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Experimental device for tribological measurement aspects in deep drawing process
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journal homepage: www.elsevier.com/locate/jmatprotec
Experimental device for tribological measurement aspects in deep drawing process X. Roizard a,∗ , J.M. Pothier a,b , J.Y. Hihn b , G. Monteil a a b
Laboratoire de Micro-Analyse des Surfaces (LMS), ENSMM, 26 chemin de l’´epitaphe, 25030 Besanc¸on, France Laboratoire de Chimie des Mat´eriaux et Interfaces (LCMI), IUT, rue de l’Observatoire, 25000 Besanc¸on, France
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Article history:
The purpose of this paper is to improve strip drawing test in order to simulate sheet/tool
Received 5 March 2007
contact conditions during a stamping operation. An original multi-pass strip drawing test
Received in revised form
in cylinder/sheet/cylinder contact geometry is presented and we emphasize the influence
10 March 2008
of certain factors rarely taken into account in the experimental simulation. In particular, we
Accepted 14 March 2008
present examples where the bulk plastic deformation and the tools average temperature influence the level of friction and the sheet surface asperities flattening. To try to readjust the friction evolutions with the modifications of surface microgeometry, the tribometer is
Keywords:
equipped with various measurements of relative sheet/tool displacement, including or not
Deep drawing
the stiffness machine and the stiffness of the contact. This was made possible by a poten-
Tribology
tiometric measurement where the tool plays the part of potentiometer and sheet that of the
Strip drawing
track, coupled to a measurement using a LVDT sensor.
Temperature contact
© 2008 Elsevier B.V. All rights reserved.
Bulk deformation Adhesive junction properties
1.
Introduction
Deep drawing is a forming process for metal sheets frequently used in the automotive industry. The conditions in which the contact takes place are especially severe: boundary lubrication in pure sliding friction, high contact pressure always implying both surface asperity and bulk plastic deformation, cyclic progressive transfer layer build-up on the dies, etc. Previously, most of the research works has been focused on the experimental and numerical simulation of the contact conditions during this operation. Many researchers – with comparisons between FEM simulations and experimental results for stamping operations – showed that friction strongly influences the formability by affecting the strain distribution in various regions of the sheet: Hortig and Schmoeckel (2001) have shown that a small friction coefficient does not always lead
∗
Corresponding author. Tel.: +33 3 81 40 28 54. E-mail address: [email protected] (X. Roizard). 0924-0136/$ – see front matter © 2008 Elsevier B.V. All rights reserved. doi:10.1016/j.jmatprotec.2008.03.023
to lower tribological loads, however Choudhury et al. (2006) have observed that, with the increase of friction coefficient the blank holder force decreases, Verma and Chandra (2006) concludes that the limiting drawing ratio (LDR) decreases with increase in the friction. On the other hand, Sutcliffe et al. (2003) and Matuszak (2000) have shown that the lubrication regime is critical for formation of the transfer layer. Therefore, Bech et al. (1999), using a transparent tool entrapment, has observed that the smaller the ratio of the estimated lubricant film thickness to the roughness is, the higher the transfer layer. Since Bowden and Tabor, a common opinion in tribology is that the pinning of the surfaces is due to cold-welded junction’s formation, so that the sliding corresponds to these junctions’ shearing. Due to economic and environmental pressures, industry is led to optimize all operating parameters by using several types of tribometers, reproducing more or less one part of
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the drawn sheet. Strip drawing test is one of these techniques and allows different tool geometries: flat/flat used by Kim et al. (2002), radial strip drawing used by Sniekers and Smits (1997), cylinder/cylinder used by Roizard and von Stebut (1995), etc. Unfortunately, most of strip drawing experiments can only simulate plastic deformation on the bulk material via increased surface pressure at the die/sheet interface: see for instance when friction tests of Wihlborg and Gunnarsson (2000) were performed in a Bending Under Tension test. To separate the bend—unbend force from the pure friction component, Sulonen (1981) has proposed a special hydraulic test rig. But even in this case, the overall bulk plastic deformation and the contact pressure applied across the surface asperities remain coupled parameters. Moreover, those tests are generally carried out at room temperature whereas the real contact temperatures rise up to 100 ◦ C after stamping of a series of sheets in industrial press shop owing to frictional self-heating. Recently, Novotny (2003) classified the suitability of lubricants and tool materials with a strip drawing test equipped with a heating device. It had been shown that fluid lubricants generate low-friction forces at room and elevated temperature if their viscosity at room temperature is high enough. This confirms the needs of tests showing the maximum of mechanical contacts similarities with industrial press shops, while keeping independence in the control of input parameters. This will make possible to optimize technological choices for a given metal sheet, as for example lubricant nature and quantity, tool material as well as allowing a better comprehension of physical behaviour of the contact. In this objective, it is possible to improve the information given by our device by improving the device equipment (heating tools, sample/tool positioning and bulk pre stressing) and by improving test methodology (multi-pass, profilometric relocalisation). In the same way, it is required to discriminate the behaviour of the contact and the device itself. This is particularly true for some experimental conditions, when the so-called stick-slip effect appears. Both experiments and simulations show that in all cases, when the static frictional is nonzero, the system exhibits a stick-slip motion at low velocities. Schematically in the stick-slip sliding regime, friction force as a function of time occurs shown a typical shape presented in Fig. 1a. In classic tribological experiments, a spring is attached to the slider, and its end connected to a base which moves with a constant velocity Vspring (Fig. 1a). The same is true for strip drawing, where the elasticity of the moving parts of the device plays the same role. The output (measured) parameter is the spring force Fspring . Let us assume that initially the system is in rest and the spring has its natural length. When the base begins to move, the spring stretches, F increases until it reaches the threshold value Fs corresponding to the static frictional force, and the block starts to move. Then, Braun and Naumovets (2006) showed that, due to inertia, the slider accelerates to catch the base. If Vspring is small, F decreases down to the “backward” threshold force Fb , and the slider stops. Then the process repeats, so the stick-slip motion occurs as shown in Fig. 1b. To understand the complete mechanisms of this phenomenon, i.e. accede to the mechanical properties of the junctions between dies and sheet, it is necessary to obtain the true relative displacement between parts. That is why our tribometer is equipped with an original measure of relative sheet/tool displacement.
