Friction of Aluminium in Deep Drawing

Transcription

1 Friction of Aluminium in Deep Drawing Wilko C. Emmens, Jan Bottema Hoogovens R&D, P.O.Box , 1970 CA IJmuiden, the Netherlands IDDRG CONGRESS 1998 Abstract Aluminium shows a large amount of asperity flattening in sliding friction contacts. This flattening causes the friction to depend on pressure. This results in a wide range of mixed lubrication; pure boundary lubrication is hard to obtain. In deep drawing friction occurs under conditions of increasing pressure. As a result, the coefficient of friction in the blankholder can be very low, near hydrodynamic lubrication. 1. Introduction Using aluminium as the metal of choice in the automotive industry will help society to meet the demands of its increasing commitment to the environment. In the past, significant gains in automobile fuel efficiency have been achieved through aerodynamic design and drive train improvements. As performance limits are approached in these areas, the recent interest is in alternate materials that will allow lighter designs. Aluminium is increasingly attractive as a candidate for material substitution because of its strength and stiffness to weight ratio. The penetration of aluminium in the automotive market is being led by the use of aluminium sheet for closures. Compared with traditional steels, aluminium rolled products for doors, closures or wings can bring a weight reduction of up to 50%. The basic requirement on automotive sheet is to have a formability as high as possible without compromising its strength. Body panels are only semistructural; the main functional requirements being dent resistance and stiffness. In order to minimise the part weight and cost, a high yield strength is required. However, the material also requires high formability and minimal spring-back. The preferred approach has been to use heat treatable aluminium alloys which have a low T4 strength, but which strengthen during paint baking finishing operations. The principle choice of aluminium closure sheet among automotive manufacturers in Europe has been AA6016, in contrast to North America where an emphasis on strength has led to the adoption of AA6111 as the alloy of choice. Since the 6016 alloy was introduced in the 80 s, various process improvements and compositional adjustments within the AA range have been made to improve its formability, hemming performance and response to lower temperature paint bake treatment. In the 80 s, EDT surface texturing was introduced for the AA6016 alloy, to improve the forming behaviour due to better lubrication. A recent development is the introduction of EBT surface texturing. This type of texture combines high roughness with low waviness which is beneficial to forming and to paint appearance. The increasing use of aluminium has lead to an increase of research on the formability of aluminium automotive sheet at Hoogovens R&D. During the tests it has been noticed that under certain conditions the deep drawing behaviour of aluminium differs significantly from that of steel. This paper de-

2 scribes in more detail why the behaviour of aluminium is different from what is traditionally observed for steel. NOTE: in this paper the behaviour of aluminium is sometimes compared to that of steel, without actually presenting results for steel. In general, uncoated steel has a classic behaviour, in such a way that the influence of process conditions (and more specifically: the pressure) is fully described by a single Stribeck curve without large additional effects. A full comparison between steel an aluminium and more detailed information however can be found in [1]. 2. Some deep draw results: defining the problem In deep draw experiments the punch force is a convenient measure to study all kinds of influencing factors. The relation between punch force and blankholder force is of particular interest, because the slope of the curve which represents this relation is roughly twice the coefficient of friction in the blankholder [1]. An example can be found in figure 1 (left) which shows results for a large cup (300 mm diameter) from aluminium. The slope of the curve is almost equal for all materials, in the range This means that the coefficient of friction in the blankholder is roughly the same for all materials. The different lines reflect differences in thickness and grade. The right hand part of figure 1 shows other situations (with a different product) in which the relation of punch force to blankholder force is not linear. At low blankholder forces the slope of the curve is in agreement with values of the coefficient of friction which is measured in friction tests. At higher blankholder forces the slope reduces to almost zero, leaving a very remarkable situation that the blankholder force could be increased to the maximum of the press without causing fracture! This effect has already been observed by Kasuga [2] and appears to be typical for aluminium. Kasuga attributed this effect to friction in the blankholder, and, to study this phenomenon more specifically, friction tests on aluminium have been carried out in our laboratory. max. punch force (kn) large cup 293 mm punch aluminium AA mm blankholder force (KN) AA mm AA mm max. punch force (kn) small cup, 75 mm punch Al AA5754, 1.0 mm blankholder force (kn) Figure 1. Examples of results of deep draw tests. The figure shows situations where the relation between (maximum) punch force and blankholder force is either linear (left), or not linear (right). 3. Friction tests: general principles. The results of friction tests will be presented in the form of Stribeck curves. Although the use of Stribeck curves is a very convenient way of mapping all kinds of effects on friction, they are rarely

