that no material removal does occur if the uncut chip thickness is below a critical value. Until the critical chip thickness is reached, both pressure and temperature in the contact zone increase to rather high values until the temperature of chip formation is reached. There is a “chip formation equilibration temperature” which depends on the main tensile strain of the workpiece material. Such behavior can also be observed in hard cutting processes. Different thermal conductivities of tool materials do not cause different levels of cutting forces so that the conclusion can be drawn that there is a constant temperature of the chip formation [3]. Because of this, it has to be assumed that there is a highly viscous

progression

boundary layer between tool and workpiece in the contact area.

The presence of a flow area of workpiece material also influences the cutting tool material. PCBN supports that boron can be solubilized to form out a boron-oxide film on the tool surface at high cutting temperatures, which can act as a layer decreasing friction. Such boron-oxide films were detected in former investigations [25] [60].

A detailed representation of wear patterns on cutting edges recovers the contact conditions at the active surface of cutting tool with highly negative rake angles in hard turning, see figure 4.6.

cutting paramelen vC=150m/mln 1 = 0 0 5 m m a p = 0 0 5 m m t = t 2 m m d,yccul

3 1 ~ ~ 7 8 4 0 ~ m IFM

Figure 4.6: Typical wear pattern of a PCBN-tool and chip in longitudinal microsection

A tool with a high negative cutting edge radius (rll,= 130”) has been selected. There, the characteristic conditions at the cutting edge can be observed in a comparatively wide area. The contact area of tool/workpiece/chip can be divided into five areas [44]. The zone 1 and 5 do not show tribological contact between the workpiece material and the flank of the tool. Also in zone 3 wear cannot be found. It seems that there is no relative velocity between tool and workpiece which indicates that a stable wear preventing flow layer has been formed. Here, the wear protection of possible boron-oxide films is possible. The flank in zone 4 is characterized by scouring areas due to abrasive wear. Elastic and plastic deformed workpiece material slides across the tool area. This zone is beyond the flow layer in the contact zone so wear occurs at the cutting tool. Zone 2 is also characterized by wear due to the material deformation and continuous contact to the chips.

The contact zone along the cutting edge radius in the transition region between major and minor flank can be divided into four areas. In zone B, there is intense contact between tool and workpiece. In the further area of the contact arc, the minimum chip thickness is increased and no chip formation occurs. Zone C is characterized by point contact between workpiece-tool material. Because of the elastic deformation of the workpiece, the zone C has a relatively wide tangential extension along the cutting edge. The boundaries to the zones A and D are reached if no contact between tool and workpiece material can be observed.

The bottom side of the chip in figure 4.6 shows a strong white layer which is formed almost independent from tool wear. This indicates the strong plastic deformation of the tool material during the chip formation [3] [61]. It c an be estimated that this area contains constituents of the flow layer from the contact zone which is essential for the chip formation in hard cutting processes with negative rake angles. A model is given in figure 4.7.

In this model, the uncut chip thickness exceeds the cutting edge radius. In this case, the minimum chip thickness hmn is decreased below the point “I”. Here, chip formation does not occur but elastic and plastic deformation of the workpiece material. A more detailed representation of this behavior is given in the detail “Y’. The point “I” is localized at the cutting edge and defined by the stagnation angle 0 = 25“ = 90”-lyeffl. After the

556

Keynote Papers

workpiece material has passed the po,int “II”, the elastic proportion Ahel springs back success~vely moving with the tool while behind the point “111” the plastic deformed proportion Ahp[ leads to a lasting deformation. The velocity profile in the flow layer (detail “ X ) is a result of the local temperatures, viscosity and shear stresses in this area [62].

Figure 4.7: Material flow at the cutting edge

As shown, a favorable wear behavior of the cutting tool in cutting hardened steel depends on the theoretical base but also on low continuous friction behavior in the contact area between tool and workpiece. Depending on the cutting tool, even cutting edges, homogenous fine grained microstructure of the tool and possibly low cutting edge radius lead to an increase tool wear. Additionally, a small cutting edge radius decreases the extension of the zones 3 and C and supports the chip formation on a wider area on the contact arc.

5 QUALITY OF WORK

5.1 Machine tool requirements Up to now, conventional lathes are usually used for hard turning. This kind of machine tools is often modified in some machine parts to be suited for industrial scale manufacturing. The modification of vital machine components like headstock, tailstock, foundation and slides show higher accuracy, and linear path-measuring systems lead to higher dimension and position accuracy. The application of such types of machine tools enables to reach workpiece qualities in the range of IT6-7 in hard cutting under appropriate condition reliability [63].

Figure 5.1 : Characteristics of high-precision lathes

Recent efforts in machine tool development lead to new machine concepts of high-precision lathes to improve the quality of hard turned workpieces significantly (see figure 5.1). The accuracy demands on these machine tools are comparable with grinding machine tools concerning static, dynamic and thermal stiffness as well as accuracy of the spindle system and slides [32] [64]. Dimensional shape and accuracy in the range of IT 4-6 becomes possible by the use of high-precision lathes [2].

suited for hard turning

Precision turning lathes with components designed for increased accuracy and high-precision turning lathes correspond with the demands of grinding machines.

Typical characteristics of precision and high-precision machines are shown in figure 5.2. The precision lathes are equipped with roller bearing spindles, slide ways and NC-controls with a resolution of 1 pm. Here, many applications reach the accuracy of grinding processes. These values can be quantified in ISO-qualities IT6-7.

Surface roughness can be realized in the area between R, = 2-6 pm. Higher demands for workpiece accuracy exceed the range of application of these machines conceptually based on conventional machine components. For high grade applications as e.g. for gear components or roller bearings with dimensional and form accuracy in the range of IT3-5 and surface roughness Rz c 1.5 pm, the precision of hard turning operations was usually insufficient, so the finishing operation must be grinding or honing.