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Fig. 1 – (a) A standard experimental setup for tribology studies and (b) friction force as a function of time in the stick-slip regime (schematically).
The aim of the present work is to present the experimental possibilities of an original device dedicated to tribological measurement aspects by keeping the maximum of similarities with industrial press shop, whereas allowing a better understanding of friction mechanisms by individual control and specific measurement methods.
2. Experimental details—setup and measurements techniques The test rig adopted is equipped with a cylinder/cylinder contact geometry very convenient for tool/sheet alignment and friction data interpretation. Its specific feature is to allow perfectly independent control of contact pressure and bulk tensile stress as well as instantaneous die/sheet sliding speed and contact temperature (Fig. 2). The actuator of a MTS® testing machine generates the relative sliding motion of the sheet with respect to both dies. The deformation induced by the tools may be qualified as macro-elastic as only the roughness is affected in a plastic manner. A particular attention was given to monitor and control mechanical parameters, i.e. to control independently contact pressure and axial tensile load along the S.D. sheet. This last one is placed between grips on a special load frame which applies a constant polarization load L to the sheet’s ends by means of rubber bellows. Because of friction, the part of the sheet on the side opposite to the dies sliding direction will be under a longitudinal tensile load equal to (L + 2 P), where is the friction coefficient and L the tensile polarization load. As a complement, our strip drawing test rig allows die temperature control ensured by means of heating coils inserted into the die grips and powered by means of a P.I.D. unit in conjunction with a platinum probe. The relative sheet/die displacement is given by a LVDT sensor, completed by an original electrical measurement in order to take into account the real displacement. Indeed, this local
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Fig. 2 – Synoptic description of strip drawing tribometer.
displacement of the tool vs. sheet is based on a potentiometric method, whose cursor and track are constituted by tool and work sheet (Fig. 3). This gives a measure as close as possible of phenomena occurring at the interface. The principle of this potentiometric method is based on the creation of a constant voltage gradient, by injecting a constant current into the constant section sample sheet. The measure itself consists in the acquisition of the difference of potential between the moving tools and the fixed electrodes. This measure is obtained through a high impedance loop so that the current passing through the tool/sheet interface is lower than 10−12 A mm−2 and allows neglecting any thermal or tribo-oxydation effects. Then, the highest resolution for this method implies the highest value for the voltage gradient, because the displacement is directly proportional to this voltage. However, the increase of voltage gradient is obtained from the increase of the injected current, while staying below the threshold of the heating by
Joule effect. As an example in the case of aluminium, the resistivity is around 2.6 × 10−8 m and for 50 mm2 section sheets, the range of magnitude of the electric resistance is about 0.5 m m−1 . So, an injection current of 50 A induce a power to be dissipated lower than 25 W m−2 . In those conditions, the experimental temperature variation never exceeds 1 ◦ C. As the potential gradient is 26 mV m−1 measured through an instrumentation amplifier with gain 3000 and that the resolution of our acquisition card is 150 V, the resolution in distance is around 2 m. Furthermore, the higher the sample resistivity is and the more the accuracy increase up to 0.8 m for steel. In order to keep the maximum of precision and repeatability in all measurements, it is necessary to respect a precise geometrical location of electrodes, in particular to insure them a “perfect” perpendicularity with the lines of electric field. This location is done by means of a template for standardization of the electrode positions. As theoretical calculations led with the resistivities of pure metals do not allow defining an exact correspondence between the measured tension and the relative position of the tool/sheet, a calibration electrode has been added for which the distance with the reference electrode is perfectly known thanks to the template. Samples used in the present work consist of low carbon steels sheets (50 mm × 300 mm × 0.7 mm) and aluminium EDT 5182 (50 mm × 300 mm × 1 mm). QUAKER® 6130 was applied as lubricant for each sample at 2 g m−2 . The velocity of the tensile machine actuator was kept quite to low values between 0.1 and 33 mm s−1 , in order to remain in boundary lubrication regime and to avoid the risk of galling. Cylindrical tools (diameter 20 mm, length 60 mm) are made of uncoated high-speed steel Z85WD06, with roughness lower than 0.1 m (arithmetic average roughness, Ra). Otherwise, the length of the sliding track has been limited to 3 mm, which is enough to systematically ensure stabilized functioning. The variables acquired during a test are sampled within a frequency of 200 Hz. They include: • The normal force FN (N) applied to the metal sheet by the tools. • The displacement (mm) of the actuator of the tensile machine, which produces a motion of the tools in relation to the metal sheet.
Fig. 3 – Potentiometric measure of sheet/tool relative displacement.
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Fig. 4 – Illustration of the multi-pass strip drawing (MPSD). Topography and friction coefficient measurements on the same scanned area of the metal sheet between two consecutive passes.
• The tangential stretching force FT (N), transmitted by the actuator of the tensile machine to generate motion. • The apparent displacement (mm) of the tool in relation to the clear tip of the work sheet. Measurement of this displacement should allow identification of tribometer compliance, presently estimated at 0.5 m/N. • The local displacement (mm) of the tool in relation to the work sheet. This value is based on the potentiometric method.
3.
Results
3.1. Multi-pass strip drawing and 3D profilometry measurement Depending on the number of times the same sheet surface area goes through frictional contact in between the same dies, one talk about Single Pass or unidirectional Multiple Pass Strip Drawing (SPSD or MPSD). SPSD is use to simulate damage mechanisms relative to cumulative build-up of sur-
Fig. 5 – Contact equivalence in strip drawing between flat/flat and cylinder/cylinder geometry.