3 used. In a Stribeck curve the measured coefficient of friction is plotted as a function of the Hershey parameter H defined as H = ηv/p, where η is the dynamic viscosity of the lubricant, v the speed (it is assumed that one of the two contacting surfaces is stationary), and P the external macroscopic pressure. H has unit length (m) and traditionally is plotted on a logarithmic scale. A general representation of the Stribeck curve is presented in figure 2 (left). Please note that both η, v and P are typical parameters which are defined by the process conditions. That means that the parameter H can been considered to be a measure for the conditions of the process studied. In a Stribeck curve three regimes of lubrication can easily be distinguished: boundary lubrication, where the total load is carried by the roughness asperities; hydrodynamic lubrication, where the total load is carried by the lubricant film; mixed lubrication, which is a mixture of boundary and hydrodynamic lubrication. Originally, Stribeck curves were constructed for friction in journal bearings. In the study of friction in forming processes it has been noticed however that the influence of pressure is not described completely by the parameter H. It is therefore customary to construct different curves for each pressure value. 0 BOUNDARY MIXED HYDRODYNAMIC friction coefficient æ O H = η v P æ hydr friction coefficient æ 0 tanhyp fit 0 H C H (m), logarithmic logh 1 logh C logh 0 logh Figure 2. The Stribeck curve. The left hand figure gives a general representation in which the three regimes of lubrication can be distinguished. The right hand figure shows how the transition points H 0 and H 1 can be calculated by fitting a tanhyp function through the measured friction data. These points are at the intersection of three lines tangent to the curve. The condition of boundary lubrication is of particular interest for the press shop. In conditions of boundary lubrication the friction hardly depends on pressure, speed, lubricant viscosity or roughness. This means that the result of the forming operation is hardly sensitive to variations in process conditions or (some) material properties. In practice the condition of boundary lubrication is usually obtained by applying a low-viscous lubricant, or by applying a very limited quantity of lubricant. For a more detailed study of the friction the so-called transition points are of interest. These are the values of H at the transition from boundary to mixed, and from mixed to hydrodynamic lubrication. This points can be obtained by plotting three tangent lines to the curve and calculating the intersection points, see figure 2, right. More precisely this is done by first fitting a tanhyp function through the measured data, the shape of a tanhyp function fits the actual friction data in nearly all observed situations (both for steel and for aluminium).

4 4. Results of friction tests The observed effect on the friction of aluminium will now be discussed on the basis of results of a series of friction tests on aluminium with different types of roughness. The conditions for these experiments are listed in table 1, and the material properties in table 2. Friction experiments have been carried out in our laboratory using a rotating friction tester. In this tester a punch with three notches is pushed against a flat sample and the sample is then rotated over some distance. Figure 3 shows the shape of the punch; a more general description of the device can be found in [1,3]. Table 1.Conditions in friction tests. Slider punch with three notches 12x12 mm² (as in figure 3) speed 1, 2, 5, 10, 20, 50, 100, 200, 500, 1000 mm/s pressure 2, 5, 10, 20 MPa lubricant three mineral oils, viscosity at 20 C: 16, 55, 300 mpa.s Table 2. Material properties in friction tests material aluminium AA6016 T4 thickness 1.2 mm roughness MF: MF type roughness, R a = 0.11 / 0.25 µm (at 0 and 90 orientation) EDT low: EDT type roughness, R a = 0.41 µm EDT high: EDT type roughness, R a = 0.80 µm 100 mm Figure 3. Punch with three notches as used in the rotating friction tester. The punch is pushed against a material sample, after which the sample is rotated over 120. friction coef. at boundary lubr pressure (MPa) MF EDT low EDT high Figure 4. Friction at boundary lubrication as a function of pressure, for three aluminium materials. The results are presented in figure 5 as four sets of Stribeck curves, one set for each pressure value. It can be seen that the area of mixed lubrication (the oblique part of the curve) is very wide, spanning three or four orders of magnitude in H. This is much more than is usually seen for steel; for steel the width of the mixed area is about one order of magnitude. Due to the wide area of mixed lubrication the areas of boundary and hydrodynamic lubrication are very small. Some important properties can be deduced from these curves. Compare now the results for both EDT materials. These two materials have a roughness with identical structure but a difference in

