- Email: [email protected]
Effects of surface-topography directionality and lubrication condition on frictional behaviour during plastic deformation
Journal of Materials Processing Technology 125–126 (2002) 379–386
Effects of surface-topography directionality and lubrication condition on frictional behaviour during plastic deformation W. Rasp, C.M. Wichern* Max-Planck-Institut fur Eisenforschung GmbH, Abteilung Mikrostrukturphysik und Umformtechnik, Max-Planck-Strasse 1, D-40237 Dussseldorf, Germany Received 20 December 2001; accepted 24 February 2002
Abstract Surface-topography, friction and plasticity form a complex system at the interface of a plastic deformation operation. Surface-topography before deformation affects the frictional behaviour at the interface, which in turn affects the level and distribution of plastic deformation. The surface-topography after deformation operations is determined by a combination of the initial surface-topography, the frictional behaviour at the interface and the level and distribution of the plastic deformation. Friction testing with the asymmetric friction upsetting (AFU) test machine and subsequent surface-topography analysis help to further clarify the relationship between surface-topography, interface friction and plasticity. The frictional conditions during testing are controlled via lubricant viscosity, film thickness, surface roughness, and deformation velocity. High and low viscosity lubricants are used in conjunction with a range of test speeds to produce frictional conditions in the boundary, mixed and hydrodynamic lubrication regimes. The effect of initial surface-topography was examined by preparing five different initial specimen surfaces and recording the surfacetopographies before testing. Five surface conditions were used: as received, etched, coarse ground perpendicular to test direction, coarse ground parallel to test direction and polished. A correlation between surface-topography directionality and frictional resistance has previously been observed by testing specimens with grooves machined into the surface at a range of angles [A newly developed test method for characterization of frictional conditions in metal forming, in: Proceedings of the Eighth International Conference on Metal Forming, Krakow, 2000, pp. 91–97; Steel Res. 69 (1998) (4–5) 154– 160; Beurteilung des Schmierungsverhaltens unterschiedlich texturierter Oberfla¨chen mit Hilfe des Streifenziehversuches, Diplomarbeit Universita¨t, Gesamthochschule Duisburg, 1996]. In the current work coarse grinding is used in place of machined grooves. The scale of the effect from the surface-topography directionality is compared to the scale of the effect of lubricants, arithmetic roughness value and friction regime. Results indicate that arithmetic roughness value and lubrication regime have greater influence than directionality. These results can be explained via the application of lubrication regime theory and the importance of each component in determining the lubrication condition. # 2002 Elsevier Science B.V. All rights reserved. Keywords: Friction; Plasticity; Surface-topography; Surface directionality
1. Introduction Surface-topography, friction and plasticity form a complex system at the interface of a plastic deformation operation. This interaction is particularly critical in rolling operations. Friction and lubrication behaviour in the roll gap are widely accepted to have a significant influence on thickness reduction, rolling efficiency and finished product surface appearance [2,4–7]. An understanding of the effect whose initial surface-topography has the lubrication regime *
Corresponding author. Tel.: þ49-211-6792-342; fax: þ49-211-6792-333. E-mail address: [email protected] (C.M. Wichern).
within the roll gap and the subsequent effect of the lubrication regime on the rolling operation and the final surface of the finished product is naturally desirable. In order to aid the understanding of roll gap behaviour mechanical simulations, which isolate individual aspects of the roll gap, must be employed. Subsequently, individual parameters within these individual aspects of the roll gap can be evaluated. Previously several studies of the effects of surface-topography directionality, otherwise known as surface lay, on frictional behaviour have been made at the Max Planck Institute for Steel Research. These studies have been performed using flat drawing tests, rolling tests and plane-strain upsetting tests. The flat drawing and upsetting tests simulate specific aspects of the roll gap while the rolling test provides
0924-0136/02/$ – see front matter # 2002 Elsevier Science B.V. All rights reserved. PII: S 0 9 2 4 - 0 1 3 6 ( 0 2 ) 0 0 3 4 6 - 1
380
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
an overall view of roll gap behaviour. In all of these tests surface lay perpendicular to the test specimen flow-direction produced the lowest friction for lubricated surfaces. Wolff et al. [1] studied the influence of surface lay on frictional behaviour using a flat drawing test. Lubricated specimens were tested with grooves machined into the surface at angles of 0 (parallel to drawing direction), 30, 45, 60 and 908. Friction was significantly lower at all speeds for the specimens with a surface lay perpendicular to the drawing direction. Another interesting result from Lordan’s work was the appearance of the machined grooves subsequent to testing. The grooves at angles 30, 45 and 608 showed decreasing levels of deformation, respectively. The decreased level of deformation is a result of the inability of the lubricant to flow out of the work zone of specimens with larger groove angles. The pressurized lubricant trapped in the grooves prevents the flow of material into grooves. Rasp and Ha¨ fele [2] studied the influence of surface lay on lubricant film thickness and workpiece reduction during cold rolling of sheets. Directional surfaces were cold rolled at a range of speeds between 0.33 and 2.83 m s1. The results showed that surfaces with a lay perpendicular to the rolling direction developed a thicker lubricant film leading to lower friction and greater thickness reduction. Velocity changes between 0.33 and 2.83 m s1 had very limited effects on the film thickness developed in the roll gap during testing of the directional surfaces. However, reduction of the workpiece during rolling increased with increasing velocity and lay angle. Lordan [3] studied the influence of surface lay on frictional behaviour using plane-strain upsetting. It was observed that surface lay perpendicular to the direction of material flow in the test specimen provided the lowest friction. The friction decreased continuously as the surface lay angle was changed from parallel to flow-direction to perpendicular to flow-direction. Additionally, Lordan [3] observed that surfaces with lay perpendicular to flow-direction retained their lay better than surfaces with lay parallel to flow-direction. The coefficient of friction, m, and the friction factor, m, provide a quantification of the frictional behaviour at the interface, however other means are necessary for qualifying and quantifying the surface-topography changes of the surfaces at the interface. Three-dimensional surface-topography analysis provides means of qualitatively describing a surface via surface contour plots. However, a quantitative method of surface description for the comparison of surfaces in an objective manner is also desirable. The three-dimensional arithmetic mean surface roughness is given by Stout et al. [8] as: Sa ¼
N X M 1 X Zðxi ; yj Þ MN j¼1 i¼1
coordinate (x, y), Sa the three-dimensional analogue of the standardized two-dimensional surface roughness [9]. Both Ra and Sa are subject to some variation resulting from data filtering and differing scan-area sizes. Whereas Sa describes only the vertical properties of the surface, other information is available which may provide more useful information about the surface with regard to lubrication and friction. Quantifying the volume below a given surface height, the number of isolated volumes (nonconnecting volumes) below a given surface height and the average size of isolated volumes below a given surface height is particularly useful for understanding the interaction between roll gap lubricant film thickness and surface roughness. The height at which the volume is calculated will have a large effect on the resulting volume parameters. Fig. 1 shows an illustration of a surface profile, which shows the effect of the height at which volume is calculated. Fig. 1(a) shows the original surface profile. In Fig. 1(b) a line has been drawn through at some given height, h, and the six volumes below h have been shaded in. When the height at which the volume is calculated is increased to h as in Fig. 1(c), the volume of interest changes significantly. Although there are still six individual volumes, the size and nature of the volumes are different. The total volume and the average volume have both increased. Additionally, some volumes that were isolated in Fig. 1(b) are now connected in Fig. 1(c). Specifically, volumes 2 and 3 in Fig. 1(b) have combined to create volume 3 in Fig. 1(c). Finally new volumes, specifically volume 1 in Fig. 1(c), are created as the height is increased. One way to ensure that volume parameters are calculated at a consistent height is to use some quantifiable height parameter. The Abbot curve offers a solution to this problem. Fig. 2 shows an example of the Abbot curve and the compensating gradient. The compensating gradient is a
(1)
where M and N are the number of points measured in the x- and y-directions respectively, Z(x, y) the surface height at
Fig. 1. Illustration of the effect of the height at which surface volumes are calculated: (a) surface profile as measured; (b) isolated volumes below height h; (c) isolated volumes below height h.