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Fig. 6 – Influence of axial tensile polarization (L). (a) Friction response vs. sliding distance and (b) surface roughness after friction test.
face flaws eventually leading to galling while MPSD aims at simulation of sheet surface and friction force modifications of a single blank area suffering repeated frictional contact at different parts of the stamping dies (Fig. 4). But Roizard and von Stebut (1993, 1995) and Roizard et al. (1999) have been clearly demonstrated in previous papers that the dynamic friction coefficient cannot be considered to be the unique characteristic parameter used to discriminate the behaviour of lubricated metal sheets in sliding friction on hard steel tools. Indeed, a relatively low-friction coefficient may mask important damage such as scratches that occur just before galling. So, it is important to combine friction measurements with 3D profilometry before and after friction – relocated with high precision – to allow analysing of the specific residual roughness and the friction-induced surface damage on the load carrying
plateaus. Whereas the dynamic friction coefficient remains stable vs. pass number, the residual roughness decrease drastically (Fig. 4). This explains the static friction peak growth, due to lack of lubricant in store while the real contact area increases. Within this scope, it is possible to develop suitable 3D surface analysis parameters for the description of metal tribological behaviour. Moreover, it is important to note that if the choice of cylinder/cylinder contact geometry could appear surprising in a first glance for press shop simulations, in previous works Guillon et al. (2001) have shown that MPSD with cylindrical dies leads to equivalent sheet surface damage and friction as compared to flat/flat contact geometry, provided that identical apparent contact pressures and identical integrated sliding lengths are chosen (Fig. 5).
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Fig. 7 – Multi-pass strip drawing of uncoated steel sheets as a function of contact temperature (25 or 100 ◦ C).
3.2. Independent control of die/sheet contact pressure and axial tensile load Without polarization load, the dynamic friction coefficient obtained for standard operating conditions (normal load 4 kN, sliding velocity 20 mm min−1 , steel sheet), the dynamic friction coefficient is about 0.17 while rising up top 0.22 by applying a polarization load L of 10 kN (Fig. 6a). The value of polarization load was chosen to give a bulk stress of 200 MPa which corresponds to the yield stress of steel. With the contact pressure being identical in all cases (normal force: P = 4 kN, leading to apparent pressure: pa = 60 MPa) this is clear-cut evidence of the importance of bulk plastic deformation for friction via a purely adhesive mechanism. This will favour sheet/die transfer which is one of the major origins of galling.
Jonasson et al. (1998) has used a bending under tension friction test in order to investigate the change in surface roughness of different sheet topographies and a relationship was determined by plotting the mean real area of the oil pockets vs. the coefficient of friction for a sheet material with an excess amount of lubricant. Kimura (1999) has studied how workpiece surface roughness is crushed by tooling in metal-forming processes and velocity fields was proposed which separately describe the bulk flow in plane strain compression and the crushing of asperities aligned along the direction of bulk flow. He showed that when superimposed, they describe asperity crushing on a bulk-flowing foundation. Fig. 6b shows that relative levelling down of the surface roughness is influenced too by the bulk stress: relative mean peak-to-valley height is more than two times smaller while applying an axial tensile load.
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When Kimura and Childs (1999) has described asperity crushing on a bulk-flowing foundation, real areas of contact were dependent of bulk strains. Therefore it is by no means sure that strip drawing experiments under pure elastic bulk deformation match up to sufficiently well press shop reality, and the application of a tensile load appears to be necessary for adequate simulative results.
3.3.
Heating the strip drawing dies
In cases of galvanized sheets, lower surface hardness related to higher relative temperature T/T0 (T0 being the fusion temperature) has been proposed to explain a higher degree of junction formation due to more severe surface shearing, in analogy with Milner’s model 1962 of cold-welding. So, such experiments at room temperature with respect to friction and sheet surface damage do not allow for realistic simulation of press shop behaviour where contact temperature may be considerably higher. So, MPSD tests were carried out with the same standard operating parameters (normal load 4 kN, sliding velocity 20 mm min−1 , and galvanized steel sheet) while controlling the dies temperature (25 and 100 ◦ C). Friction coefficients vs. displacement are shown in Fig. 7 for five successive passes combined with a 3D profilometry analysis. It clearly appears that the behaviour of the tool/sheet contact is influenced by the dies temperature and that for 100 ◦ C, galling is observed leading to a stick-slip phenomenon more and more marked. The major mechanism proposed to describe this more severe contact at high temperature is the decrease of lubricant viscosity leading to more chemical reactivity of the surface which enhanced junction formation. With a laboratory strip reduction test simulating the tribological conditions of an ironing process, Andreasen et al. (1997) were capable of simulating varying process conditions such tool temperature. He has quantified the onset of breakdown of the lubricant film and subsequent galling with tool temperature.
3.4.
Measure of the real displacement sheet/tool
In order to understand the mechanisms occurring at the interface, a good knowledge of the surface relative movements is needed. The measurement of the actuator alone is clearly not sufficient, because it is not representative of the real displacements as many phenomenons are not taken into account such as tribometer stiffness. Then, it is usual to use an inductive sensor LVDT—to obtain the relative movement sheet/die and to plot friction coefficients as a function of this displacement. The original potentiometric measurement was used as a complementary displacement measurement and results are shown in Figs. 8 and 9 for hot dip galvanized mild steels during strip drawing with high-speed steel dies (Z160CDV12). Both tool and/or sheet surfaces present different initial states (used or virgin). Within this scope, three signals were recorded (Fig. 8): • The drawing force, which give the friction coefficient, linked to the strength provide by the actuator to produce a relative sheet/tool displacement.