5 roughness height. The Stribeck curves for these two materials are more or less parallel. The coefficient of friction at boundary lubrication (the friction value at the far left) is a little higher for the higher roughness. The major difference however is that the curves for the higher roughness are shifted to the right. This means that for process conditions in the mixed regime, material with higher roughness will exhibit higher friction. This relation is generally valid for all kinds of materials. The results for MF are not directly comparable to the results for EDT. At low pressures the friction of MF equals that of the friction of the low EDT. At higher pressures the friction of MF 'travels' between the friction values of the two EDT types, the slope of the curve in the mixed area is steeper than for EDT. coefficient of friction coefficient of friction MPa MPa EDT high EDT low MF EDT high EDT low MF coefficient of friction coefficient of friction MPa MPa EDT high EDT low MF EDT high Figure 5. Results of friction tests on three aluminium materials with different roughness, plotted as a function of pressure and the Hershey parameter H. The lines are tanhyp functions fitted though the data. EDT low MF The influence of pressure can also be noticed in the friction of EDT. Firstly, the coefficient of friction at boundary conditions increases with increasing pressure. This can be deduced from figure 5, but better from figure 4 where the values of the coefficient of friction at boundary lubrication as determined using the tanhyp fits of figure 5, have been plotted as a function of pressure. This figure again shows that the friction of MF differs from that of both EDT types of roughness. Secondly the pressure also has an influence on the position of the Stribeck curve. This is more difficult to see from figure 5, therefore the transition points as defined in figure 2 have been calculated, the results are shown in figure 6 (left). This figure shows the large influence of pressure on the transition point. The influence of pressure is larger for the MF roughness than for the EDT roughness. The influence of pressure on the location of the transition point H 1 (boundary - mixed lubrication) is particularly strong. Increasing the pressure by a factor 10 (from 2 to 20 MPa) decreases the value of H 1 by

6 more than a factor 100, logh 1 reduces by approximately 2.2. In the right hand part of figure 6 results from another series of tests are presented, illustrating that this strong influence of the pressure is also found in other types of roughness such as EBT, and is a general property of aluminium. Some results obtained for uncoated steel are presented as well for comparison. This strong influence of pressure however strongly affects the behaviour of aluminium in deep draw operations, as will be shown now H 0-8 H H 1 log H -11 EDT high EDT low MF H pressure (MPa) -11 steel EBT EDT pressure (MPa) Figure 6. Transition points as defined in figure 2, as functions of pressure. Boundary lubrication is below H 1, mixed lubrication is between H 1 and H 0, and hydrodynamic lubrication is above H 0. The left hand side shows results obtained from the tests described above; the right hand side from similar tests in which EBT roughness was compared with EDT roughness, some results for uncoated steel are presented as well for comparison. log P Figure 7. Schematic representation of a transition diagram by analogy with figure 6. B, M and H denote the lubrication regimes (boundary, mixed and hydrodynamic). P1 and P2 denote conditions during deep drawing, see text 5. Implications of influence of pressure To fully understand the consequences of the influence of pressure on friction we have to realise what is happening during deep drawing. The main source of friction in a deep draw operation is in the blankholder. During the deep draw operation, material is pulled out of the blankholder area. Whereas in many cases the blankholder force remains constant, this movement of material causes the normal pressure on the workpiece in the blankholder to increase during the drawing operation. Due to plastic deformation the material will thicken in the blankholder as depicted in figure 8 (left). This may proceed so far that the inner part of the flange fully looses contact with the tool. This will further reduce the area of contact in the blankholder, and, due to the thickening, the blankholder force will be concentrated on the outer edge of the flange. To get a better idea of the actual conditions, finite element simulations have been carried out using the tool geometry and material properties as used in the tests described hereafter. In figure 8 (right) the normal pressure on the outer edge of the flange has been plotted as a function of the position of the edge of the flange during the deep draw operation. If the material would not thicken, the normal pressure would be equal to the blankholder force divided by the total area of the flange, this pressure