The substitution of these processes by hard turning requires an adapted machine conception (see figure 5.2). Special attention must be paid to the geometrical and kinematical accuracy as well as on the stiffness of the machine tool. Additionally, the stress relief of the workpiece due to the possibility of sparking-out processes in grinding must be realized by higher precision of hard turning machines. Each deviation of the relative position between cutting tool and workpiece causes inaccuracies of the workpiece [42] [65] [66].

machine type

machine bed

spindle drive

spindle bearing

max revlmin power

transverse slide

longitudinal slidc

cooling system

IT-quality

precision lathes

slant bed 30’ 45’ 60’ vertically i ptck-up

rneehanttecasl / stlicateConcrete mineralile ConcrBIe filled gray cast iron

wnv bell dnve dired dnve

preoslon roller bearing (steel ceramic). hydrostatical beanng

4000.6300 llmin

16/22 - 4 3 5 7 kW (100%/50% WT)

slide way linear roller Slide

slide way linear roller slide hydrostatic slide

spindle cooling circuit (waterlair). 011 coaling urwbt (spmdle slmcl~re column)

5 7 mainly 6 (measunng sensor shafl encoder longitudinal. linear scale for plannmg

1 high precision lathes

slant bea 4~ honzantaie

natural granile

direct dnve

hydrosbtial beartng aetostalical bearing

3500. 10000 t /mw 1 5 7 5 k W

hydroslallc siMe

hydroslalic slide

011 cooling

5 5 (measuring by linear scale) 31324648 0 IFW

Figure 5.2: Characteristics of hard turning lathes

Recent developments of high-precision turning lathes adapted to the special demands in hard turning lead to an increase in workpiece quality [42]. These high- precision turning lathes are characterized by hydrostatic guideways as well as spindles and supports which are arranged on both sides of the machine bed of natural granite. Hydrostatic bearing systems embody the possibility to increase the deviation of radial run out of the spindle center axis on values distinctly lower as 1 pm. Furthermore, a sensitive adjustment of position and motion with a resolution in the area of 0.1 pm is enabled preventing stick-slip. To increase the thermal stability of the machine tool, heat producing components like electric and hydraulic devices are located separately from the machine.

High-precision turning lathes equipped with components comparable to the technology of grinding machines enable to gain workpiece accuracy of grinding processes by hard turning. In this case it must be mentioned that the reproducibility of higher workpiece accuracy also depends on the tool wear so the constraints of process monitoring is more extensive than in comparable grinding processes.

5.2 Geometric accuracy A main criterion of the workpiece quality especially for finishing processes is the surface roughness. In many applications, conventional lathes deliver sufficient surface qualities in the range of R, = 2-4 prn.

557

the theoretical path has to be implemented in the NC- control. With linear interpolation of these bases the required profile can be achieved.

wnvex microprofile numenc wntrol

0 2 mm 4 & 8 12

6 0 4 2 ~ ~ cumng 1001 rnatenal PCBN unworn 1001

0 005 0 1 mm 02 feed f

~ j 7 . ~ / ; . I r l s a n j rG a ,=o lmm 6' j -26O 14" I55 ' / 6' 1 2 0 mm

6 0 005 0 1 mm 0 2 feed f

3131201878 0 IFW

Figure 5.3: Surface quality on high-precision lathes

Higher requirements on surface quality can be fulfilled by the use of high-precision lathes. Figure 5.3 shows that a surface roughness in the range of Rz = 0.5-1 pm can be achieved. Fine grinding or honing qualities are defined up to a surface roughness of R, < 1 pm. Feed rates of f = 0.1 mm deliver surface qualities in hard turning comparable to fine grinding or honing [42] [67].

Besides the surface roughness, for some precision components like bearings or gear parts the surface microprofile is of high importance for the functional behavior. Some of these precision parts are characterized by a very fine and smooth contour profile. Fine grinding machine tools can realize the production of fine contour requirements by highest NC-control resolution and an appropriate dressing of the grinding tool. Due to the fact that hard turning processes only offer limited possibilities in case of selecting the tool geometry, the precision of the surface profile completely depends on the NC-control resolution [42] [68]. To realize sophisticated requirements on microprofiles, the application of high-precision lathes is necessary, as figure 5.4 shows.

By the use of conventional lathes, steps in the microprofile in the range of the NC-control resolution occur. In the case shown in figure 5.4, the NC-control resolution is about 1 pm which is a typical value for conventional turning lathes. Using high-precision lathes with an NC-control resolution of about 0.05 pm, single steps cannot be recognized.

conventional lathe NC-control resol~lion 1 pm

high-precision lathe NCconIrol resoIuI10n 0 05 um

I I

1 I Jf"s.cllr-

c u m 9 Parameter workpie~e malerial

a, = 0 05 mm f =O05mm

cuning material PCD vC = 150 d m i n CuZn35N12 geometry SPUN 120304 yellow brass 270 e laa1 313122309e 0 IFW

wne

d, = d, + 10 pm d,

Figure 5.4: Surface microprofile in cone turning

Because of the hydrostatic bearing, the ability of the slides to move in steps in the range of 0.05 pm avoiding stick-slip behavior is the pre-condition to achieve a convex microprofile. Due to this property it is possible to generate microprofiles with shape accuracy required e.g. by roller bearings to avoid unfavorable edge pressures.

An example for the application of high-precision lathes is the production of an inner ring of a roller bearing with very close tolerances for the profile contour. As figure 5.5 shows, a smooth profile without recognizable NC-control- steps can be realized after machining. Hence, cylindrical roller bearings permit the use of hard turning as a finishing process. In order to machine this microprofile,

max crowoing 1 6 pm

requ'red raceway profile - r'i

workpiece matenal

60 62 HRC bearing wning malerial PCBN tool geometry DNMA 150620

cylindnca roller

NU208 nner nng

1WC6 AlSl 52100

313i20192e 0 FW

Figure 5.5: Microprofile of a cylindrical roller bearing

One of the main benefits of hard turning is the possibility to avoid the use of cooling lubricants. On the other hand, in the case of dry machining, form errors gain more importance. Regarding hard turning as a finishing process, low tolerances have to be fulfilled. One important source of error that has not been focussed so far is the effect of process heat in hard cutting. The form errors of the workpiece occur due to the thermal expansion of the workpiece and cutting material and lead to a reduction of workpiece diameter in feed direction [69]. Additionally, the thermal effects lead to deviation of parallelism on the surface lines.

Since being able to move in steps in the range of 0.05 pm, newly developed high-precision lathes offer the possibility to compensate form errors due to thermal effects. This is a requirement to implement a correction in the NC-control and to receive a smooth profile without recognizable NC-control steps after dry turning.