• The displacement measured by LVDT sensor called D1. • The displacement measured by potentiometric method called D2. The test A concerns a virgin blank sheet and new tools, polished with grade 1200 emery paper just before strip drawing. Fig. 8A-a gives the friction coefficient vs. actuator displacement (i.e. time for a constant velocity). Two different phases can be identified, a first one which shows a quick increase, generally called static phase and typical of the elasticity of the system (tribometer stiffness, etc.) and a second which describes the dynamic friction phase. The same signal is plotted vs. D1 (Fig. 8A-b) and D2 (Fig. 8A-c). Both methods allow suppressing the static phase due to the device elasticity and show the same global behaviour. This is confirmed by plotting the displacement measured by LVDT vs. the displacement measured by potentiometric method in Fig. 8A-d. Then, the same experiment was carried out by means of a virgin blank sheet and a tool used in previous tests (three passes) and which presents metal pick-up at its surface (test B). Like in the test A, the behaviour of the friction coefficient vs. actuator displacement show a static peak and a dynamic friction phase, with a more stable plateau at a higher level ( = 0.25). This is characteristic of friction between the same material due to the metal pick-up on tool surface (Fig. 8B-a). The beginning of the dynamic part is not easy to detect, because of a break presence discernable in the initial slope. This break point is the beginning of dynamic friction phase in the case of displacement measured by the help of the potentiometric method (Fig. 8B-c) whereas it starts after if using LVDT (Fig. 8B-b). This difference could be quantified by plotting the displacement given by the LVDT vs. the ones given by potentiometric sensors by an offset of 300 m (Fig. 8B-d). After this initial offset, any difference could be seen between the displacement measurement methods. In test C, at the contrary than in test B, the measurements are made by the help of new tools, polished with grade 1200 emery paper just before strip drawing and fore used sheets, ironed down roughness plateaux by three passes. As expected, a transitory phase appears important before the steady-state dynamic behaviour corresponding to the build-up of a metal pick-up on the tool surface, visible for both D1 and D2 records (Fig. 9C-b and C-c). One again, a gap appears between the beginnings of the dynamic phase as it is measured by LVDT or potentiometric methods and can be quantified in Fig. 9Cd. The value obtained is a little bit smaller (150 m), but of the same sign. Finally, experiments are carried out by the help of a tool used in previous tests (three passes) combined with fore used sheets (ironed down roughness plateaux by three passes) and the results are shown in Fig. 9D. The transitory dynamic phase is not more present (because of the initial presence of transfer pick-up on tools), and the global behaviour is close to the one observed in Fig. 8B. The beginning of the dynamic part is once again not easy to detect. The beginning of dynamic friction phase in the case of displacement measured by the help of the potentiometric method (Fig. 9D-c) starts after if using LVDT (Fig. 9D-b). The quantification of this difference by plotting the displacement given by the LVDT vs. the one‘s given by potentiometric sensors give an offset of 450 m (Fig. 9D-d). It is important to notice that this value correspond to the sum of the offset in the case where just the tool was
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Fig. 8 – Results of tests A and B: (a) evolution of the friction coefficient () with respect to time (sampling rate: 100 Hz), (b) evolution of with respect to the measured displacement D1 (LVDT), (c) evolution of with respect to the real displacement D2 (potentiometer) and (d) evolution of D1 vs. D2.
used (Fig. 8B-d) and the case where just the sheet was ironed (Fig. 9C-d). In all cases, the displacements of the tools measured by both LVDT and potentiometric sensors begin after the tribometer actuator move, due to the relatively high device compliance (0,7 m N−1 ), i.e. the actuator is moving before an actual sheet/tools displacements. This actual sheet/tools moving takes place when the spring force of the system is
superior to the shear resistance of the created junctions. Then, all offsets obtained by plotting D1 vs. D2 displacement (Fig. 8Ad, B-d, C-d and D-d) allows to quantify the plastic deformations of the adhesive junctions created during the static phase. Indeed, higher the distance for which the signal D2 records no variation is, higher the deformations are, underlining strongly adhesive and ductile connections at the sheet steel/tools interface.
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Fig. 9 – Results of tests C and D: (a) evolution of the friction coefficient () with respect to time (sampling rate: 100 Hz), (b) evolution of with respect to the measured displacement D1 (LVDT), (c) evolution of with respect to the real displacement D2 (potentiometer) and (d) evolution of D1 vs. D2.
Due to its principle, the potentiometric method contains both real relative movement sheet/tool and local elasto-plastic deformations of the interface. So, the comparison with LVDT method allows discriminating real sliding from plastic deformations of the interface. Berthier et al. (1988) has shown that in “dry” friction, velocity can be accommodated through 20 different mechanisms (5 sites and 4 modes per site) known as velocity accommodation mechanisms. So potentiometric method is very interesting in order to allow a better understanding of the velocity accommodation mechanisms between sheet and tool.
For instance, static peak is linked to the shear resistance of adhesive junctions, which are proportional to the metal transfer on the tool (mode M2, i.e. shearing—location S2, i.e. tool screen as mentioned by Berthier et al. (1989)). The evolution of friction coefficient can be discriminate in three successive phases: specific to the own tribometer properties, specific to the surface history and specific to steady-state dynamic phase (Fig. 10). Thanks to the methodology and measurement methods proposed here, it is possible to make their quantifications:
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rial and/or tests made at room temperature is likely to be insufficient. Finally, an innovating method for displacement measurements has been developed, using the property of conductivity for metals. This method allows the analyses of the local displacement of the tool in relation to the metal sheet. By comparison with an apparent displacement measurement of the tool in relation to the clear tip of the metal sheet (by a LVDT sensor), it is indeed possible to identify both tribometer compliance and mechanical properties of adhesive junctions. This will lead to further determination of tool/sheet velocity accommodation mechanisms.
references
Fig. 10 – Dynamic measurements: D1, LVDT macroscopic sheet/tool displacement measurement; D2, potentiometric real sheet/tool displacement measurement; force measurement given by the traction test rig allowing the relative displacement between the sheet and the tool to be done.
• the drawing force linked to the actuator gives the relative displacement sheet/tool, allowing the tribometer compliance determination, • the displacement measurement by LVDT sensor gives the macroscopic sheet/tool displacement, called D1, corresponding to the effective sliding, • the potentiometric measurement gives the beginning of sheet/tool displacement, called D2, containing both junctions shearing and effective sliding.
4.