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
Fig. 2. Example of an Abbot curve showing the compensating gradients Z1 and Z2.
straight line of any slope, which is placed on the Abbot curve at the position where it gives the best fit over 40% of the data. The height at which the compensating gradient intersects the surface height axis at data-cut values of 0 and 100% are Z1 and Z2, respectively. The surface heights between Z1 and Z2 are the heights of the surface that bear most of the load during tribological contact. Thus in these areas the die and the workpiece are considered to be in intimate contact and any lubricant at the interface other than a monolayer will either be forced out of the interface or into the valleys below Z2. In accord with these assumptions the current work uses the volume below Z2 to describe surface void volume before and after testing.
2. Experimental method 2.1. Asymmetric friction upsetting (AFU) Lordan [3] have previously described the test apparatus used in this work in greater detail. The basic test methodology is identical to that of a traditional plane-strain upsetting test apparatus, which is often used to simulate the deformation that occurs in the roll gap [10]. The simulated rolling direction is the length direction, left to right in Fig. 3, and the simulated transverse direction is the width direction, into the page in Fig. 3. The test apparatus for the current work measures friction not via the traditional method of measuring the upset at a given force or the force necessary to obtain
381
a given upset [9], but instead via the angle formed in the test specimen when the upper and lower die–workpiece interfaces have different frictional conditions. This technique, referred to as the AFU which is illustrated in Fig. 3 where the upper die–workpiece interface has a higher coefficient of friction than the lower interface. The lower friction at bottom die–workpiece interface allows more lateral flow on the bottom of the specimen and an angle, a is formed. The angle a is used to calculate the difference in the friction number, Dm, between the top and bottom die–workpiece interfaces. The Dm value is calculated from the upset strain, j, and a via Eq. (2) as developed by Lordan [3] Dm ¼
2a 3j
(2)
2.2. Surface-topography All surface-topographies were measured using a commercial three-dimensional surface-profilometer. The field of view of the measurements was 1:54 mm 1:6 mm with a lateral resolution of 3 mm. The test specimen surface-topographies were measured in the same location before and after upsetting. Commercial software was used to produce surface-topography contour-plots. The three-dimensional arithmetic mean surface roughness (Sa), the total surface volume (VT), the average volume (VA), the number of volumes (NV), the Abbot curve, the compensating gradient, Z1 and Z2 were all calculated using FORTRAN programs developed by the authors. No filtering was performed on the surface data before the surface parameters were calculated. 2.3. Specimen preparation Test specimens were prepared from 99.5 aluminium with a yield stress s0 of 140 MPa. The specimens were 4 mm 30 mm 55 mm in size. The specimen surfaces were tested in the following forms: as received, polished to a mirror surface, chemically etched, abrasively scratched with 200 grit paper parallel and transverse to the simulated rolling direction. These surface conditions are henceforth referred to as AR, POL, ET, RD and TD, respectively. 2.4. Test matrix
Fig. 3. Schematic diagram of AFU test apparatus showing the angle, a, formed due to the frictional differences between the upper and lower die– workpiece interfaces.
The current work used the AFU in conjunction with a drop hammer for very high strain rates and in conjunction with a hydraulic press for lower strain rates. Upsetting velocities ranged from 6.4 to 0.0005 m s1, which correspond to strain rates of 1600 to 0.125 s1. Two lubricants of similar chemical makeup, but differing viscosity were used. Lubricant 1 has a viscosity of 8.8 mm2 s1 and lubricant 2 has a viscosity of 89.2 mm2 s1. Testing was performed in the boundary, mixed and hydrodynamic lubrication regimes by using different combinations
382
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
of upsetting speed and lubricant viscosity. Different combinations of upper and lower test specimen surface preparations were used to allow direct and indirect comparison of the surface-topography changes and frictional behaviour.
3. Results Key results of the surface-topography analysis from three test specimens are given in Figs. 4–6. In all surface contourplots the simulated rolling direction runs laterally across the page and the simulated transverse direction runs vertically across the page. Surface contours for a test specimen with TD and RD surfaces on the opposing sides before and after upsetting with no lubricant on either side are shown in Fig. 4. The untested surfaces for both sides clearly show the scratches and the high directionality of the surfaces (Fig. 4(a) and (c)). After testing the TD surface retains very little of its directional feature and is subject to large-scale smoothing (Fig. 4(b)). The extent of the smoothing is also seen in the decrease of Sa from 1.67 to 0.45 mm (Table 1). After testing the RD surface, Fig. 4(d) also shows large-scale smoothing and an Sa decrease from 1.93 to 0.33 mm. The RD surface, however, shows more retained
directionality and also develops previously non-existent features transverse to the simulated rolling direction. The AR surface shows nearly no change in Sa as a result of upsetting without lubricant (Table 1). The second group of results (Fig. 5) provides insight into the effect of lubricant on surface-topography changes during upsetting without the influence of the initial surface-topography. A test specimen with a polished surface (Fig. 5(a)) shows severe roughening as a result of upsetting (Fig. 5(b)). The Sa value of the surface increases from 0.13 mm before upsetting to 1.53 mm after upsetting. The volume parameters for this surface also change significantly during upsetting. Originally the surface in Fig. 5(a) had a large number, NV ¼ 1417 of very small, VA ¼ 1 mm3, isolated volumes. After upsetting the surface has fewer, NV ¼ 484, but larger, VA ¼ 2866 mm3, isolated volumes. The third group of results provides insight into the effects of both lubrication and initial surface-topography on the surface-topography changes during upsetting. Fig. 6 shows a test specimen with TD and RD surfaces on the opposing sides before and after upsetting using lubricant 2 on both sides. Again, the initial surface directionality is obvious in Figs. 6(a) and (c). After upsetting the TD surface retains its directionality, but the morphology of the scratches is dif-
Fig. 4. Surface contour-plots of test specimens before and after upsetting without lubrication at an upset speed of 4.5 m s1: (a) untested surface with transverse direction scratches; (b) surface a after testing; (c) untested surface with rolling direction scratches; (d) surface c after testing.