Figure 5.6 gives an example of the compensation of form errors caused by thermal expansion in machining the microprofile of the cylindrical roller bearing mentioned above in figure 5.5.

raceway

requ,red racway pmfl1e canectfon referewe prollle ?+t'::,!:: pfi- -.

d, = d, f K -

tml entry d,

0 4 9 5

wohpnu, L C d arlbnp paramelw WkplBCemaIalal cy lhdml mller beannp cutllng mslenal PCBN vC i t50mlmin tCCC6 AlSl 52100 NU208 inner r q geanetry DNMA 150616 S ap = 0 05 mrn 6042 HRC

I = 0 0 5 m m 313122310e 0 IFW VB, = I00 prn

Figure 5.6: Compensation of form errors due to thermal

For the compensation, a correction reference profile with different gradients described by the correction factor K is used. Lower or higher K-factors lead to unallowable deviations of parallelism. Thus, by using high-precision lathes it is possible to compensate form errors, above all in case of dry machining, to fulfill the high requirements of workpiece quality.

A more common way to increase workpiece quality preventing form errors due to thermal effects is the use of cooling lubricants.

In figure 5.7, the differences in dimensional accuracy between dry and coolant supported cutting processes are shown [69]. The thermal expansion of tool and workpiece material causes a diameter deviation in a range of 12 pm

expansion

558

Keynote Papers

respecting a feed length of 40 mm. Respecting the thermal expansion of the tool and workpiece material, the application of cooling lubricants exerts a positive effect, so that no deviation of diameter does occur during the cutting. This means that the dimensional accuracy is not infected by the process heat, and a compensation of the NC-control is not necessary. Hence, the use of cooling lubricants is recommended if the control resolution is not sufficient to compensate form errors.

vc = 180 mlmin f =008mm ap=02rnm

Aagneuc chuck

11

b . , . I 10 20 mm 40

feed length sourca J Liarmann

1wI matenal

lml geometry CNMA 120412 /_I VB, = 0.09 rnm 1 workpiece: 1W)Cffi. AlSl 52100 62 HRC workpiece geameQy Do=lOOmm D, = 80 mm

Figure 5.7: Form errors in hard turning due to thermal effects

5.3 Physical properties The functional characteristics of a machined part essentially consist of volume effects, workpiece geometry and surface roughness as well as surface integrity. Especially geometry and surface attributes are strongly influenced by the cutting process. Regarding hard cutting is a finishing process, the influence on the machined part is very important for the functional behavior of the workpiece. To achieve high applicability of machined parts, the final cutting operation must be carried out with small cutting depth. So, the influence of mechanical and thermal loads on the functional characteristics of the workpiece is decreased. Therefore, deviations in dimensions should be reduced due to low cutting forces, simultaneously low values of surface roughness can be achieved and influences on the surface integrity are minimized.

Main criteria of the functional workpiece quality are the geometrical surface integrity (surface roughness, dimensional and form accuracy) and the physical surface integrity (microstructure, hardness, residual stresses), see figure 5.8.

Figure 5.8: Stress distribution and surface integrity in

Due to the combination of extremely high temperatures and mechanical loads in the contact zone, the austenite temperature of the workpiece material is reached in contact times of approximately 0.1 ms. The workpiece material is self-quenched, structural changes will have to be expected and residual stresses influence the surface integrity. In this case, the relation to the tool wear must be taken into account. Increased tool wear leads to higher thermal process energy related to increased

hard cutting

friction. This is indicated by higher cutting forces which usually increase linear to the tool wear. Here, the back force shows the steepest gradient. The development of structural changes in connection with increasing tool wear is compiled in figure 5.9 [70] [71].

In relation to the tool wear, a white layer connected with a tempering zone is formed out. In this rehardened zone, an increased hardness of approximately 1000 HV0.025 can be found in contrast to the hardness of 800 HV00.25 in the primary structure [70]. Depending on increasing flank wear land VBc, the back force and the structural changes increase. Furthermore, in all cases tensile residual stresses are found directly on the surface. The amount of tool wear leads to an increase of tensile residual stresses, here up to 600 MPa for VB, = 200 pm. Not only the surface value of residual stresses is influenced, but also depth profiles are measured in order to investigate residual stresses in deeper heat affected material regions, see figure 5.10. Depending on the tool wear, the residual stresses beneath the machined surface are measured from the surface to a depth of 90 pm. The cutting tool material has no significant influence on the depth profiles, the crucial criterion is the tool wear.

Figure 5.9: Surface structure as a function of tool wear

With increasing flank wear, residual stresses on the surface shift to tensile range. Close beneath the surface, compressive residual stresses are detected. The maximum level occurs in small depth. Increasing flank wear shifts the maximum to deeper material regions. Tensile stresses at the surface are caused by thermal influence due to friction and plastic deformation. On the other hand, mechanical loads lead to compressive residual stresses which admittedly influence deeper material regions [3] [36] [68] [72] [73] [74] [75].

800

MPa

Mx)

400

200

0

200

-400 ' 1 I . l 0 10 20 30 40 50 60 70 pm 90

depth yxim 0 Brandl IFW 317111038~0 IFW 9113

Figure 5.10: Residual stress as a function of tool wear

Besides the influence on surface integrity in case of residual stresses, the formation of white layers characterizes the surface of hard cutting processes. In order to discuss the mechanisms of transformation in the material structure, it is necessary to analyse the material composition and structure.

The distribution of the chemical elements near the surface is an important criterion for the examination of the effects during the turning process. In case of hard

559

turning, it is of interest whether the development of white layers changes the distribution of chemical elements in the transition region to the annealing zone compared to the crude material structure. Figure 5.1 1 shows the results of measurements with optical glow-discharge spectroscopy on a workpiece.

workpiece material 15 MnCr S5 (ASTM 5115) 50 62HRC

9. = 145 rn rnin 1 = 0 1 m m

s, a , = 0 2 m m

measurement technique

glow discharge aplical emlssion spdrometty

317112985 0 IFLV

Figure 5.1 1 : Element distribution in hard turned surface

The white layers with a thickness of 10 pm are produced by a worn tool of mixed ceramic (VBc = 280 pm). All elements which have been examined (Mn, Cr, C, Si, S) are equally distributed related to the removed workpiece material, and neither in the zone near the white layer nor in the other material structure changes in the distribution can be found. References for diffusion of elements or chemical reactions can not be derived.

Figure 5.12 compares the result of EDX-examinations of workpieces which have been machined with differently worn PCBN-tools. Additionally, the microstructure of subsurface area and the remaining austenite is determined and compared to the crude workpiece material.