Conclusions
In this paper, a strip drawing test is described in order to simulate sheet/tool contact conditions during a stamping operation. Its novelty consists in proposing both new equipments and methodological aspects. Multi-pass strip drawing tests with a cylinder/sheet/cylinder contact geometry, associated to specific measurement equipments, allows emphasizing the influence of factors rarely taken into account in experimental simulation. It has been shown the strong evidence of adhesive friction due to temperature contact and bulk plastic deformation of the strip drawing metal sheet and its consequences with respect to ironing roughness. The levelling down of the sheet surface during the repeated ironing action of the tools is more pronounced. The general friction level in the presence of bulk plastic deformation or during tests carry out within realistic contact temperature is about 20% higher. So, for realistic simulation of deep drawing by strip drawing, purely elastic deformation of the bulk mate-
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journal homepage: www.elsevier.com/locate/jmatprotec
Experimental device for tribological measurement aspects in deep drawing process X. Roizard a,∗ , J.M. Pothier a,b , J.Y. Hihn b , G. Monteil a a b
Laboratoire de Micro-Analyse des Surfaces (LMS), ENSMM, 26 chemin de l’´epitaphe, 25030 Besanc¸on, France Laboratoire de Chimie des Mat´eriaux et Interfaces (LCMI), IUT, rue de l’Observatoire, 25000 Besanc¸on, France
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Article history:
The purpose of this paper is to improve strip drawing test in order to simulate sheet/tool
Received 5 March 2007
contact conditions during a stamping operation. An original multi-pass strip drawing test
Received in revised form
in cylinder/sheet/cylinder contact geometry is presented and we emphasize the influence
10 March 2008
of certain factors rarely taken into account in the experimental simulation. In particular, we
Accepted 14 March 2008
present examples where the bulk plastic deformation and the tools average temperature influence the level of friction and the sheet surface asperities flattening. To try to readjust the friction evolutions with the modifications of surface microgeometry, the tribometer is
Keywords:
equipped with various measurements of relative sheet/tool displacement, including or not
Deep drawing
the stiffness machine and the stiffness of the contact. This was made possible by a poten-
Tribology
tiometric measurement where the tool plays the part of potentiometer and sheet that of the
Strip drawing
track, coupled to a measurement using a LVDT sensor.
Temperature contact
© 2008 Elsevier B.V. All rights reserved.
Bulk deformation Adhesive junction properties
1.
Introduction
Deep drawing is a forming process for metal sheets frequently used in the automotive industry. The conditions in which the contact takes place are especially severe: boundary lubrication in pure sliding friction, high contact pressure always implying both surface asperity and bulk plastic deformation, cyclic progressive transfer layer build-up on the dies, etc. Previously, most of the research works has been focused on the experimental and numerical simulation of the contact conditions during this operation. Many researchers – with comparisons between FEM simulations and experimental results for stamping operations – showed that friction strongly influences the formability by affecting the strain distribution in various regions of the sheet: Hortig and Schmoeckel (2001) have shown that a small friction coefficient does not always lead
∗
Corresponding author. Tel.: +33 3 81 40 28 54. E-mail address: [email protected] (X. Roizard). 0924-0136/$ – see front matter © 2008 Elsevier B.V. All rights reserved. doi:10.1016/j.jmatprotec.2008.03.023
to lower tribological loads, however Choudhury et al. (2006) have observed that, with the increase of friction coefficient the blank holder force decreases, Verma and Chandra (2006) concludes that the limiting drawing ratio (LDR) decreases with increase in the friction. On the other hand, Sutcliffe et al. (2003) and Matuszak (2000) have shown that the lubrication regime is critical for formation of the transfer layer. Therefore, Bech et al. (1999), using a transparent tool entrapment, has observed that the smaller the ratio of the estimated lubricant film thickness to the roughness is, the higher the transfer layer. Since Bowden and Tabor, a common opinion in tribology is that the pinning of the surfaces is due to cold-welded junction’s formation, so that the sliding corresponds to these junctions’ shearing. Due to economic and environmental pressures, industry is led to optimize all operating parameters by using several types of tribometers, reproducing more or less one part of
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the drawn sheet. Strip drawing test is one of these techniques and allows different tool geometries: flat/flat used by Kim et al. (2002), radial strip drawing used by Sniekers and Smits (1997), cylinder/cylinder used by Roizard and von Stebut (1995), etc. Unfortunately, most of strip drawing experiments can only simulate plastic deformation on the bulk material via increased surface pressure at the die/sheet interface: see for instance when friction tests of Wihlborg and Gunnarsson (2000) were performed in a Bending Under Tension test. To separate the bend—unbend force from the pure friction component, Sulonen (1981) has proposed a special hydraulic test rig. But even in this case, the overall bulk plastic deformation and the contact pressure applied across the surface asperities remain coupled parameters. Moreover, those tests are generally carried out at room temperature whereas the real contact temperatures rise up to 100 ◦ C after stamping of a series of sheets in industrial press shop owing to frictional self-heating. Recently, Novotny (2003) classified the suitability of lubricants and tool materials with a strip drawing test equipped with a heating device. It had been shown that fluid lubricants generate low-friction forces at room and elevated temperature if their viscosity at room temperature is high enough. This confirms the needs of tests showing the maximum of mechanical contacts similarities with industrial press shops, while keeping independence in the control of input parameters. This will make possible to optimize technological choices for a given metal sheet, as for example lubricant nature and quantity, tool material as well as allowing a better comprehension of physical behaviour of the contact. In this objective, it is possible to improve the information given by our device by improving the device equipment (heating tools, sample/tool positioning and bulk pre stressing) and by improving test methodology (multi-pass, profilometric relocalisation). In the same way, it is required to discriminate the behaviour of the contact and the device itself. This is particularly true for some experimental conditions, when the so-called stick-slip effect appears. Both experiments and simulations show that in all cases, when the static frictional is nonzero, the system exhibits a stick-slip motion at low velocities. Schematically in the stick-slip sliding regime, friction force as a function of time occurs shown a typical shape presented in Fig. 1a. In classic tribological experiments, a spring is attached to the slider, and its end connected to a base which moves with a constant velocity Vspring (Fig. 1a). The same is true for strip drawing, where the elasticity of the moving parts of the device plays the same role. The output (measured) parameter is the spring force Fspring . Let us assume that initially the system is in rest and the spring has its natural length. When the base begins to move, the spring stretches, F increases until it reaches the threshold value Fs corresponding to the static frictional force, and the block starts to move. Then, Braun and Naumovets (2006) showed that, due to inertia, the slider accelerates to catch the base. If Vspring is small, F decreases down to the “backward” threshold force Fb , and the slider stops. Then the process repeats, so the stick-slip motion occurs as shown in Fig. 1b. To understand the complete mechanisms of this phenomenon, i.e. accede to the mechanical properties of the junctions between dies and sheet, it is necessary to obtain the true relative displacement between parts. That is why our tribometer is equipped with an original measure of relative sheet/tool displacement.