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
383
Fig. 5. Surface contour-plots of test specimens before and after upsetting with lubricant 2 at an upset speed of 6.3 m s1: (a) untested polished surface; (b) surface a after testing.
ferent. The scratches are now fewer, but deeper and wider. In this case Sa increases, from 1.57 to 3.54 mm, similarly to the behaviour observed with the polished surface. Additionally, NV decreases from 658 to 251 and VA increases from 3010 to 11 444 mm3. These trends were also observed on the polished surfaces before and after upsetting.
After upsetting the RD surface (Fig. 6(d)) loses some of its initial surface directionality and actually develops transverse direction surface-features after upsetting. The RD surface shows less roughening than the TD surface, with Sa increasing from 1.93 to 2.03 mm. The volume parameters of the RD surface also behave differently with NV and VA both decreas-
Fig. 6. Surface contour-plots of test specimens before and after upsetting with lubricant 2 at an upset speed of 4.5 m s1: (a) untested surface with transverse direction scratches; (b) surface a after testing; (c) untested surface with rolling direction scratches; (d) surface c after testing.
384
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
Table 1 Summary of all of the surface parameters reported for all surface treatment and upsetting conditions Velocity (m s1)
Viscosity (mm2 s1)
VT
NV (mm3)
As received/RD Untested Test 1 Test 2
NA 4.5/6.3 4.5
NA None 89.2
5.42E þ 04 3.38E þ 04 3.93E þ 05
3433 2365 873
Polished/POL Untested Test 1 Test 2
NA 6.3 6.3
NA 89.2 8.8
1.31E þ 03 4.09E þ 06 1.39E þ 06
Etched/ET Untested Test 1
NA 6.3
NA 89.2
1.55E þ 06 2.40E þ 06
Rolling direction Untested Test 1 Test 2 Test 3 Test 4
scratches/RD NA 0.001 4.5 6.3 0.001
NA None 89.2 89.2 89.2
1.38E 8.31E 1.13E 3.30E 8.61E
þ þ þ þ þ
NA None 89.2 89.2 89.2
1.46E 1.21E 2.91E 5.34E 4.28E
þ þ þ þ þ
Transverse direction scratches/TD Untested NA Test 1 0.001/4.5 Test 2 4.5 Test 3 6.3 Test 4 0.001
ing, from 595 to 474 mm3 and from 2788 to 2562 mm3, respectively. Another effect of the surfaces changes observed in Figs. 4–6 is a change in the size and nature of the intimate contact area between the die and the workpiece, which is observed in the Abbot bearing area curve. Fig. 7 shows the Abbot curves for the test specimen shown in Fig. 4. The TD surface (Fig. 7(a)), shows a numerical increase in the compensating gradient slope, mGC from 0.05 to 0.01. Additionally, the Abbot curve for the tested surface is ‘‘flatter’’ in that the linear portion of the curve is larger—indicative of greater intimate contact. The RD surface (Fig. 7(b)) shows a similar flatness and a similar mGC increase from 0.06 to 0.01. The Abbot curves for the specimen with the polished surface (Fig. 8) behave differently in that mGC decreases
VA (mm3)
Z2 (mm)
mGC (mm)
Sa (mm)
16 14 451
0.56 0.57 0.98
0.01 0.01 0.03
0.35 0.34 0.98
1417 354 484
1 11560 2866
0.31 0.83 0.49
0.00 0.07 0.05
0.13 2.57 1.53
280 194
5552 12350
2.93 2.81
0.11 0.12
3.52 4.11
06 04 06 06 04
595 2850 474 236 3042
2788 29 2562 14123 29
1.19 0.48 1.55 3.69 0.53
0.06 0.01 0.06 0.15 0.01
1.93 0.33 2.03 5.30 0.37
06 05 06 06 05
658 2036 251 170 2205
3010 213 11444 31430 60
0.67 0.61 1.87 2.64 0.41
0.05 0.01 0.11 0.16 0.01
1.67 0.45 3.54 6.23 0.58
from 0.00 to 0.07 and the Abbot curve becomes less flat after upsetting using lubricant 2. The Abbot curves resulting from the upsetting of RD and TD surfaces using lubricant 2 are shown in Fig. 9. Upon upsetting, the TD surface behaves similarly to the polished surface showing a decrease in mGC from 0.00 to 0.11 and a loss of flatness in the Abbot curve (Fig. 9(a)). The RD surface, however, is nearly identical before and after upsetting (Fig. 9(b)). The mGC for the RD surface is constant at 0.06 and the flatness of the Abbot curves are the same. The effects of upsetting on the surface-topography of the AR and ET surfaces are similar to the effects observed on the TD surface. Upsetting at a speed of 4.5 m s1 using lubricant 2 results in an increase in Sa and a decrease in mGC for both the AR and ET surfaces (Table 1). Also similarly to TD
Fig. 7. Abbot bearing area plots of test specimens before and after upsetting without lubrication at an upset speed of 4.5 m s1: (a) untested and tested surfaces with transverse direction scratches; (b) untested and tested surfaces with rolling direction scratches.
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
385
difference between the RD and TD surfaces, Dm ¼ 0. However, at velocities above and below 4.5 m s1 the TD surface has a lower m value for both the high and the low viscosity lubricant.
4. Discussion
Fig. 8. Abbot bearing area plots of test specimens before and after upsetting with lubricant 2 at an upset speed of 6.3 m s1.
Table 2 Dm values for different upset conditions for specimens with RD and TD surfaces on opposing sides Velocity (m s1)
Surface pair
Dm
No lube both sides 4.5 0.001
RD/TD RD/TD
0.046 0.077
Lube 2 both sides 4.5 0.001 0.0005
RD/TD RD/TD RD/TD
0 0.015 0.005
Lube 1 both sides 6.3 4.5 0.001
RD/TD RD/TD RD/TD
0.013 0 0.016
surfaces, the AR and ET surfaces both exhibit a decrease in NV and increase in VA after upsetting. The Dm values for the RD and TD surface combinations tested under different upset conditions are given in Table 2. In Table 2 Dm is reported as: Dm ¼ mRD mTD
(3) 1
Without lubrication at low upset velocity, 0.001 m s , the RD surface has a lower m value, Dm ¼ 0:077. At higher upset velocity, 4.5 m s1, the TD surface has a lower m value, Dm ¼ 0:046. For upsetting at 4.5 m s1 neither high nor the low viscosity lubricant result in any frictional
The surfaces upset without lubrication (Fig. 4 and Table 1) are nearly identical in terms of mGC and Sa independent of speed or initial surface preparation. The differences between the resulting RD and TD surfaces are limited to slight differences in VT and VA and the appearance of new surface-features at 908 to the original scratches on the RD surface after testing (Fig. 4(d)). These features found in Fig. 4(d) may be imprints from the die surface or effects from the creation of new free surface during upsetting. Nonetheless, it is clear that the resulting surfaces have very low roughness, have very low surface void volume and offer large areas of intimate contact between die and workpiece. At very low upsetting speeds, 0.001 m s1, the lubricant has more time to flow out of the interface in spite of the surface lay differences. As a result a large area of intimate contact is formed and the differences between the RD and TD surfaces after upsetting is very small as seen in Table 1. The lack of lubricant load bearing ability but the potential for lubricant presence and thus chemical action at the interface is indicative of the boundary lubrication regime. The opposite extreme of a large intimate contact area is complete separation of the die and workpiece, which is the case for the specimens with polished surfaces upset at high upsetting speed with high viscosity lubricant (Fig. 5). Under these upset conditions the lubricant is unable to escape from the die–workpiece interface and a lubricant cushion prevents intimate contact between the die and workpiece. In this case the surface roughens due to either pressure differentials within the lubricant at the interface or roughness-scale local in-homogeneties in the deformability of the surface. This combination of conditions, complete load bearing by the lubricant and surface roughening, indicates that the specimen in Fig. 5 was upset in the hydrodynamic lubrication regime.
