Akkerman, Remko

de Rooij, Matthias B.

Sachs, Ulrich

ten Thije, R.H.W.

English

A previously developed mesoscopic friction model for glass/PP textile composite laminates during forming is evaluated for glass and carbon/PPS laminates, at higher temperatures and lower viscosities than before. Experiments were performed for tool/ply and ply/ply configurations in a new friction test set-up. The experimental results indicate that this contact is indeed in the hydrodynamic regime. The model results are sensitive to inaccuracies in the geometric representation of the textile structure, whereas the experimental results are sensitive to even slight misalignments

University of Twente Research Information

C. Binetruy, F. Boussu (Eds.) - Recent Advances in Textile Composites, 2010 p.271-279       Friction in Textile Thermoplastic Composites Forming   R. AKKERMAN, R. TEN THIJE, U. SACHS and M. DE ROOIJ     ABSTRACT  A previously developed mesoscopic friction model for glass/PP textile composite laminates during forming is evaluated for glass and carbon/PPS laminates, at higher temperatures and lower viscosities than before. Experiments were performed for tool/ply and ply/ply configurations in a new friction test set-up. The experimental results indicate that this contact is indeed in the hydrodynamic regime. The model results are sensitive to inaccuracies in the geometric representation of the textile structure, whereas the experimental results are sensitive to even slight misalignments.    INTRODUCTION  Textile reinforced thermoplastics offer a cost reduction compared to their thermoset counterparts due to promising fast production methods such as diaphragm forming and press forming, usually starting from pre-consolidated laminates. Friction plays an important role in these forming processes [1]. The constraints imposed by friction between subsequent plies and between the laminate and the tools are a major factor in the laminate deformations (such as wrinkling and folding) generated during composite forming. This friction depends on the forming process parameters such as the pressure, the temperature and the sliding velocity. In addition, it depends on the material properties of the fibres and the resin, the fibre distribution and the reinforcement architecture, as for any composite property.  A model was developed to describe the frictional phenomena in the general case, taking into account the effects of the various governing parameters [2]. While previous studies have resulted in empirical models [3,4,5,6,7] or report experimentally obtained values [8], this model predicts the friction between thermoplastic laminates and a steel tool by assuming hydrodynamic lubrication on a meso-mechanical level. Thus, the frictional properties can be calculated, solely based on the rheological properties of the matrix constituent and the fabric weave geometry. The model was validated experimentally with a novel pull-through friction tester, in which a laminate is pulled at constant velocity, while clamped by two stiff blocks at C. Binetruy, F. Boussu (Eds.) - Recent Advances in Textile Composites, 2010 p.271-279 processing temperature. Force displacement diagrams are generated from these pull-through experiments (Fig. 1). These diagrams typically show an overshoot after which (in general) the friction force attains a steady state value. The steady state forces can be used to determine a friction coefficient, which usually increases with velocity and decreases with temperature and pressure. The model predictions in general agree well with our experimental results. Our results on glass-PP Twintex fabrics (around 200ºC) compare well with Harrison et al [9], but opposite trends were found by Lebrun et al [8]. This has been the main reason for starting a benchmark exercise on this particular topic, initiated at the Esaform 2010 conference.  Further measurements were performed on glass/PPS and on carbon/PPS Cetex laminates (around 300ºC). The lower viscosity leads to lower predicted film thicknesses, approaching the limits of hydrodynamic lubrication. Here, we will discuss the results and highlight the limitations of our approach.    FRICTION MODEL  A meso-scale model was developed [9, 10] based on a geometrical description of the tows within the fabric. One of the advantages of the model is that the film thickness can be predicted from the normal pressure and velocity. This avoids the use of some arbitrary thickness of this lubrication film.  Figure 2 presents a schematic cross section of the composite material. Hydrodynamic lubrication is assumed between the bundles and the tool surface. The total friction force per unit width follows by integrating the surface shear stresses over the length of the cross section, disregarding the bundle curvatures out of the plane of this cross section for the time being. The contributions of the longitudinal warp and transverse weft yarns can be analysed separately and added up to the total friction force.  Figure 2. Schematic cross section of a 2x2 Twill ply on a tool surface. FnFf , U pull out length [mm]Force [N]Force [N] Figure 1. Pull-through experiment (left) and typical force-displacement graph (right). transverse bundle longitudinal bundlematrixtoolC. Binetruy, F. Boussu (Eds.) - Recent Advances in Textile Composites, 2010 p.271-279     Figure 3. Schematic pressure distribution underneath a bundle. The Reynolds’ equation describes the relation between the pressure and thickness distributions in thin film lubrication. The simple one dimensional steady state situation is given by 36 .h p hUx x xη ∂ ∂ ∂= ∂ ∂ ∂  (1) A Cross-WLF viscosity model was used to characterise the stationary shear viscosity [ 11]. The pressure distribution can be solved for a given film thickness distribution using the following boundary conditions (see Figure 3). ( ) ( ) ( )0 00; 0; 0;pp L p x xx∂− = = =∂  (2) where the pressures are assumed to be non-negative due to cavitation in the fluid.  Then, the bearing and friction force per unit width follow as, respectively, ( )0;xBLF p x dx−= ∫  ( )0 0.2x xfL Lh p UF x dx dxx hτ η− −∂= = +∂∫ ∫  (3a,b) The “1½D” mesoscopic model predicts the bearing and friction forces FB and Ff as a function of the input parameters temperature T, velocity U and minimum film thickness h0. The model was used inversely, iteratively adapting h0 such that the integrated bearing force over all fibres was equal to the prescribed normal load Fn. This procedure also leads to the integral pull-out force, which can be compared to the experimental results.   