If the tool wear is low, the microstructure can be compared to the unmachined part. The amount of remaining austenite does not essentially differ from this part. But as soon as the white layers are formed, a considerable increase of the amount of austenite can be observed in the subsurface. On the confirmed supposition that the measured values have mostly been taken within the area of the white layer, the percentage by mass of austenite has been 73%. White layers which have been gained by means of hard turning consist of about 213 of austenite notwithstanding the applied cutting tool material.

Figure 5.12: Relation between subsurface

White layers consist of over 60% austenite which is a non-etching white component in contradiction to the dark martensite scorching. The martensite components also show a mostly tetragonal twisted lattice, and can hardly be etched. The fact must be emphasized that no amorphous layer exists. Apart from the EDX- examination, additional links come from metallographical analyses. Another argument confirming that white layers are crystal-structured is the following: Amorphous material is not conductive, but the conductivity of the

microstucture and austenite content

surface of hard-turned workpieces with white layers is, as examinations confirm, not reduced.

In order to receive white layers by the machining of hard turned materials, two conditions must be given:

The contact temperature must exceed austenizing level of the workpiece material. The cooling of the surface area must take place in very short time.

Therefore, a description of the mechanism about the formation of white layers must be able to explain the extremely fast heating phase which is connected with austenization. In hard cutting, the contact time between tool and workpiece is often less than 0.1 ms even in case of increasing tool wear. For this reason, the conditions leading to changes in the material structure fundamentally differ from the conditions in the grinding operations.

In hard turning, the contact time is very short as well as the time for heat conduction. Besides, just a single contact takes place between tool and regarded element on the workpiece surface. Based on these conditions, the reason for the high temperatures of the machining process and the short time of heating must be a combination between friction and plastic deformation.

Plastic deformations can produce a high amount of heat in a short time which also heats up the workpiece. Moreover, the time which is necessary to austenize is considerably short.

The following thesis resulting from these reflections will be discussed now:

"In hard turning without plastic deformations no white layers are produced."

Comparing the processes on the rake face and flank face in case of different tool wear, the absolute and specific effective forces are on a similar level. Merely considering chip formation and chip deflection in the contact area, the wear of the flank exerts just small influence. This can be regarded in the shape of the chip bottom sides in the chip roots of figure 5.13 where significant white areas are formed out. Here, the high process temperature in combination with the high grade of plastic deformation is significant [3] [76].

Figure 5.13: Chip roots in hard turning depending on

Regarding the subsurface of the workpiece material after turning, the formation of white layers significantly depends on the tool wear. The plastic deformation caused by unworn tools appears as material separation directly in the chip formation zone. Increased tool wear changes the deformation conditions in the contact area as well as at the region of the tool flank.

The effects of wear cause increased upsetting deformation and chip deflection as well as shear by friction. Thus, the increase of the wear of the tool flank influences the formation of the workpiece subsurface zone with respect to two main aspects:

tool wear

560

Keynote Papers

Increased frictional heat is caused by extended contact time due to geometrical change of the flank of the tool.

Plastic deformations lead to higher temperatures and support the lattice shearing so the austenization of the workpiece material is supported in the machined subsurface area.

The combination of high plastic deformations and extreme short heat up times cause the lattice shearing with the formation of a very fine austenitic grain structure. Grain growth is almost impossible due to the fast chilling after tool contact. The combination of martensite and extremely fine grained austenite structure constituents causes the high hardness in the area of the white layer.

Due to the microstructural changes during hard cutting, the surface integrity is affected. The residual stresses of the austenite material components are clearly shifted towards compressive strain. In the white layer, measurements of martensite show more than 600 MPa, the dominating austenite is only about 150 MPa (figure 5.14).

Figure 5.14: Structure and residual stress of white

The martensite tensile residual stresses are caused by lower specific volume and consequently higher density of the austenite. The surrounding martensite is obligatory set under tensile stress while the austenite is forming.

Because white layers only occur directly on the surface, the tensile residual stresses are restricted to the boundary layer of the workpiece, too. Under the surface and as a consequence of high mechanical stress, compressive residual stress occurs. Its maximum is localized under the surface. From these overlaying influences, the curve of the residual stress which is typical for hard turned work pieces with increasing tool wear is derived.

layers

5.4 Functional behavior The reliable behavior of machined components, especially the functional surfaces, is an important criterion in many technical applications. Hard cutting processes have decisive influence on the surface integrity. This affects the fatigue behavior and is a crucial aspect for the application. Important applications are rolling contact and durable alteration of loads [74].

In motive application, the rolling load of the surfaces of components is common. According to this, the evaluation of rolling fatigue life concerning hard turned surfaces is a decisive prerequisite for applicability. An adapted testing method represents the long time investigation in roller type test stands [77] [78]. Test conditions for rolling fatigue life are performed at a Hertzian stress of po = 3600 MPa. The limit yf fatigue strength is achieved at a number of N = 5 x 10 overruns in case the area of pittings is smaller than 1% of the contact area.

Figure 5.15 shows the influence on the subsurface microstructure in rolling contact loads. The original functional surfaces represented by the sample with N = 0 overruns have been turned with a wear affected cutting tool (VB, = 100 vm) and a feed rate o f f = 0.1 mm. Due to

these conditions a white layer of small thickness is generated. Interesting conditions of obseyation are the beginning of rolling contact load 9 N = 10 overruns and at the end of the test at N = 5 x 10 overruns.

Figure 5.15: Microstructure after hard turning and

The rolling contact load does not show an influence on the microstructure within the first N = lo4 ovejruns. At the end of the rolling contact test (N = 5 x 10 overruns), petty damages in the guide ways can be observed. The micrographic analysis of the stressed area reveals the damages as planishing and removal of material. In macroscopic view, the damages are localized as lines of pittings following the peeks of the feed marks which occur in turning. Additional tests using tools with increased wear rate or higher feed rates for preparation of the test samples showed that the origin of pitting- appearance is connected with the geometry of feed mark and the highest surface roughness peaks. This indicates that by the mere presence of white layers, the surface integrity is of minor relevance for roller bearing fatigue life of machined parts. Of essential effect is the surface geometry the latter mainly influenced by the feed and wear rate in the production process.