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Fig. 1 – (a) A standard experimental setup for tribology studies and (b) friction force as a function of time in the stick-slip regime (schematically).
The aim of the present work is to present the experimental possibilities of an original device dedicated to tribological measurement aspects by keeping the maximum of similarities with industrial press shop, whereas allowing a better understanding of friction mechanisms by individual control and specific measurement methods.
2. Experimental details—setup and measurements techniques The test rig adopted is equipped with a cylinder/cylinder contact geometry very convenient for tool/sheet alignment and friction data interpretation. Its specific feature is to allow perfectly independent control of contact pressure and bulk tensile stress as well as instantaneous die/sheet sliding speed and contact temperature (Fig. 2). The actuator of a MTS® testing machine generates the relative sliding motion of the sheet with respect to both dies. The deformation induced by the tools may be qualified as macro-elastic as only the roughness is affected in a plastic manner. A particular attention was given to monitor and control mechanical parameters, i.e. to control independently contact pressure and axial tensile load along the S.D. sheet. This last one is placed between grips on a special load frame which applies a constant polarization load L to the sheet’s ends by means of rubber bellows. Because of friction, the part of the sheet on the side opposite to the dies sliding direction will be under a longitudinal tensile load equal to (L + 2 P), where is the friction coefficient and L the tensile polarization load. As a complement, our strip drawing test rig allows die temperature control ensured by means of heating coils inserted into the die grips and powered by means of a P.I.D. unit in conjunction with a platinum probe. The relative sheet/die displacement is given by a LVDT sensor, completed by an original electrical measurement in order to take into account the real displacement. Indeed, this local
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Fig. 2 – Synoptic description of strip drawing tribometer.
displacement of the tool vs. sheet is based on a potentiometric method, whose cursor and track are constituted by tool and work sheet (Fig. 3). This gives a measure as close as possible of phenomena occurring at the interface. The principle of this potentiometric method is based on the creation of a constant voltage gradient, by injecting a constant current into the constant section sample sheet. The measure itself consists in the acquisition of the difference of potential between the moving tools and the fixed electrodes. This measure is obtained through a high impedance loop so that the current passing through the tool/sheet interface is lower than 10−12 A mm−2 and allows neglecting any thermal or tribo-oxydation effects. Then, the highest resolution for this method implies the highest value for the voltage gradient, because the displacement is directly proportional to this voltage. However, the increase of voltage gradient is obtained from the increase of the injected current, while staying below the threshold of the heating by
Joule effect. As an example in the case of aluminium, the resistivity is around 2.6 × 10−8 m and for 50 mm2 section sheets, the range of magnitude of the electric resistance is about 0.5 m m−1 . So, an injection current of 50 A induce a power to be dissipated lower than 25 W m−2 . In those conditions, the experimental temperature variation never exceeds 1 ◦ C. As the potential gradient is 26 mV m−1 measured through an instrumentation amplifier with gain 3000 and that the resolution of our acquisition card is 150 V, the resolution in distance is around 2 m. Furthermore, the higher the sample resistivity is and the more the accuracy increase up to 0.8 m for steel. In order to keep the maximum of precision and repeatability in all measurements, it is necessary to respect a precise geometrical location of electrodes, in particular to insure them a “perfect” perpendicularity with the lines of electric field. This location is done by means of a template for standardization of the electrode positions. As theoretical calculations led with the resistivities of pure metals do not allow defining an exact correspondence between the measured tension and the relative position of the tool/sheet, a calibration electrode has been added for which the distance with the reference electrode is perfectly known thanks to the template. Samples used in the present work consist of low carbon steels sheets (50 mm × 300 mm × 0.7 mm) and aluminium EDT 5182 (50 mm × 300 mm × 1 mm). QUAKER® 6130 was applied as lubricant for each sample at 2 g m−2 . The velocity of the tensile machine actuator was kept quite to low values between 0.1 and 33 mm s−1 , in order to remain in boundary lubrication regime and to avoid the risk of galling. Cylindrical tools (diameter 20 mm, length 60 mm) are made of uncoated high-speed steel Z85WD06, with roughness lower than 0.1 m (arithmetic average roughness, Ra). Otherwise, the length of the sliding track has been limited to 3 mm, which is enough to systematically ensure stabilized functioning. The variables acquired during a test are sampled within a frequency of 200 Hz. They include: • The normal force FN (N) applied to the metal sheet by the tools. • The displacement (mm) of the actuator of the tensile machine, which produces a motion of the tools in relation to the metal sheet.
Fig. 3 – Potentiometric measure of sheet/tool relative displacement.
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Fig. 4 – Illustration of the multi-pass strip drawing (MPSD). Topography and friction coefficient measurements on the same scanned area of the metal sheet between two consecutive passes.
• The tangential stretching force FT (N), transmitted by the actuator of the tensile machine to generate motion. • The apparent displacement (mm) of the tool in relation to the clear tip of the work sheet. Measurement of this displacement should allow identification of tribometer compliance, presently estimated at 0.5 m/N. • The local displacement (mm) of the tool in relation to the work sheet. This value is based on the potentiometric method.
3.
Results
3.1. Multi-pass strip drawing and 3D profilometry measurement Depending on the number of times the same sheet surface area goes through frictional contact in between the same dies, one talk about Single Pass or unidirectional Multiple Pass Strip Drawing (SPSD or MPSD). SPSD is use to simulate damage mechanisms relative to cumulative build-up of sur-
Fig. 5 – Contact equivalence in strip drawing between flat/flat and cylinder/cylinder geometry.