7 has also been plotted in figure 8 ("mean stress"). At the start of the operation the material is perfectly flat and the pressure at the outer edge equals the lower limit. When the deep draw operation starts and the material is being pulled out of the blankholder area (the flange edge moves inward), the material starts to thicken at the edge and the load gets concentrated quickly at the outer edge: the normal pressure rises quickly. This continues to a situation were the inner part of the flange looses contact with the tool; in that case the pressure on the outer edge is more than twice the mean stress. When the operation proceeds the amount of plastic deformation in the flange increases and the actual area of contact reduces again. In the final stages the pressure distribution on the flange becomes more even and the pressure at the outer edge approximates the lower limit again. 120 mean stress x2 PUNCH BLANKHOLDER DIE normal stress (MPa) position of die radius mean stress in total blankholder area calculated stress at flange edge from simulation actual position of flange (mm) Figure 8. Conditions at the flange in deep drawing. Due to plastic deformation the outer flange edge will thicken, causing a concentration of the blankholder force on the outer part of the flange (left). The right hand part shows results of a finite element simulation of a deep draw operation, showing that the actual normal pressure during the drawing operation will increase significantly. During the operation the flange will be drawn inwards, the actual position of the flange edge is plotted as the abscissa. The mean stress is defined as the blankholder force divided by the actual total area of the flange. A very high blankholder force has been used in the calculations. Now let's see what the effects of increasing pressure are on friction. Figure 7 presents a schematic transition diagram, in the same way as figure 6. The labels B, M and H denote the three lubrication regimes: boundary, mixed and hydrodynamic lubrication. Some possible H-P relations have been indicated by the lines 1-4. The lines 1 and 3 represent the situation for materials which show only a small influence of pressure on the transition points, as for steel. The lines 2 and 4 represent the situation for aluminium according to the results shown in figure 6. A distinct tribological condition, defined by viscosity, speed and pressure, is presented in the diagram by a given location. Suppose that the conditions in the blankholder at the start of the draw operation are indicated by point P1. As has been shown, during the draw operation the (effective) pressure in the flange will increase, the location of the condition shifts to position P2. The line connecting these two points has a slope -1. In the situation from figure 7 the trajectory P1-P2 intersects line 3, entering the regime of boundary lubrication. This means that in situations where pressure hardly affects the transition points, increase of the pressure will eventually cause the friction to operate under conditions of boundary lubrication. This is happening when processing steel.

8 For aluminium the situation as presented by the lines 2 and 4. The trajectory P1- P2 does not intersect these lines. This means that a situation of boundary condition will not be reached. When also the transition mixed-hydrodynamic shows a strong influence of pressure almost pure hydrodynamic lubrication might occur. Some evidence for that can be found in the results for EDT in figure 6 (left), for higher pressures. The implication of all this is: with aluminium it is not possible to shift from mixed lubrication to boundary lubrication just by increasing the pressure coefficient of friction MPa 200 mm/s Figure 9. Friction results illustrating effect of pressure. Increasing pressure will reduce the friction whereas the Stribeck curves shift to the left. 50 This statement is illustrated by the results of figure 9, which have been obtained from friction tests on anodised aluminium not described in this work. The data of figure 9 for one value of the pressure form a part of a Stribeck curve. When we look at results obtained for one value of speed then increasing the pressure is seen to shift the points in the Stribeck curve to the left (follow the arrows). However, whereas the Stribeck curves themselves shift further to the left, increasing pressure actually reduces the friction! 6. Friction tests at high pressure The influence of pressure on the friction of aluminium has been studied further by performing friction experiments at (very) high pressures. This was achieved by making a punch as in figure 3 with very small notches (in fact two punches were used). The tests were carried out by increasing the pressure in small steps on different samples. For the very small notches this process had to stop at a pressure where suddenly severe scoring occurred. Inspection of the samples revealed that at that point bulk plastic deformation under the notch occurred. Results obtained on aluminium 6000 EDT are presented in figure 10, left. The two punches with different notch sizes used in the tests yielded different results. This has been observed on many occasions, elsewhere as well, but so far no good explanation have been found as to why in friction tests with flat sliders the friction depends on the size of the slider. Of more interest here is the influence of speed. If increasing the pressure would eventually lead to boundary lubrication, the influence of speed would vanish. However the results show that even at very high pressures the coefficient of friction decreases with increasing speed, indicating that lubrication is still mixed. Actually, the influence of speed does not depend on pressure, as can be seen in figure 10, right.