Higher demands on surfaces in rolling contact loads are given if slip contact is possible. In figure 5.16, the effects of slip-load in rolling contact is described. The slip free contact allows the achievement of the defined roller contact fatigue life even if the test sample is turned with the relatively high feed rate o f f = 0.15 mm and a tool with increased wear (VB, = 200 pm).

roller contact fatigue test

Figure 5.16: Surface topography after rolling contact testings concerning slip-load

When slip load occurs, the samples parts fail, even if surface roughness of the sample is decreased by the reduced feed rate to f = 0.1 mm. Tolerable surface damage in the rolling contact fatigue life tests under slip load is achieved by turning the test sample with low wear and feed rates. Here, the surface roughness of the roller exercises the main influence on the rolling contact fatigue life.

Besides microstructural influences from hard turning, residual stresses in the subsurface area of functional surface are of high interest concerning the fatigue life of components. Figure 5.1 7 describes the development of

561

residual stresses in dependence on the load time in the roller time test stand. The samples were produced concerning low feed (f = 0.05 mm) and tool wear (VBc = 0.1 mm), slip and slip-free contact conditions can be observed.

The slip-contact conditions show that surface tensile stresses parallel to the cutting direction 0 1 1 decrease strongly after the first overruns so the area of compressive stresses is reached very soon. In contrast to this, the reduction of tensile residual stress 0 1 1 in slip- free load is small. The residual stresses perpendicular to the cutting direction were originally in the compressive area and change just slightly in compressive direction. In case of roller contact loads, residual stresses in hard turned surfaces change to values in direction of compressive stresses which is known for its positive influence in case of the fatigue live of components.

sample matenat lOOC6 AlSl52100 6 M 2 HRC

cullmg paramelem vc= 125 dmin aD= 0 15 mm f =O05mm

rdler wntacl set up w m 2960 limii tteman pressure 2400 MPa SllP 0 x 4%

om, parallel lo vC

A99 I 0 10 100 10’ 1 r s load time t o perpendicular lo vC

e ~ n m e n l s WZL. Aachen 31300527 0 IFW

Figure 5.1 7: Surface residual stresses depending on load time in roller contact load

Another important component characteristic is the fatigue strength against bending or oscillating loads. Here, sensitive influences from the area of the subsurface, surface topography, microstructure and residual stresses are known. As an example, case hardened components are common in automotive applications and the reliability is very important. Bending fatigue strength is a main criterion for components because the maximum of the load directly occurs in the surface, the machined area. Here, the changes of the surface integrity are also lying on a maximum. In this case, the subsurface residual stresses caused by hard machining are of special meaning.

For investigations on the influence of hard turning to bending fatigue life, sample pieces with different residual stresses by the turning processes were selected. Stress cycle diagrams of a Wohler fatigue test reveal the influence of hard cutting conditions to the oscillating fatigue strength, see figure 5.18.

0 10’ 1 0’ 106 10’ source IW TU ciausma1 oscillation number N 31326M3 0 IFW

Figure 5.18: Residual stress in relation to the oscillation

The results of the investigations show that the residual stresses are almost not influenced by the number of oscillating loads. A reduction of residual stresses as shown in rolling contact loads does not occur.

number

On sample pieces turned with worn tools, significant differences in the fatigue life could not be found, see figure 5.19.

8 W

MPa d 0 U

600 - n

2 500

E ? -

4 w

300 1 0‘ 1 o5 l o 6 10’

Figure 5.19: Oscillation fatigue life of hard turned

In contrast to this, sample pieces produced with unworn tools which cause more compressive residual stresses in the subsurface show much higher values of oscillation fatigue strength. Here, the striking decrease of fatigue strength that characterizes the samples which have been turned with worn tool cannot be found. The examination of a relation to the surface roughness of the samples did not show significant influence on the fatigue strength of oscillating loads, in contrast to the rolling contact fatigue life.

As shown in figure 5.20, on characteristic fracture surfaces wedge shaped fissures occur on the sample pieces which show tensile residual stresses. The fracture shape of sample pieces characterized by compressive residual subsurface stresses are very homogeneous so that consequently premature failures can be avoided. For this reason, the requirements for hard cutting processes to achieve components with high oscillating fatigue strength have to be chosen in the way that compressive subsurface residual stresses are generated.

oscillation number N 31Y26651 ‘C IFM source IMAB TU Clauslhal

samples

Figure 5.20: Surfaces of fracture of hard turned Wohler- fatigue tested sample pieces

6 POTENTIAL OF HARD CUTTING

6.1 Economics Apart from further influencing criteria, the material removal rate is a most important economical aspect to evaluate the productivity of a cutting process. For finishing operations, additionally the generated surfaces are of high importance. Figure 6.1 shows the relations between machining conditions and material removal rate. In grinding, the defining quantity is the width related volume rase. For finishing, it might be less than Q,’ = 1 mm @ms, in roughing high values of more than Q,’ = 20 mm lmms can be reached. Feasible process removal rates are able to achieve values exceeding 200 mm3/mms. From these values, the derived material removal rates Q, for grinding are at least as high as in hard cutting. Besides, the volume related material

562

Keynote Papers

characteristic parameter

removal rate A,., is an important criterion concerning productivity. It may be implied from the calculation that in grinding the surface rate is superior to hard cutting. This points out that the grinding time is often shorter than the cutting time in turning [42].

a~ 5 . 20mm a = 0.05-0.3mm f ' = 0.05-0.2 mm 3 vc = 150dmin

" = 1 m i r n.

I arindina I cuttina I

typical values

a f " a =eL W I

= d- II v, spec material removal rate

a,, = 2 - 12 mm lqms = 22. 242 m p Imms F: Qw = 10.240 mm Is I 0 = 6.150 rnm Is1 8

= 5000-20000mm Is A: = 125-MOmm lo

surface rate A - = f v

0 gnnding 2 machines hydraulic component workptece matenal 16MnCr5 52-56 HRC

internal I external grinding caolant

hard turning 1 precision lathe one Selup

machined surface dry machining

'qdev@&vv@ ,@' &#* 0" 8' ,/@' 4%' &Q$ +.&> .#= a.6 &+ e\*

source Mannermann ReirOlh 31Y20526c Q IFW

Figure 6.2: Turning versus grinding - hydraulic

Compared to the former production by grinding operations, the production sequence is reduced noticeably. Only one machine tool and one single set up is necessary. Workpiece quality achieved in hard turning is at least at the level of the grinding operation. Significant improvements can be seen concerning production times and costs. For instance machining time for batch size is about 60% shorter compared to grinding. In comparison to hard turning, grinding is a more complex process. The repeated conditioning of the grinding wheel occupies the machine itself while tool preparation in hard turning can be done off-line.

component

An additional bonus of hard turning is avoiding cutting fluids. The possibility of dry machining means saving considerable costs otherwise caused by buying, monitoring, treatment and disposal of cutting fluids.