Fig. 9. Abbot bearing area plots of test specimens before and after upsetting with lubricant 2 at an upset speed of 4.5 m s1: (a) untested and tested surfaces with transverse direction scratches; (b) untested and tested surfaces with rolling direction scratches.
386
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
As stated by Rasp and co-workers [1] there is a competition between the hydrodynamic lubrication effects and the surface-topography effects. When surface lay and roughness effects are combined with lubrication regime effects at high upsetting velocity, the effects of the mixed, hydrostatic microhydrodynamic [11] and fully hydrodynamic lubrication interact. The result of this interaction is a wide variation of frictional behaviour and surface-topography changes depending on the initial surface-topography of the test specimen. For example, under the above conditions the TD surface roughens significantly more and shows a much greater change in mGC than the RD surface. The reason for this disparity is the same one proposed by previous authors [1–4]: the lubricant is unable to escape from the interface as easily on the TD surface. The high roughness surfacefeatures present on the RD and TD surfaces break through the lubricant film and contact the die surface, an effect recognizable via the flattened peaks (Fig. 6). The scratches on the RD surface provide isolated, but and short channels for the lubricant to flow to the edge of the interface. The scratches on the TD surface, however, provide longer isolated channels making it more difficult for lubricant to exit the interface resulting in more lubricant at the interface and higher pressure in the lubricant. The RD and TD surfaces are in two different lubrication regimes. The interaction of the lubrication regime effects and the surface-topography effects complicates the interpretation of the frictional results used to compare the different surfaces. As can be seen in Table 2 at high speeds, 6.3 m s1, and low speeds, 0.001 m s1, the TD surface has a lower m value when lubricant is used. When lubricant is not used the RD surface has a lower m value at low speed and the TD has a lower m value at high speed. Further complicating the situation is the result that the TD and RD surfaces perform equally during lubricated upset at 4.5 m s1. However, careful consideration and rationalization of the tribological conditions at the interfaces provides a possible explanation. The tribological interaction, which occurs at an upset velocity of 4.5 m s1 without lubricant, is purely mechanical resulting in a lower m value for the RD surface. The RD surfaces after testing with and without lubricant, Figs. 6(d) and 4(d)) appear very similar implying that despite the presence of lubricant the tribological behaviour with and without lubricant is also similar. It has already been proposed that the TD surfaces provide better conditions for the hydrodynamic, mixed or hydrostatic micro-hydrodynamic lubrication regimes to occur. In the hydrodynamic lubrication regime that prevails during upsetting at 6.3 m s1 with lubricant 2 the TD surface has a lower m value. During testing at 4.5 m s1 the RD surface is still predominantly controlled by mechanical tribological behaviour where it has a lower m than the TD surface. The TD surface under the same upset conditions is closer to the hydrodynamic lubrication regime where it has a lower m than the RD surface and can thus match the frictional performance of the RD surface. This balance of opposing frictional regimes nullifies the
effects of the surface lay. At lower upset velocities this balancing effect may not be present or observable due to the dependency of boundary lubrication regime behaviour on contact time and velocity at the interface. 5. Summary The change in the slope of the compensating gradient, mGC, is a good gauge for determining whether a tribological interface is moving toward or away from the hydrodynamic lubrication regime. Volume parameters, which measure the volume of the voids on a surface at a specific height, can be used in conjunction with mGC to describe the quantity, location and effect of a lubricant at an interface during forming operations. Surface lay transverse to the rolling direction promotes lubricant entrapment at the interface and promotes roughening and surface void formation. The effects of surface lay on frictional behaviour during upset without lubricant vary with upset velocity. At high upset velocity the TD surface has a lower m value than the RD surface and at low upset velocity the RD surface has a lower m value than the TD surface. As a result of their mutual dependency lubrication regimes and surface-topography must be carefully considered when comparing the tribological behaviour of different surfaces. References [1] C. Wolff, O. Pawelski, W. Rasp, A newly developed test method for characterization of frictional conditions in metal forming, in: Proceedings of the Eighth International Conference on Metal Forming, Krakow, 2000, pp. 91–97. [2] W. Rasp, P. Ha¨ fele, Investigation into tribology of cold strip rolling, Steel Res. 69 (4–5) (1998) 154–160. [3] J. Lordan, Beurteilung des Schmierungsverhaltens unterschiedlich texturierter Oberfla¨ chen mit Hilfe des Streifenziehversuches, Diplomarbeit Universita¨ t, Gesamthochschule Duisburg, 1996. [4] O. Pawelski, W. Rasp, S. Draese, P. Ha¨ fele, Influence of hydrodynamic effects and surface roughness on tribological phenomena in cold strip rolling, in: Proceedings of the Fifth ICTP, Columbus, OH, 1996, pp. 31–34. [5] J.V. Reid, J.A. Schey, Full fluid film lubrication in aluminium strip rolling, ASLE Trans. 21 (3) (1978) 191–200. [6] W.R.D. Wilson, W. Lee, Mechanics of surface roughening in metal forming processes, Trans. ASME, J. Manuf. Sci. Eng. 123 (2001) 279–283. [7] W.R.D. Wilson, Friction and lubrication in bulk metal-forming processes, J. Appl. Metalworking 1 (1979) 7–19. [8] K.J. Stout, P.J. Sullivan, W.P. Dong, E. Mainsah, N. Luo, T. Mathia, H. Zahouani, The Development of Methods for the Characterization of Roughness in Three Dimensions, EUR 15178, ECSC-EEC-EAEC, Brussels-Luxemburg, 1993. [9] ISO 4287, Geometrical product specifications (GPS)—surface texture, Profile method—terms, definitions and surface texture parameters, 1997. [10] J.A. Schey, Tribology in Metalworking—Friction, Lubrication and Wear, American Society for Metals, Cleveland, OH, 1984. [11] H. Kudo, A. Azushima, Interaction of surface microstructure and lubricant in metal forming tribology, in: Proceedings of the Second ICTP, Stuttgart, 1987, pp. 373–384.