9 coefficient of friction mm/s 20 mm/s 100 mm/s pressure (MPa) medium notches small notches coefficient of friction MPa 140 MPa speed (mm/s) Figure 10. Results of friction tests at (very) high pressures. Left: friction as a function of pressure. Right: friction as a function of speed. 7. Deep drawing experiments Table 3. Conditions in deep draw experiments punch 293 mm diameter, flat top, radius 20 mm die gap 3,5 mm, radius 20 mm blank diameter 480 and 520 mm (deep draw ratio: 1.64 and 1.77) speed low = 4-8 mm/s, high = mm/s blankholder force various lubricant preserving oil, viscosity at 20 C: 55 mpa.s Table 4. Material properties in deep draw tests material aluminium AA6016 T4 thickness 1.2 mm roughness EDT: EDT type roughness, R a = 0.70 µm EBT1: EBT type roughness, R a = 0.95 µm EBT2: EBT type roughness, R a = 1.15 µm yield stress 125 MPa tensile strength 230 MPa n 0.23 To further investigate the effects of friction at high pressure, deep draw experiments have been carried out. These experiments were carried out on a large cylindrical cup, and apart from the blankholder force also the speed, the amount of lubricant and the blank diameter were varied. Test conditions are stated in table 3, and an overview of material properties is given in table 4. The speed of the press and the blankholder force showed a large interaction. Therefore the speed could not be held constant during the series. However, two ranges of speeds have been obtained with a difference large enough to allow conclusions about the influence of speed to be drawn. Two dosages of lubricant were applied, hereafter simply called wet and dry condition. At the wet condition an overdose was generously applied to the blank. At the dry condition a very small amount of lubricant was applied to the blank and evenly distributed, after which the blank was wiped 'clean'

10 maximum punch force (kn) maximum punch force (kn) dry dry with a dry cloth. Moreover the tool was wiped clean before each pressing. This method left just enough lubricant on the sheet to prevent metal-to-metal contact (galling!) and ensure sparse lubrication. However, the actual amounts of lubricant have not been measured. Results are presented in figure 11, in general both blank diameters show identical results. An influence of the speed is only present in the wet lubrication condition. At that condition, the higher speed clearly results in a lower punch force. At the dry condition however there is no influence of speed, and the punch force is much higher than at the wet lubrication condition. Moreover, an influence of roughness can only be seen at the wet condition, and in particular at the smallest blank (less critical process). The slope of the punch force - blankholder force relation is very low, values of were calculated for the combination of small blank, high speed and wet condition. This indicates that the coefficient of friction was very low in the blankholder, and probably the condition of friction was not far away from pure hydrodynamic lubrication. Consequently, the blankholder force could be increased until the limit of the press without causing product failure. The observed influences of speed and amount of lubricant also indicate that the observed phenomena are caused by friction effects. 8. Discussion 480 mm wet, slow blankholder force (kn) wet, slow 520 mm wet, fast So far we have shown that the extreme pressure dependency of the transition points may cause the friction in deep drawing of aluminium to decrease significantly. This is generally valid for all applications where the friction during the operation is subjected to a (continuously) increasing pressure. It has not become clear where that large pressure dependency stems from. The first author has already shown that aluminium is subjected to a severe amount of flattening of the asperities in friction [3,4]. Figure 12 has been taken from those publications showing to what extent the roughness height may decrease under friction conditions. It has also been shown that after EBT1 EBT2 EDT blankholder force (kn) wet, fast EBT1 EBT2 EDT Figure 11. Results of friction tests on a large cylindrical cup, for three types of aluminium with different roughness. Two blank diameters were used (480 mm and 520 mm).