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Fig. 6 – Influence of axial tensile polarization (L). (a) Friction response vs. sliding distance and (b) surface roughness after friction test.
face flaws eventually leading to galling while MPSD aims at simulation of sheet surface and friction force modifications of a single blank area suffering repeated frictional contact at different parts of the stamping dies (Fig. 4). But Roizard and von Stebut (1993, 1995) and Roizard et al. (1999) have been clearly demonstrated in previous papers that the dynamic friction coefficient cannot be considered to be the unique characteristic parameter used to discriminate the behaviour of lubricated metal sheets in sliding friction on hard steel tools. Indeed, a relatively low-friction coefficient may mask important damage such as scratches that occur just before galling. So, it is important to combine friction measurements with 3D profilometry before and after friction – relocated with high precision – to allow analysing of the specific residual roughness and the friction-induced surface damage on the load carrying
plateaus. Whereas the dynamic friction coefficient remains stable vs. pass number, the residual roughness decrease drastically (Fig. 4). This explains the static friction peak growth, due to lack of lubricant in store while the real contact area increases. Within this scope, it is possible to develop suitable 3D surface analysis parameters for the description of metal tribological behaviour. Moreover, it is important to note that if the choice of cylinder/cylinder contact geometry could appear surprising in a first glance for press shop simulations, in previous works Guillon et al. (2001) have shown that MPSD with cylindrical dies leads to equivalent sheet surface damage and friction as compared to flat/flat contact geometry, provided that identical apparent contact pressures and identical integrated sliding lengths are chosen (Fig. 5).
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Fig. 7 – Multi-pass strip drawing of uncoated steel sheets as a function of contact temperature (25 or 100 ◦ C).
3.2. Independent control of die/sheet contact pressure and axial tensile load Without polarization load, the dynamic friction coefficient obtained for standard operating conditions (normal load 4 kN, sliding velocity 20 mm min−1 , steel sheet), the dynamic friction coefficient is about 0.17 while rising up top 0.22 by applying a polarization load L of 10 kN (Fig. 6a). The value of polarization load was chosen to give a bulk stress of 200 MPa which corresponds to the yield stress of steel. With the contact pressure being identical in all cases (normal force: P = 4 kN, leading to apparent pressure: pa = 60 MPa) this is clear-cut evidence of the importance of bulk plastic deformation for friction via a purely adhesive mechanism. This will favour sheet/die transfer which is one of the major origins of galling.
Jonasson et al. (1998) has used a bending under tension friction test in order to investigate the change in surface roughness of different sheet topographies and a relationship was determined by plotting the mean real area of the oil pockets vs. the coefficient of friction for a sheet material with an excess amount of lubricant. Kimura (1999) has studied how workpiece surface roughness is crushed by tooling in metal-forming processes and velocity fields was proposed which separately describe the bulk flow in plane strain compression and the crushing of asperities aligned along the direction of bulk flow. He showed that when superimposed, they describe asperity crushing on a bulk-flowing foundation. Fig. 6b shows that relative levelling down of the surface roughness is influenced too by the bulk stress: relative mean peak-to-valley height is more than two times smaller while applying an axial tensile load.
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When Kimura and Childs (1999) has described asperity crushing on a bulk-flowing foundation, real areas of contact were dependent of bulk strains. Therefore it is by no means sure that strip drawing experiments under pure elastic bulk deformation match up to sufficiently well press shop reality, and the application of a tensile load appears to be necessary for adequate simulative results.
3.3.
Heating the strip drawing dies
In cases of galvanized sheets, lower surface hardness related to higher relative temperature T/T0 (T0 being the fusion temperature) has been proposed to explain a higher degree of junction formation due to more severe surface shearing, in analogy with Milner’s model 1962 of cold-welding. So, such experiments at room temperature with respect to friction and sheet surface damage do not allow for realistic simulation of press shop behaviour where contact temperature may be considerably higher. So, MPSD tests were carried out with the same standard operating parameters (normal load 4 kN, sliding velocity 20 mm min−1 , and galvanized steel sheet) while controlling the dies temperature (25 and 100 ◦ C). Friction coefficients vs. displacement are shown in Fig. 7 for five successive passes combined with a 3D profilometry analysis. It clearly appears that the behaviour of the tool/sheet contact is influenced by the dies temperature and that for 100 ◦ C, galling is observed leading to a stick-slip phenomenon more and more marked. The major mechanism proposed to describe this more severe contact at high temperature is the decrease of lubricant viscosity leading to more chemical reactivity of the surface which enhanced junction formation. With a laboratory strip reduction test simulating the tribological conditions of an ironing process, Andreasen et al. (1997) were capable of simulating varying process conditions such tool temperature. He has quantified the onset of breakdown of the lubricant film and subsequent galling with tool temperature.
3.4.
Measure of the real displacement sheet/tool
In order to understand the mechanisms occurring at the interface, a good knowledge of the surface relative movements is needed. The measurement of the actuator alone is clearly not sufficient, because it is not representative of the real displacements as many phenomenons are not taken into account such as tribometer stiffness. Then, it is usual to use an inductive sensor LVDT—to obtain the relative movement sheet/die and to plot friction coefficients as a function of this displacement. The original potentiometric measurement was used as a complementary displacement measurement and results are shown in Figs. 8 and 9 for hot dip galvanized mild steels during strip drawing with high-speed steel dies (Z160CDV12). Both tool and/or sheet surfaces present different initial states (used or virgin). Within this scope, three signals were recorded (Fig. 8): • The drawing force, which give the friction coefficient, linked to the strength provide by the actuator to produce a relative sheet/tool displacement.