6.2 Environmental aspects The following example shows the ecological benefits of cutting compared to grinding, (figure 6.3). Assumed that 5000 parts of a gear component have to be produced per year, approximately 50 kg chips will result from the overmeasure machining. This is independent from the chosen manufacturing process.

However, in grinding a consumption of coolant has to be considered additionally which reaches up to 8 tons per year. Furthermore, degendent on grinding wheel composition, about 20 cm of abrasive and bond particles are generated during machining and dressing. These particles are mixed with coolant, chips and filter material. In industrial application, it is nearly impossible to separate these materials. Thus, the waste consists of a variety of different, usually detrimental to health and environmentally harmful waste materials, the whole amount has to be disposed under special security conditions.

\ - - - - - . 'i machined', , , mmpMent *

disp0le.I I m I I q m ' F y 8 ua

/'

1117291ciDIFW lW10

Figure 6.3: Hard turning versus grinding -ecological

Cutting leads to more favorable conditions. Due to the possibility of dry machining, there are only chips consisting of not contaminated workpiece material that can easily be recycled. Very small amounts of tool material are to be neglected because they are dissolving easily in the steel matrix. Tool can be disposed (ceramics) or reused after sharpening, but there is no mixture of different materials. Thus, cutting of hard materials can be considered as a very efficient possibility for protection of environment.

aspects

7 SUMMARY First of all, the availability of hard and super hard cutting tool materials enabled to machine hardened steels. Investigations in the cutting tool sector accompanied with developments and design of machine tools supports the successful story of this relatively new production technology.

Independent from cutting parameters high compressive stresses due to relatively high rake angles are adjusted in the tool tip region. High hydrostatic pressure is defined to be the most important physical quantity of plastic deformation of hardened ferrous materials which is essential for manufacturing technical surfaces [23]. By increasing the distance to the tool tip compressive stress is increased [25] and plastic deformation cannot take place. Shear localization occurs and results in segmented chip formation. Local thermal softness, crack initiation in the free workpiece surface and microcracks in the shear zone represent the most denunciated phenomena initiating shear localization [12] [26] [30] [31] [32] [34]. In literature, different theories exist concerning the crack growth. Crack initiation and crack growth are subject to temperature distribution, stress distribution, strain and strain rate in the work area, and on time and location. Depending on resulting time and location

563

dependent thermo-mechanical loads, crack initiation and crack growth is decisively influenced.

If the workpiece hardness is increased in a range up 50 HRC, the machining force and the energy consumption are increased [36] [39]. The increase of converted energy in the work area results in higher temperatures and influences the thermal load of the tool and of the workpiece subsurface [44] [46]. The passive force F is the largest cutting force component. With an increashg cutting length leading to an increasing width of flank wear, a significant rise of the passive force F can be observed. Consequently, the force ratio FJP,, is reversed in contrast to conventional cutting [42]. Newly developed high precision lathes offer the possibility to improve the quality of hard turned workpiece significantly. Turning tests on roller bearings with high demands on surface roughness and shape accuracy show the potential of precision hard turning. Due to the ability of high-precision lathes to move in steps less than 1 pm especially fine and smooth microprofiles can be realized which are necessary for some precision parts. Furthermore, compensation of form errors as a result of thermal expansion of the workpiece becomes possible.

Changes of boundary conditions require the consideration of ecological aspects in manufacturing. In order to promote the development of new manufacturing concepts, detailed analyses of single processes as well as whole production chains are necessary although in future, the acceptance of ecological manufacturing technologies will depend on their economical benefits. Using dry machining, the environmental pollution as well as the production costs can be decreased. For this, necessarily a consequent adaptation of tools and machines is required. Every application has to be checked for economical suitability.

A comparison between hard cutting and grinding shows that the first offers several important advantages. Different surfaces and shapes can be formed applying only a single tool, and the surface quality is as well comparable, as might be shown in finishing of a hydraulic component. Therefore, hard turning turns out to have high potential to replace grinding operations.

8 REFERENCES Chryssolouris, G.: Einsatz hochharter polykristalliner Schneidstoffe zum Drehen und Frasen, Dr.-lng. Dissertation, IFW-Hannover 1979 Koch, K. F.: Technologie des Hochprazisions- Hartdrehens, Dr.-lng. Dissertation, RWTH-Aachen

Brandt, D.: Randzonenbeeinflussung beim Hartdrehen, Dr.-lng. Dissertation, Universitat- Hannover, 1995 VDEh: Werkstoffkunde Stahl Band 1, Grundlagen, Springer Verlag, Berlin 1984 Schumann, H.: Metallographie, Deutscher Verlag fur Grundstoffindustrie, Leipzig 1989 Rose, A.: Eigenspannung als Ergebnis von Warmebehandlung und Umwandlungsverhalten,

Schatt, W.: Einfuhrung in die Werkstoffwissenschaft, VEB-Deutscher Verlag, Leipzig 1973 Macherauch, E., Voehringe, 0.: Verformungsverhalten geharteter Stahle. HTM 41

Razim, C.: Restaustenit zum Kenntnisstand uber Ursachen und Auswirkungen bei einsatzgeharteten Stahlen, HTM 40 (1985) 4, pp. 150-164 Habig, K.-H.: Verschleil3 und Harte von Werkstoffen, Carl Hanser Verlag, Munchen Wien. 1980

HTM 21 (1966) 1, pp. 150-164

(1986) 2, pp. 71-91

Liedke, D., Jonsson, R.: Warmebehandlung - Grundlagen und Anwendungen fur Eisenwerkstoffe. Expert Verlag, Ehningen, 1991 Komanduri, R., Brown, R. H.: On the Mechanics of Chip Segmentation in Machining, Journal of Engineering for Industry, Trans. ASME, 103 (1981);

Lemaire, J. C., Backofen, W. A.: Metall. Trans., 3

Ueda, N., Matsuo, T., Uehara, K.: An Analysis of Saw-Toothed Chip Formation, Ann. CIRP, 31

Toenshoff, H.K., Ben Amor, R., Andrea, P.: Chip Formation in High Speed Cutting (HSC), SME- Technical Paper, 3‘ International Machining an Grinding Conference, October 4-7, 1999, Cincinnati-Ohio, MR99-253 Koenig, W., Wand, Th.: Hartdrehen von Walzlagerstahl mit PKD und Schneidkeramik, IDR

Baik, Ch.-M.: Drehen von Stellite 6 mit Hartmetall und PKD, IDR 21 (1987) 2, pp. 95-99 Koenig, W., Iding, M., Link, R.: Feindrehen und Bohren geharteter Stahlwerkstoffe, IDR (1989) 1,

Meister, D., Theisen, W.: Einflul3 der Drehbearbeitung von Hartlegierungen auf die Oberflachenrandzone, IDR (1990) 3, pp. 143-148 Kaiser, M.: Hardfeinbearbeitung von einsatzgehartetem Stahl mit PKB-Werkzeugen, IDR

Biermann, D., Segtrop, K., Theisen, W.: Spanende Bearbeitung verschleitlbestandiger Hartlegierungen, VDI-Berichte 1021 (1993), pp.