Effects of surface-topography directionality and lubrication condition on frictional behaviour during plastic deformation W. Rasp, C.M. Wichern* Max-Planck-Institut fur Eisenforschung GmbH, Abteilung Mikrostrukturphysik und Umformtechnik, Max-Planck-Strasse 1, D-40237 Dussseldorf, Germany Received 20 December 2001; accepted 24 February 2002
Abstract Surface-topography, friction and plasticity form a complex system at the interface of a plastic deformation operation. Surface-topography before deformation affects the frictional behaviour at the interface, which in turn affects the level and distribution of plastic deformation. The surface-topography after deformation operations is determined by a combination of the initial surface-topography, the frictional behaviour at the interface and the level and distribution of the plastic deformation. Friction testing with the asymmetric friction upsetting (AFU) test machine and subsequent surface-topography analysis help to further clarify the relationship between surface-topography, interface friction and plasticity. The frictional conditions during testing are controlled via lubricant viscosity, film thickness, surface roughness, and deformation velocity. High and low viscosity lubricants are used in conjunction with a range of test speeds to produce frictional conditions in the boundary, mixed and hydrodynamic lubrication regimes. The effect of initial surface-topography was examined by preparing five different initial specimen surfaces and recording the surfacetopographies before testing. Five surface conditions were used: as received, etched, coarse ground perpendicular to test direction, coarse ground parallel to test direction and polished. A correlation between surface-topography directionality and frictional resistance has previously been observed by testing specimens with grooves machined into the surface at a range of angles [A newly developed test method for characterization of frictional conditions in metal forming, in: Proceedings of the Eighth International Conference on Metal Forming, Krakow, 2000, pp. 91–97; Steel Res. 69 (1998) (4–5) 154– 160; Beurteilung des Schmierungsverhaltens unterschiedlich texturierter Oberfla¨chen mit Hilfe des Streifenziehversuches, Diplomarbeit Universita¨t, Gesamthochschule Duisburg, 1996]. In the current work coarse grinding is used in place of machined grooves. The scale of the effect from the surface-topography directionality is compared to the scale of the effect of lubricants, arithmetic roughness value and friction regime. Results indicate that arithmetic roughness value and lubrication regime have greater influence than directionality. These results can be explained via the application of lubrication regime theory and the importance of each component in determining the lubrication condition. # 2002 Elsevier Science B.V. All rights reserved. Keywords: Friction; Plasticity; Surface-topography; Surface directionality
1. Introduction Surface-topography, friction and plasticity form a complex system at the interface of a plastic deformation operation. This interaction is particularly critical in rolling operations. Friction and lubrication behaviour in the roll gap are widely accepted to have a significant influence on thickness reduction, rolling efficiency and finished product surface appearance [2,4–7]. An understanding of the effect whose initial surface-topography has the lubrication regime *
Corresponding author. Tel.: þ49-211-6792-342; fax: þ49-211-6792-333. E-mail address: [email protected] (C.M. Wichern).
within the roll gap and the subsequent effect of the lubrication regime on the rolling operation and the final surface of the finished product is naturally desirable. In order to aid the understanding of roll gap behaviour mechanical simulations, which isolate individual aspects of the roll gap, must be employed. Subsequently, individual parameters within these individual aspects of the roll gap can be evaluated. Previously several studies of the effects of surface-topography directionality, otherwise known as surface lay, on frictional behaviour have been made at the Max Planck Institute for Steel Research. These studies have been performed using flat drawing tests, rolling tests and plane-strain upsetting tests. The flat drawing and upsetting tests simulate specific aspects of the roll gap while the rolling test provides
0924-0136/02/$ – see front matter # 2002 Elsevier Science B.V. All rights reserved. PII: S 0 9 2 4 - 0 1 3 6 ( 0 2 ) 0 0 3 4 6 - 1
380
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
an overall view of roll gap behaviour. In all of these tests surface lay perpendicular to the test specimen flow-direction produced the lowest friction for lubricated surfaces. Wolff et al. [1] studied the influence of surface lay on frictional behaviour using a flat drawing test. Lubricated specimens were tested with grooves machined into the surface at angles of 0 (parallel to drawing direction), 30, 45, 60 and 908. Friction was significantly lower at all speeds for the specimens with a surface lay perpendicular to the drawing direction. Another interesting result from Lordan’s work was the appearance of the machined grooves subsequent to testing. The grooves at angles 30, 45 and 608 showed decreasing levels of deformation, respectively. The decreased level of deformation is a result of the inability of the lubricant to flow out of the work zone of specimens with larger groove angles. The pressurized lubricant trapped in the grooves prevents the flow of material into grooves. Rasp and Ha¨ fele [2] studied the influence of surface lay on lubricant film thickness and workpiece reduction during cold rolling of sheets. Directional surfaces were cold rolled at a range of speeds between 0.33 and 2.83 m s1. The results showed that surfaces with a lay perpendicular to the rolling direction developed a thicker lubricant film leading to lower friction and greater thickness reduction. Velocity changes between 0.33 and 2.83 m s1 had very limited effects on the film thickness developed in the roll gap during testing of the directional surfaces. However, reduction of the workpiece during rolling increased with increasing velocity and lay angle. Lordan [3] studied the influence of surface lay on frictional behaviour using plane-strain upsetting. It was observed that surface lay perpendicular to the direction of material flow in the test specimen provided the lowest friction. The friction decreased continuously as the surface lay angle was changed from parallel to flow-direction to perpendicular to flow-direction. Additionally, Lordan [3] observed that surfaces with lay perpendicular to flow-direction retained their lay better than surfaces with lay parallel to flow-direction. The coefficient of friction, m, and the friction factor, m, provide a quantification of the frictional behaviour at the interface, however other means are necessary for qualifying and quantifying the surface-topography changes of the surfaces at the interface. Three-dimensional surface-topography analysis provides means of qualitatively describing a surface via surface contour plots. However, a quantitative method of surface description for the comparison of surfaces in an objective manner is also desirable. The three-dimensional arithmetic mean surface roughness is given by Stout et al. [8] as: Sa ¼
N X M 1 X Zðxi ; yj Þ MN j¼1 i¼1
coordinate (x, y), Sa the three-dimensional analogue of the standardized two-dimensional surface roughness [9]. Both Ra and Sa are subject to some variation resulting from data filtering and differing scan-area sizes. Whereas Sa describes only the vertical properties of the surface, other information is available which may provide more useful information about the surface with regard to lubrication and friction. Quantifying the volume below a given surface height, the number of isolated volumes (nonconnecting volumes) below a given surface height and the average size of isolated volumes below a given surface height is particularly useful for understanding the interaction between roll gap lubricant film thickness and surface roughness. The height at which the volume is calculated will have a large effect on the resulting volume parameters. Fig. 1 shows an illustration of a surface profile, which shows the effect of the height at which volume is calculated. Fig. 1(a) shows the original surface profile. In Fig. 1(b) a line has been drawn through at some given height, h, and the six volumes below h have been shaded in. When the height at which the volume is calculated is increased to h as in Fig. 1(c), the volume of interest changes significantly. Although there are still six individual volumes, the size and nature of the volumes are different. The total volume and the average volume have both increased. Additionally, some volumes that were isolated in Fig. 1(b) are now connected in Fig. 1(c). Specifically, volumes 2 and 3 in Fig. 1(b) have combined to create volume 3 in Fig. 1(c). Finally new volumes, specifically volume 1 in Fig. 1(c), are created as the height is increased. One way to ensure that volume parameters are calculated at a consistent height is to use some quantifiable height parameter. The Abbot curve offers a solution to this problem. Fig. 2 shows an example of the Abbot curve and the compensating gradient. The compensating gradient is a
(1)
where M and N are the number of points measured in the x- and y-directions respectively, Z(x, y) the surface height at
Fig. 1. Illustration of the effect of the height at which surface volumes are calculated: (a) surface profile as measured; (b) isolated volumes below height h; (c) isolated volumes below height h.