11 correcting the measured coefficient of friction for this amount of flattening, the friction of aluminium resembles that of steel for which the phenomena described in this report do not occur. The influence of pressure on friction could be attributed to the influence of pressure on the flattening of the roughness. Recently a model for mixed lubrication has been developed which takes into account the flattening of the asperities [1]. Although the model is very simple, it successfully predicts the observed phenomena. Figure 14 shows results of calculations compared to actually recorded Stribeck curves. The sharp transition from mixed to hydrodynamic lubrication predicted by the model stems from the fact that the model assumes all asperities to have the same height, whereas of course in practice the asperity heights show a statistical distribution. Because the model cannot predict the friction at boundary lubrication, the recorded curves have been normalised to a coefficient of friction at boundary lubrication of 1. The agreement between predicted and measured friction is satisfactory. R PM (æm) MPa 5 MPa 10 Mpa 20 MPa Figure 12. Reduction of roughness height (depicted by R pm ) in friction contacts [4]. R a / R a, travel distance (mm) Figure 13. Relative reduction of R a in the friction tests as a function of slider travel distance relative friction coefficient relative friction coefficient logh (m) logh (m) Figure 14. Influence of pressure on Stribeck curves as predicted by a simple friction model (left), and as actually measured for aluminium in the form of tanhyp fits through the data points (right). The numbers at the lines denote pressures in MPa. Figure from [1]. Relative friction is defined as µ/µ 0. The question still remains why for aluminium the asperities flatten so much more than for steel. Micro hardness measurements indicate that the hardness of aluminium is roughly half of the value for

12 soft steel (for the materials used in our tests). But both are much softer than the tool, which is roughly five times as hard as of soft steel. So for both materials we have a situation that a soft working material slides against a hard tool. This still does not account for the large difference in flattening for steel and aluminium. Recently the first author has suggested [1,5] that the flattening mechanism proposed by Tabor [6] might occur. This mechanisms implies that flattening happens almost instantaneously when sliding begins. However, the results presented in figure 13 show that a certain amount of sliding is necessary for full flattening, contradictory to the mechanism proposed by Tabor. So it still may be just a matter of mechanical properties. Please note that the results presented in figure 13 have been obtained from very rude tests and are only to be used as an indication. More research is needed. Kasuga proposed that the phenomena observed are caused by hydrostatic lubrication due to lubricant trapped in isolated pockets. This mechanism is not accepted everywhere, but evidence for this mechanism has been found in earlier experiments in our laboratory [1,4]. A possible contribution of hydrostatic lubrication can only occur when lubricant is actually trapped in pockets, and the severe flattening of roughness at aluminium can cause such state. It is recommended to put more emphasis on the study of (additional) hydrostatic lubrication in metal forming. 9. Conclusions Aluminium shows a strong flattening of roughness asperities in sliding contact. The flattening of asperities induces a strong influence of pressure on the process conditions at which mixed lubrication occurs. For aluminium, it is not possible to go from mixed lubrication to boundary lubrication just by increasing the pressure. In deep drawing of aluminium, situations may occur in which the product does not fracture, even at the highest blankholder force. 10. Literature [1] Emmens, W.C., Tribology of Flat Contacts, and its Application in Deep Drawing, Thesis University of Twente, the Netherlands, 1997 [2] Kasuga, Y., Yamaguchi, K., Friction and lubrication in the deformation of metals, 1 st report, Bulletin JSME, vol. 11, no. 44, 1968, pp [3] Emmens, W.C., A Novel Design Friction Tester, IDDRG Working Group Meetings, Pisa, Italy, May 1991 [4] Emmens, W.C., Schoepen, F., Some frictional aspects of aluminium in sheet metal forming, Proceedings 14th IDDRG Congress, Eger, Hungary, June pp [5] Emmens, W.C., Some Frictional Aspects of Aluminium in Deep Drawing. Proceedings 1st International Congress on Tribology of Manufacturing Processes (ICTMP), Gifu, Japan, October 1997, pp [6] Tabor, D., Junction Growth in Metallic Friction: the Role of Combined Stresses and Surface Contamination, Proceedings of the Royal Society of London A, vol. 251, 1959, pp