D. Biermann, M. Liedschulte: Plasmaunterstutztes Drehen von Hartlegierungen auf Eisenbasis mit

Toenshoff, H. K., Wobker, H.-G., Brandt, D.: Hartbearbeitung aus der Sicht der Forschung, VDI- Berichte 988 (1993), pp. 189-209 Beitz, W., Kuttner, K.-H.: Dubbel “Taschenbuch fur den Maschinenbau” 16. Auflage, Springer Verlag, Berlin 1987 Fallboehmer, P.: Advanced Cutting Tools for the Finishing of Dies and Molds, Dr.-lng. Dissertation, Uni-Hannover, 1998 Wobker, H.-G.: Hartbearbeitung, Habilationsschrift Uni-Hannover, 1997 Komanduri, R., Zhen-Bin, H.: On a Thermo- mechanical Model of Shear Instability in Machining, Annals of the ClRP 44 (1995) 1, pp. 69-73 Recht, R. F.: Catastrophic Thermoplastic Shear, Trans ASME, 86 (1964), pp. 189-193 Semiatin, S. L., Rao, S. B.: Shear Localization during metal cutting, Material science and Engineering, 61 (1983), pp. 185-192 Recht, R. F.: A dynamic analysis of high speed machining, ASME Journal of Engineering For Industry, 107(1985), pp. 309-315 Nakayama, K., Arai, M., Kanda, T.: Machining Characteristic of Hardened Steels, Annals of the

Elbestawi, M. A., Srivastava, A. K., El-Wardany, T. I.: A Model For Chip Formation During Machining of Hardened Steel, Annals of the ClRP 45 (1996) 1,

pp. 33-51

(1972), pp. 477-481

(1982) 1, pp. 81-84

(1989) 2, pp. 67-74

pp. 22-33

(1 992) 1, pp. 24-33

509-51 2

PKD, IDR (1994) 2, pp. 71-77

CIRP, 37 (1988) 1, pp. 89-92

pp. 71-76

564

Keynote Papers

[33] Sih, G. C.: Strain-Energy-Density Factor Applied to Mixed Mode Crack Problems, Int. J. of Fracture

[34] Shaw, M. C., Vyas, A.: Mechanics of Saw-Tooth

10 (1974) 3, pp. 305-321

Chip Formation in Metal Cutting, ASME Journal of Manufacturing Science and Engineering, 121 (1999), pp. 163-172

[35] Shaw, M. C., Ramaraj, T. C., Santhanam, S.: Foot Formation, ASME Journal Of Engineering For Industry, 110 (1988), pp. 333-338

[36] Matsumoto, Y., Barash, M. M., Liu, C. R.: Cutting mechanisms during machining of hardened steels, Material science and technology, 3 (1987), pp. 299- 305

[37] Chao, B. T., Trigger, K. J.: Cutting Temperature and Metal-Cutting Phenomena, Trans. ASME (1951) 73, pp. 777-793

[38] Aspinwall, D. K., Ng, E. G.: Schnittkraft und Temperatur beim Hartdrehen von gehartetem Gesenkstahl, lndustrie Diamanten Rundschau,

[39] Eda, H., Kishi, K., Hashimoto, S.: In cutting tool materials, American society for metals, 1980, pp.

[40] Liu, C. R., Mittal, S.: Single-Step Superfinish Hard Machining: Feasability and feasible Cutting Conditions, Robotics and Computer-Integrated Manufacturing, 12 (1996) 1, pp. 15-27

[41] Toenshoff, H. K., Wobker, H.-G., Brandt, D.: Hard Turning - Substitute for Grinding Operations, Eurometalworking, Udine, September 28-30, 1994,

33 (1 999) 3, pp. 260-273

265-286

pp. 6011-6018

[42] Toenshoff, H. K.: Environmental and Ecological Aspects of Cutting -State, Developments, Potential- , I.C.E.M. Meeting, Ecology and Economy in Metal Forming and Cutting, s'-Hetogenbosch, Netherlands 1997

[43] Lindenberg, B., Lindstroem, B.: Measurements of the Segmentation Frequency in the Chip Formation Process, Annals of the CIRP, 32(1983)1, pp. 17-20

[44] Schmidt, J.: Mechanische und thermische Wirkungen beim Drehen geharteter Stahle, Dr.-lng. Dissertation Universitat Hannover, 1999

[45] Toenshoff, H.K.: Spanen, Grundlagen, Springer- Verlag, 1995

[46] Ueda, T., Al Huda, M., Yamada, K., Nakayama, K.: Temperature Measurement of CBN Tool in Turning of Hardened Steel: Annals of the CIRP, 48 (1999)

[47] Winands, N.: Hartdrehen aus der Umformwarme geharteter Walzlagerringe. Dr.-lng. Dissertation RWTH Aachen, 1996

[48] Koenig, W., Komanduri, R., Toenshoff, H.K., Ackerschott, G.: Machining of Hard Materials, Annals of the ClRP Vol. 3 (1984) 2, pp. 417-427

[49] Weinert, K., Albermann, F., Finke, M., Guntermann, G.: Prozekoptirnierung bei der Formenherstellung. Wt - Werkstattechnik 89 (1999) 718

[50] Enselmann, A.: HSC-Hartfrasen von Formen und Gesenken, Dr.-lng. Dissertation, Universitat Dortmund, 1998

[51] Wertheim, R.: Development and application of new cutting tool materials. ClRP General Assembly, Montreux, Switzerland, August 1999