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
Fig. 2. Example of an Abbot curve showing the compensating gradients Z1 and Z2.
straight line of any slope, which is placed on the Abbot curve at the position where it gives the best fit over 40% of the data. The height at which the compensating gradient intersects the surface height axis at data-cut values of 0 and 100% are Z1 and Z2, respectively. The surface heights between Z1 and Z2 are the heights of the surface that bear most of the load during tribological contact. Thus in these areas the die and the workpiece are considered to be in intimate contact and any lubricant at the interface other than a monolayer will either be forced out of the interface or into the valleys below Z2. In accord with these assumptions the current work uses the volume below Z2 to describe surface void volume before and after testing.
2. Experimental method 2.1. Asymmetric friction upsetting (AFU) Lordan [3] have previously described the test apparatus used in this work in greater detail. The basic test methodology is identical to that of a traditional plane-strain upsetting test apparatus, which is often used to simulate the deformation that occurs in the roll gap [10]. The simulated rolling direction is the length direction, left to right in Fig. 3, and the simulated transverse direction is the width direction, into the page in Fig. 3. The test apparatus for the current work measures friction not via the traditional method of measuring the upset at a given force or the force necessary to obtain
381
a given upset [9], but instead via the angle formed in the test specimen when the upper and lower die–workpiece interfaces have different frictional conditions. This technique, referred to as the AFU which is illustrated in Fig. 3 where the upper die–workpiece interface has a higher coefficient of friction than the lower interface. The lower friction at bottom die–workpiece interface allows more lateral flow on the bottom of the specimen and an angle, a is formed. The angle a is used to calculate the difference in the friction number, Dm, between the top and bottom die–workpiece interfaces. The Dm value is calculated from the upset strain, j, and a via Eq. (2) as developed by Lordan [3] Dm ¼
2a 3j
(2)
2.2. Surface-topography All surface-topographies were measured using a commercial three-dimensional surface-profilometer. The field of view of the measurements was 1:54 mm 1:6 mm with a lateral resolution of 3 mm. The test specimen surface-topographies were measured in the same location before and after upsetting. Commercial software was used to produce surface-topography contour-plots. The three-dimensional arithmetic mean surface roughness (Sa), the total surface volume (VT), the average volume (VA), the number of volumes (NV), the Abbot curve, the compensating gradient, Z1 and Z2 were all calculated using FORTRAN programs developed by the authors. No filtering was performed on the surface data before the surface parameters were calculated. 2.3. Specimen preparation Test specimens were prepared from 99.5 aluminium with a yield stress s0 of 140 MPa. The specimens were 4 mm 30 mm 55 mm in size. The specimen surfaces were tested in the following forms: as received, polished to a mirror surface, chemically etched, abrasively scratched with 200 grit paper parallel and transverse to the simulated rolling direction. These surface conditions are henceforth referred to as AR, POL, ET, RD and TD, respectively. 2.4. Test matrix
Fig. 3. Schematic diagram of AFU test apparatus showing the angle, a, formed due to the frictional differences between the upper and lower die– workpiece interfaces.
The current work used the AFU in conjunction with a drop hammer for very high strain rates and in conjunction with a hydraulic press for lower strain rates. Upsetting velocities ranged from 6.4 to 0.0005 m s1, which correspond to strain rates of 1600 to 0.125 s1. Two lubricants of similar chemical makeup, but differing viscosity were used. Lubricant 1 has a viscosity of 8.8 mm2 s1 and lubricant 2 has a viscosity of 89.2 mm2 s1. Testing was performed in the boundary, mixed and hydrodynamic lubrication regimes by using different combinations
382
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
of upsetting speed and lubricant viscosity. Different combinations of upper and lower test specimen surface preparations were used to allow direct and indirect comparison of the surface-topography changes and frictional behaviour.
3. Results Key results of the surface-topography analysis from three test specimens are given in Figs. 4–6. In all surface contourplots the simulated rolling direction runs laterally across the page and the simulated transverse direction runs vertically across the page. Surface contours for a test specimen with TD and RD surfaces on the opposing sides before and after upsetting with no lubricant on either side are shown in Fig. 4. The untested surfaces for both sides clearly show the scratches and the high directionality of the surfaces (Fig. 4(a) and (c)). After testing the TD surface retains very little of its directional feature and is subject to large-scale smoothing (Fig. 4(b)). The extent of the smoothing is also seen in the decrease of Sa from 1.67 to 0.45 mm (Table 1). After testing the RD surface, Fig. 4(d) also shows large-scale smoothing and an Sa decrease from 1.93 to 0.33 mm. The RD surface, however, shows more retained
directionality and also develops previously non-existent features transverse to the simulated rolling direction. The AR surface shows nearly no change in Sa as a result of upsetting without lubricant (Table 1). The second group of results (Fig. 5) provides insight into the effect of lubricant on surface-topography changes during upsetting without the influence of the initial surface-topography. A test specimen with a polished surface (Fig. 5(a)) shows severe roughening as a result of upsetting (Fig. 5(b)). The Sa value of the surface increases from 0.13 mm before upsetting to 1.53 mm after upsetting. The volume parameters for this surface also change significantly during upsetting. Originally the surface in Fig. 5(a) had a large number, NV ¼ 1417 of very small, VA ¼ 1 mm3, isolated volumes. After upsetting the surface has fewer, NV ¼ 484, but larger, VA ¼ 2866 mm3, isolated volumes. The third group of results provides insight into the effects of both lubrication and initial surface-topography on the surface-topography changes during upsetting. Fig. 6 shows a test specimen with TD and RD surfaces on the opposing sides before and after upsetting using lubricant 2 on both sides. Again, the initial surface directionality is obvious in Figs. 6(a) and (c). After upsetting the TD surface retains its directionality, but the morphology of the scratches is dif-
Fig. 4. Surface contour-plots of test specimens before and after upsetting without lubrication at an upset speed of 4.5 m s1: (a) untested surface with transverse direction scratches; (b) surface a after testing; (c) untested surface with rolling direction scratches; (d) surface c after testing.