• The displacement measured by LVDT sensor called D1. • The displacement measured by potentiometric method called D2. The test A concerns a virgin blank sheet and new tools, polished with grade 1200 emery paper just before strip drawing. Fig. 8A-a gives the friction coefficient vs. actuator displacement (i.e. time for a constant velocity). Two different phases can be identified, a first one which shows a quick increase, generally called static phase and typical of the elasticity of the system (tribometer stiffness, etc.) and a second which describes the dynamic friction phase. The same signal is plotted vs. D1 (Fig. 8A-b) and D2 (Fig. 8A-c). Both methods allow suppressing the static phase due to the device elasticity and show the same global behaviour. This is confirmed by plotting the displacement measured by LVDT vs. the displacement measured by potentiometric method in Fig. 8A-d. Then, the same experiment was carried out by means of a virgin blank sheet and a tool used in previous tests (three passes) and which presents metal pick-up at its surface (test B). Like in the test A, the behaviour of the friction coefficient vs. actuator displacement show a static peak and a dynamic friction phase, with a more stable plateau at a higher level ( = 0.25). This is characteristic of friction between the same material due to the metal pick-up on tool surface (Fig. 8B-a). The beginning of the dynamic part is not easy to detect, because of a break presence discernable in the initial slope. This break point is the beginning of dynamic friction phase in the case of displacement measured by the help of the potentiometric method (Fig. 8B-c) whereas it starts after if using LVDT (Fig. 8B-b). This difference could be quantified by plotting the displacement given by the LVDT vs. the ones given by potentiometric sensors by an offset of 300 m (Fig. 8B-d). After this initial offset, any difference could be seen between the displacement measurement methods. In test C, at the contrary than in test B, the measurements are made by the help of new tools, polished with grade 1200 emery paper just before strip drawing and fore used sheets, ironed down roughness plateaux by three passes. As expected, a transitory phase appears important before the steady-state dynamic behaviour corresponding to the build-up of a metal pick-up on the tool surface, visible for both D1 and D2 records (Fig. 9C-b and C-c). One again, a gap appears between the beginnings of the dynamic phase as it is measured by LVDT or potentiometric methods and can be quantified in Fig. 9Cd. The value obtained is a little bit smaller (150 m), but of the same sign. Finally, experiments are carried out by the help of a tool used in previous tests (three passes) combined with fore used sheets (ironed down roughness plateaux by three passes) and the results are shown in Fig. 9D. The transitory dynamic phase is not more present (because of the initial presence of transfer pick-up on tools), and the global behaviour is close to the one observed in Fig. 8B. The beginning of the dynamic part is once again not easy to detect. The beginning of dynamic friction phase in the case of displacement measured by the help of the potentiometric method (Fig. 9D-c) starts after if using LVDT (Fig. 9D-b). The quantification of this difference by plotting the displacement given by the LVDT vs. the one‘s given by potentiometric sensors give an offset of 450 m (Fig. 9D-d). It is important to notice that this value correspond to the sum of the offset in the case where just the tool was
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Fig. 8 – Results of tests A and B: (a) evolution of the friction coefficient () with respect to time (sampling rate: 100 Hz), (b) evolution of with respect to the measured displacement D1 (LVDT), (c) evolution of with respect to the real displacement D2 (potentiometer) and (d) evolution of D1 vs. D2.
used (Fig. 8B-d) and the case where just the sheet was ironed (Fig. 9C-d). In all cases, the displacements of the tools measured by both LVDT and potentiometric sensors begin after the tribometer actuator move, due to the relatively high device compliance (0,7 m N−1 ), i.e. the actuator is moving before an actual sheet/tools displacements. This actual sheet/tools moving takes place when the spring force of the system is
superior to the shear resistance of the created junctions. Then, all offsets obtained by plotting D1 vs. D2 displacement (Fig. 8Ad, B-d, C-d and D-d) allows to quantify the plastic deformations of the adhesive junctions created during the static phase. Indeed, higher the distance for which the signal D2 records no variation is, higher the deformations are, underlining strongly adhesive and ductile connections at the sheet steel/tools interface.
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Fig. 9 – Results of tests C and D: (a) evolution of the friction coefficient () with respect to time (sampling rate: 100 Hz), (b) evolution of with respect to the measured displacement D1 (LVDT), (c) evolution of with respect to the real displacement D2 (potentiometer) and (d) evolution of D1 vs. D2.
Due to its principle, the potentiometric method contains both real relative movement sheet/tool and local elasto-plastic deformations of the interface. So, the comparison with LVDT method allows discriminating real sliding from plastic deformations of the interface. Berthier et al. (1988) has shown that in “dry” friction, velocity can be accommodated through 20 different mechanisms (5 sites and 4 modes per site) known as velocity accommodation mechanisms. So potentiometric method is very interesting in order to allow a better understanding of the velocity accommodation mechanisms between sheet and tool.
For instance, static peak is linked to the shear resistance of adhesive junctions, which are proportional to the metal transfer on the tool (mode M2, i.e. shearing—location S2, i.e. tool screen as mentioned by Berthier et al. (1989)). The evolution of friction coefficient can be discriminate in three successive phases: specific to the own tribometer properties, specific to the surface history and specific to steady-state dynamic phase (Fig. 10). Thanks to the methodology and measurement methods proposed here, it is possible to make their quantifications:
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rial and/or tests made at room temperature is likely to be insufficient. Finally, an innovating method for displacement measurements has been developed, using the property of conductivity for metals. This method allows the analyses of the local displacement of the tool in relation to the metal sheet. By comparison with an apparent displacement measurement of the tool in relation to the clear tip of the metal sheet (by a LVDT sensor), it is indeed possible to identify both tribometer compliance and mechanical properties of adhesive junctions. This will lead to further determination of tool/sheet velocity accommodation mechanisms.
references
Fig. 10 – Dynamic measurements: D1, LVDT macroscopic sheet/tool displacement measurement; D2, potentiometric real sheet/tool displacement measurement; force measurement given by the traction test rig allowing the relative displacement between the sheet and the tool to be done.
• the drawing force linked to the actuator gives the relative displacement sheet/tool, allowing the tribometer compliance determination, • the displacement measurement by LVDT sensor gives the macroscopic sheet/tool displacement, called D1, corresponding to the effective sliding, • the potentiometric measurement gives the beginning of sheet/tool displacement, called D2, containing both junctions shearing and effective sliding.
4.
Conclusions
In this paper, a strip drawing test is described in order to simulate sheet/tool contact conditions during a stamping operation. Its novelty consists in proposing both new equipments and methodological aspects. Multi-pass strip drawing tests with a cylinder/sheet/cylinder contact geometry, associated to specific measurement equipments, allows emphasizing the influence of factors rarely taken into account in experimental simulation. It has been shown the strong evidence of adhesive friction due to temperature contact and bulk plastic deformation of the strip drawing metal sheet and its consequences with respect to ironing roughness. The levelling down of the sheet surface during the repeated ironing action of the tools is more pronounced. The general friction level in the presence of bulk plastic deformation or during tests carry out within realistic contact temperature is about 20% higher. So, for realistic simulation of deep drawing by strip drawing, purely elastic deformation of the bulk mate-
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