[52] Boemcke, A.: Ein Beitrag zur Ermittlung der Verschleikmechanisrnen beim Zerspanen mit hochharten polykristallinen Schneidstoffen. Dr.-lng. Diss. RWTH Aachen. 1989

1, pp. 63-66

[53] Goldstein, M.: Optimierung der Fertigungsfolge ,Kalffliekpressen - Spanen' durch Hartdrehen als Feinbearbeitungsverfahren fur einsatzgehartete Prekteile. Dr.-lng. Diss. RWTH Aachen, 1991

[54] Steinmetz, K.: PKB-Wendeschneidplatten im Einsatz: Harte Eisenwerkstoffe wirtschaftlich zerspanen, lndustrie Anzeiger 106 (1984) 30, pp. 22-25

[55] Stanske, C.: Einsatz moderner Schneidstoffe fur die Hartbearbeitung. Werkzeuge fur die moderne Fertigung, Reihe Kontakt und Studium, Fertigung, Band 370 (1993) Ehningen: Expert, pp. 188-216, 223

[56] Chou, Y.-S.: Wear mechanisms of cubic boron nitride tools in precision turning of hardened steels. Diss. Purdue Univ., West Lafayette, USA, 1995

[57] Koenig, W., Neises, A.: Hartbearbeitung mit PKB - Auf den Schneidstoff kommt es an. lndustrie Diamanten Rundschau, Band 29 (1995) Heft 2,

[58] Kishawy, H.A., Elbastawi, M. A.: Effects of process parameters on material side flow during hard turning. Int. Journ. of Machine Tools & Manufacture

pp.108-I 17

39 (1999), pp. 1017-1030

[59] Werner, G.: Kinematik und Mechanik des

[60] Klimenko, S. A. et al.: On the Wear Mechanisms of

Schleifprozesses. Dr.-lng. Diss. Hannover, 1993

Cubic Boron Nitride Base Cutting Tool. Wear 157, 1992

Hardened Steel. Annals of the ClRP Vol. 42/1/1993

[62] Schaller, E.: Beitrag zur Untersuchung von Spannungen und dynamischen Vorgangen in der Grenzschicht zwischen Werkzeug und Span bei der Zerspanung mit Hartmetallwerkzeugen. Dr.-lng. Diss. RTWH Aachen 1964

[63] Toenshoff, H.K., Karpuschewski, B., Borbe, C.: Prazisions-Hgrtdrehen von Walzlagerringen. 12.-14. Mai 1997, 14 International Plansee Seminar '97, Reutte, Tirol

[64] Weck, M., Wieners, A.: Maschinen fur die Hoch- und Ultraprazision, Technica Nr. 13/14/95, S. 14-19

[65] Klocke, F., Koch, K. F.: High-precision Turning of Hardened Steel Part. Int. Prog. in Precision Engineering Proceedings 1995

[66] Martens, M., van Luttervelt, C. A.: Precisje- harddraaien in de kleinseriefabricage. MB Jg. 64, No. 617 1998

[67] Davies, M. A., Burns, T.J., Evans C. J.: On the Chip Morphology, Tool Wear and Cutting Mechanics in Finish Hard Turning. Annals of the ClRP Vol.

[68] Liu, C. R., Agha, S. R.: Experimental Study of the Pre-stressing Capability through residual Stresses in Superfinish machining of Hardened Steel. MED- Vol. 6-1, Manufacturing Science and Technology, Vol. 1 ASME 1997

[69] Liermann J.: Hartdrehen walzbelasteter Bauteile. Dr.-lng. Dissertation RWTH Aachen, 1997

[70] Toenshoff, H.K., Wobker, H.-G., Brandt, D.: Eigenspannungen und Randzonenausbildungen beim Hartd re hen. Harterei-Tec hnisc he Mitteil u ng en -HTM, Band 50 (1995) Heft 3, S. 176-181

[71] Fleming, M. A., Sweeney, M. A., Valentine, T. J., Simpkin, R.: Werkstuckrandzonenbeeinflussung beim Hartdrehen mit PKB. IDR 33 (1999) Nr.1

[72] Barash, M. M., Lui, C. R.: The Mechanical State of the Sublayer of a Surface Generated by Chip

[61) Shaw, M.C.: Chip Formation in the Machining of

pp.29-33

451111996 pp. 77-82

565

Removal Process. Journal of Engineering for industry, Vol. 98, No. 4, November 1976, pp. 1192- 1208.

[73] Toenshoff, H.K., Karpuschewski, B., Matzenell, C.: Sfrface generation by hard machining. Proceedings 4' European Conference on Resifual Stresses. Cluny en Bourgogne, France 4.-6' June, Volume 1

[74] Toenshoff, H. K., Karpuschewski, B., Borbe, C.: Hartbearbeitung: Stand der Forschung, Hard machining - State of Research, VDI Berichte Nr. 1399; 1998

[75] Brinksmeier, E: Anwendungsgerechte Hartbearbei- tung von Walzkomponenten. AWT-Tagung "Randschichtermudung im Walzkontakt" May 1992 Suhl.

[76] Brinksmeier, E., Kroning R.: Generation and Propertie? of Surfaces when Cutting Hardened Steels. 7' Int. Conf. on Metrology and Properties of Engineering Surfaces 1997, pp 45-52.

[77] Agha, S. R., Liu, R.: The Effect of cutting Conditions on Rolling Contact fatigue Life of Superfinish Hard Turned Surfaces. Proceedings of NSF and manufacturing Conference, Jan. 1999

Bauteilverhalten beim Hartdrehen, DGM-Seminar ,,Werkstoffgefuge und Zerspanung" Hannover 1999.

[79] Laub, H.: Hartdrehen ersetzt Schleifen. WB - Fertigen im Fahtzeugbau, 132 (1999) 9

[80] Grosspietsch, J.: Gewinde Hartschalen. Werkstatt und Betrieb, 131 (1998) 12, pp.1122-1126

[81] Momper, F.: Trocken und hart: Werkstatt und Betrieb, 131 (1998) 12

[78] Borbe, C.: Randzoneneigenschaften und

[82] Kwanka, W.: Hartdrehen von Hinterachswellen. VDI-Z 141 (1999) 9/10

[83] Brinksmeier, E., Kroning R.: Anwendung einschneidiger Reibwerkzeuge zur Hartfeinbearbeitung von Bohrungen. Werkstattstechnik wt 88 (1988) H 3 pp 139-141

566