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
383
Fig. 5. Surface contour-plots of test specimens before and after upsetting with lubricant 2 at an upset speed of 6.3 m s1: (a) untested polished surface; (b) surface a after testing.
ferent. The scratches are now fewer, but deeper and wider. In this case Sa increases, from 1.57 to 3.54 mm, similarly to the behaviour observed with the polished surface. Additionally, NV decreases from 658 to 251 and VA increases from 3010 to 11 444 mm3. These trends were also observed on the polished surfaces before and after upsetting.
After upsetting the RD surface (Fig. 6(d)) loses some of its initial surface directionality and actually develops transverse direction surface-features after upsetting. The RD surface shows less roughening than the TD surface, with Sa increasing from 1.93 to 2.03 mm. The volume parameters of the RD surface also behave differently with NV and VA both decreas-
Fig. 6. Surface contour-plots of test specimens before and after upsetting with lubricant 2 at an upset speed of 4.5 m s1: (a) untested surface with transverse direction scratches; (b) surface a after testing; (c) untested surface with rolling direction scratches; (d) surface c after testing.
384
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
Table 1 Summary of all of the surface parameters reported for all surface treatment and upsetting conditions Velocity (m s1)
Viscosity (mm2 s1)
VT
NV (mm3)
As received/RD Untested Test 1 Test 2
NA 4.5/6.3 4.5
NA None 89.2
5.42E þ 04 3.38E þ 04 3.93E þ 05
3433 2365 873
Polished/POL Untested Test 1 Test 2
NA 6.3 6.3
NA 89.2 8.8
1.31E þ 03 4.09E þ 06 1.39E þ 06
Etched/ET Untested Test 1
NA 6.3
NA 89.2
1.55E þ 06 2.40E þ 06
Rolling direction Untested Test 1 Test 2 Test 3 Test 4
scratches/RD NA 0.001 4.5 6.3 0.001
NA None 89.2 89.2 89.2
1.38E 8.31E 1.13E 3.30E 8.61E
þ þ þ þ þ
NA None 89.2 89.2 89.2
1.46E 1.21E 2.91E 5.34E 4.28E
þ þ þ þ þ
Transverse direction scratches/TD Untested NA Test 1 0.001/4.5 Test 2 4.5 Test 3 6.3 Test 4 0.001
ing, from 595 to 474 mm3 and from 2788 to 2562 mm3, respectively. Another effect of the surfaces changes observed in Figs. 4–6 is a change in the size and nature of the intimate contact area between the die and the workpiece, which is observed in the Abbot bearing area curve. Fig. 7 shows the Abbot curves for the test specimen shown in Fig. 4. The TD surface (Fig. 7(a)), shows a numerical increase in the compensating gradient slope, mGC from 0.05 to 0.01. Additionally, the Abbot curve for the tested surface is ‘‘flatter’’ in that the linear portion of the curve is larger—indicative of greater intimate contact. The RD surface (Fig. 7(b)) shows a similar flatness and a similar mGC increase from 0.06 to 0.01. The Abbot curves for the specimen with the polished surface (Fig. 8) behave differently in that mGC decreases
VA (mm3)
Z2 (mm)
mGC (mm)
Sa (mm)
16 14 451
0.56 0.57 0.98
0.01 0.01 0.03
0.35 0.34 0.98
1417 354 484
1 11560 2866
0.31 0.83 0.49
0.00 0.07 0.05
0.13 2.57 1.53
280 194
5552 12350
2.93 2.81
0.11 0.12
3.52 4.11
06 04 06 06 04
595 2850 474 236 3042
2788 29 2562 14123 29
1.19 0.48 1.55 3.69 0.53
0.06 0.01 0.06 0.15 0.01
1.93 0.33 2.03 5.30 0.37
06 05 06 06 05
658 2036 251 170 2205
3010 213 11444 31430 60
0.67 0.61 1.87 2.64 0.41
0.05 0.01 0.11 0.16 0.01
1.67 0.45 3.54 6.23 0.58
from 0.00 to 0.07 and the Abbot curve becomes less flat after upsetting using lubricant 2. The Abbot curves resulting from the upsetting of RD and TD surfaces using lubricant 2 are shown in Fig. 9. Upon upsetting, the TD surface behaves similarly to the polished surface showing a decrease in mGC from 0.00 to 0.11 and a loss of flatness in the Abbot curve (Fig. 9(a)). The RD surface, however, is nearly identical before and after upsetting (Fig. 9(b)). The mGC for the RD surface is constant at 0.06 and the flatness of the Abbot curves are the same. The effects of upsetting on the surface-topography of the AR and ET surfaces are similar to the effects observed on the TD surface. Upsetting at a speed of 4.5 m s1 using lubricant 2 results in an increase in Sa and a decrease in mGC for both the AR and ET surfaces (Table 1). Also similarly to TD
Fig. 7. Abbot bearing area plots of test specimens before and after upsetting without lubrication at an upset speed of 4.5 m s1: (a) untested and tested surfaces with transverse direction scratches; (b) untested and tested surfaces with rolling direction scratches.
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
385
difference between the RD and TD surfaces, Dm ¼ 0. However, at velocities above and below 4.5 m s1 the TD surface has a lower m value for both the high and the low viscosity lubricant.
4. Discussion
Fig. 8. Abbot bearing area plots of test specimens before and after upsetting with lubricant 2 at an upset speed of 6.3 m s1.
Table 2 Dm values for different upset conditions for specimens with RD and TD surfaces on opposing sides Velocity (m s1)
Surface pair
Dm
No lube both sides 4.5 0.001
RD/TD RD/TD
0.046 0.077
Lube 2 both sides 4.5 0.001 0.0005
RD/TD RD/TD RD/TD
0 0.015 0.005
Lube 1 both sides 6.3 4.5 0.001
RD/TD RD/TD RD/TD
0.013 0 0.016
surfaces, the AR and ET surfaces both exhibit a decrease in NV and increase in VA after upsetting. The Dm values for the RD and TD surface combinations tested under different upset conditions are given in Table 2. In Table 2 Dm is reported as: Dm ¼ mRD mTD
(3) 1
Without lubrication at low upset velocity, 0.001 m s , the RD surface has a lower m value, Dm ¼ 0:077. At higher upset velocity, 4.5 m s1, the TD surface has a lower m value, Dm ¼ 0:046. For upsetting at 4.5 m s1 neither high nor the low viscosity lubricant result in any frictional
The surfaces upset without lubrication (Fig. 4 and Table 1) are nearly identical in terms of mGC and Sa independent of speed or initial surface preparation. The differences between the resulting RD and TD surfaces are limited to slight differences in VT and VA and the appearance of new surface-features at 908 to the original scratches on the RD surface after testing (Fig. 4(d)). These features found in Fig. 4(d) may be imprints from the die surface or effects from the creation of new free surface during upsetting. Nonetheless, it is clear that the resulting surfaces have very low roughness, have very low surface void volume and offer large areas of intimate contact between die and workpiece. At very low upsetting speeds, 0.001 m s1, the lubricant has more time to flow out of the interface in spite of the surface lay differences. As a result a large area of intimate contact is formed and the differences between the RD and TD surfaces after upsetting is very small as seen in Table 1. The lack of lubricant load bearing ability but the potential for lubricant presence and thus chemical action at the interface is indicative of the boundary lubrication regime. The opposite extreme of a large intimate contact area is complete separation of the die and workpiece, which is the case for the specimens with polished surfaces upset at high upsetting speed with high viscosity lubricant (Fig. 5). Under these upset conditions the lubricant is unable to escape from the die–workpiece interface and a lubricant cushion prevents intimate contact between the die and workpiece. In this case the surface roughens due to either pressure differentials within the lubricant at the interface or roughness-scale local in-homogeneties in the deformability of the surface. This combination of conditions, complete load bearing by the lubricant and surface roughening, indicates that the specimen in Fig. 5 was upset in the hydrodynamic lubrication regime.
Fig. 9. Abbot bearing area plots of test specimens before and after upsetting with lubricant 2 at an upset speed of 4.5 m s1: (a) untested and tested surfaces with transverse direction scratches; (b) untested and tested surfaces with rolling direction scratches.
386
W. Rasp, C.M. Wichern / Journal of Materials Processing Technology 125–126 (2002) 379–386
As stated by Rasp and co-workers [1] there is a competition between the hydrodynamic lubrication effects and the surface-topography effects. When surface lay and roughness effects are combined with lubrication regime effects at high upsetting velocity, the effects of the mixed, hydrostatic microhydrodynamic [11] and fully hydrodynamic lubrication interact. The result of this interaction is a wide variation of frictional behaviour and surface-topography changes depending on the initial surface-topography of the test specimen. For example, under the above conditions the TD surface roughens significantly more and shows a much greater change in mGC than the RD surface. The reason for this disparity is the same one proposed by previous authors [1–4]: the lubricant is unable to escape from the interface as easily on the TD surface. The high roughness surfacefeatures present on the RD and TD surfaces break through the lubricant film and contact the die surface, an effect recognizable via the flattened peaks (Fig. 6). The scratches on the RD surface provide isolated, but and short channels for the lubricant to flow to the edge of the interface. The scratches on the TD surface, however, provide longer isolated channels making it more difficult for lubricant to exit the interface resulting in more lubricant at the interface and higher pressure in the lubricant. The RD and TD surfaces are in two different lubrication regimes. The interaction of the lubrication regime effects and the surface-topography effects complicates the interpretation of the frictional results used to compare the different surfaces. As can be seen in Table 2 at high speeds, 6.3 m s1, and low speeds, 0.001 m s1, the TD surface has a lower m value when lubricant is used. When lubricant is not used the RD surface has a lower m value at low speed and the TD has a lower m value at high speed. Further complicating the situation is the result that the TD and RD surfaces perform equally during lubricated upset at 4.5 m s1. However, careful consideration and rationalization of the tribological conditions at the interfaces provides a possible explanation. The tribological interaction, which occurs at an upset velocity of 4.5 m s1 without lubricant, is purely mechanical resulting in a lower m value for the RD surface. The RD surfaces after testing with and without lubricant, Figs. 6(d) and 4(d)) appear very similar implying that despite the presence of lubricant the tribological behaviour with and without lubricant is also similar. It has already been proposed that the TD surfaces provide better conditions for the hydrodynamic, mixed or hydrostatic micro-hydrodynamic lubrication regimes to occur. In the hydrodynamic lubrication regime that prevails during upsetting at 6.3 m s1 with lubricant 2 the TD surface has a lower m value. During testing at 4.5 m s1 the RD surface is still predominantly controlled by mechanical tribological behaviour where it has a lower m than the TD surface. The TD surface under the same upset conditions is closer to the hydrodynamic lubrication regime where it has a lower m than the RD surface and can thus match the frictional performance of the RD surface. This balance of opposing frictional regimes nullifies the
effects of the surface lay. At lower upset velocities this balancing effect may not be present or observable due to the dependency of boundary lubrication regime behaviour on contact time and velocity at the interface. 5. Summary The change in the slope of the compensating gradient, mGC, is a good gauge for determining whether a tribological interface is moving toward or away from the hydrodynamic lubrication regime. Volume parameters, which measure the volume of the voids on a surface at a specific height, can be used in conjunction with mGC to describe the quantity, location and effect of a lubricant at an interface during forming operations. Surface lay transverse to the rolling direction promotes lubricant entrapment at the interface and promotes roughening and surface void formation. The effects of surface lay on frictional behaviour during upset without lubricant vary with upset velocity. At high upset velocity the TD surface has a lower m value than the RD surface and at low upset velocity the RD surface has a lower m value than the TD surface. As a result of their mutual dependency lubrication regimes and surface-topography must be carefully considered when comparing the tribological behaviour of different surfaces. References [1] C. Wolff, O. Pawelski, W. Rasp, A newly developed test method for characterization of frictional conditions in metal forming, in: Proceedings of the Eighth International Conference on Metal Forming, Krakow, 2000, pp. 91–97. [2] W. Rasp, P. Ha¨ fele, Investigation into tribology of cold strip rolling, Steel Res. 69 (4–5) (1998) 154–160. [3] J. Lordan, Beurteilung des Schmierungsverhaltens unterschiedlich texturierter Oberfla¨ chen mit Hilfe des Streifenziehversuches, Diplomarbeit Universita¨ t, Gesamthochschule Duisburg, 1996. [4] O. Pawelski, W. Rasp, S. Draese, P. Ha¨ fele, Influence of hydrodynamic effects and surface roughness on tribological phenomena in cold strip rolling, in: Proceedings of the Fifth ICTP, Columbus, OH, 1996, pp. 31–34. [5] J.V. Reid, J.A. Schey, Full fluid film lubrication in aluminium strip rolling, ASLE Trans. 21 (3) (1978) 191–200. [6] W.R.D. Wilson, W. Lee, Mechanics of surface roughening in metal forming processes, Trans. ASME, J. Manuf. Sci. Eng. 123 (2001) 279–283. [7] W.R.D. Wilson, Friction and lubrication in bulk metal-forming processes, J. Appl. Metalworking 1 (1979) 7–19. [8] K.J. Stout, P.J. Sullivan, W.P. Dong, E. Mainsah, N. Luo, T. Mathia, H. Zahouani, The Development of Methods for the Characterization of Roughness in Three Dimensions, EUR 15178, ECSC-EEC-EAEC, Brussels-Luxemburg, 1993. [9] ISO 4287, Geometrical product specifications (GPS)—surface texture, Profile method—terms, definitions and surface texture parameters, 1997. [10] J.A. Schey, Tribology in Metalworking—Friction, Lubrication and Wear, American Society for Metals, Cleveland, OH, 1984. [11] H. Kudo, A. Azushima, Interaction of surface microstructure and lubricant in metal forming tribology, in: Proceedings of the Second ICTP, Stuttgart, 1987, pp. 373–384.