EXPERIMENTAL  We consider friction testing in a straight movement as depicted in Figure 1. The ply is clamped with a constant force Fn between two (tooling) blocks with a friction area A. This friction area reduces when a pull-out test is performed, whereas this area remains constant when a pull-through configuration is used, in which virgin material is being pulled between the blocks during testing. The constant sliding speed U is imposed in a single step when the test is started. The temperature T is kept constant during a test and the friction force Ff  is measured.  A pull-out set-up was developed at the University of Twente, in conjunction with the friction model as summarized above. Results were presented first in 2007 [ 11], C. Binetruy, F. Boussu (Eds.) - Recent Advances in Textile Composites, 2010 p.271-279 with good agreement between measurements on glass/PP and the model predictions, despite the simple 1½D modeling. However, when turning to other material systems e.g. PPS composites, we discovered less correlation between test and model results. In addition, we needed to go to higher temperatures than 300ºC for high end thermoplastics (PEEK, PEKK etc) than possible with the pull-out set-up. A critical evaluation revealed that lower viscosities of PPS lead to smaller film thicknesses (in the order of 10 µm), which in combination with small deflections of the friction tool can lead to erroneous results of the experiments [ 12]. A pull-out set-up further leads to the development of a pressure gradient during the test, which makes interpretation of the test results complex, as the laminate friction properties depend strongly on pressure, amongst other parameters. Additionally, temperature control is critical due to the nature of thermoplastic viscosities which typically vary exponentially with temperature. The critical evaluation resulted in a new design of a friction tester in a pull-through configuration.  The new friction tester set-up  A schematic design of the new friction tester is shown in Figure 4. The set-up operates in a standard mechanical test system. A flexible pneumatic actuator supplies the compressive load in a self-aligning system. Thick blocks minimize the tool deflection, whereas the overlapping edges are used to pre-heat the laminate before it enters the contact area. The spacing between the tool blocks is measured at four corners with micrometer accuracy and temperatures are measured in both blocks with multiple thermocouples. The laminate pressure is measured with three load cells. The set-up can be used for tool-ply and ply-ply friction experiments. Figure 5 shows a picture of the setup. Steel blocks are used with an area of 50x50mm and a thickness of over 40 mm. We observed that a clean surface needs to be used at every test; otherwise we find a gradual increase of friction. Hence a fresh disposable metal foil is used as a tool surface for all experiments. The set-up was designed to reduce the deviations in the pressure profile to less than 3% at a nominal pressure of 10 bars. The air actuator allows free movement for all 6 degrees of freedom. Some key values on the new set-up properties: • powered by 4 electrical cartridge heaters • heating from 20 to 400 °C in 10 minutes • max. laminate temperature variations during the experiment:  ± 0.5 °C at 200 °C up to ± 1.5 °C at 400 °C    Figure 4. Schematic representation of the new friction tester set-up. Figure 5. Picture of the new friction tester setup C. Binetruy, F. Boussu (Eds.) - Recent Advances in Textile Composites, 2010 p.271-279 • ‘Pyrotek’ ceramic insulation keeps the sensors and actuators below a temperature of 45 °C at an operating temperature of 400 °C. • range: 0.05 – 1.2 bars • initial accuracy before test is started:  0.005 bar Pressure • accuracy during test: ± 0.02 bar Velocity • velocity variations within 1%  Measurements with the new setup give results that significantly differ from the results obtained with the old setup. This is attributed mostly to the improved homogeneity of the pressure profile in the new setup [ 12]. For this paper, experiments were performed on PPS laminates with displacements in the warp direction. The test matrix is presented in Table 1. Materials 8H satin glass/PPS 5H satin carbon/PPS  Surface Orientation warp dominant (//) weft dominant (⊥)  Configuration tool/ply ply/ply  Temperature 290ºC 310ºC  Pressure 10 kPa 50 kPa  Velocity 20 mm/min 100 mm/min 500 mm/min Table 1. Experimental matrix for textile composite friction tests.   RESULTS AND DISCUSSION   P = 10 kPa P = 50 kPa T = 290ºC  00.20.40.60.811.21.420 100 500coeff. of friction   00.10.20.30.40.50.60.720 100 500coeff. of friction tool/ply  00.20.40.60.811.21.420 100 500pul l-out velocitycoeff. of friction  00.10.20.30.40.50.60.720 100 500coeff. of friction T = 310ºC ply/ply vs tool/ply  00.20.40.60.811.21.420 100 500pull-out velocity [mm/min]coeff. of friction  00.10.20.30.40.50.60.720 100 500pull-out velocity [mm/min]coeff. of frictionG/PPS ply/ ply ⊥G/PPS tool/ply ⊥C/PP S tool/ply  //C/PP S ply /ply // coeff. of friction Table 2. Summary of experimental results for steady state friction. G/PPS tool/ply //G/PPS tool/ply ⊥C/PPS tool/ply //C/PPS tool/ply ⊥C. Binetruy, F. Boussu (Eds.) - Recent Advances in Textile Composites, 2010 p.271-279 The steady state experimental results are summarized in Table 2. As was also found for glass/PP composites [ 2], it is observed that the coefficient of friction: • decreases when the pressure increases, • increases when the velocity increases and • (mostly) decreases when the temperature increases. The friction for 8HS glass/PPS is lower when the bundles at the surface are oriented predominantly transverse to the pulling direction. The coarser 5HS carbon fabric does not show this effect. The 290ºC ply/ply experiments failed, as the edges outside the tool adhered to each other below the melting temperature. In general, we observe slightly higher ply/ply friction than for the corresponding tool/ply configuration. Theoretically, we would expect a lower ply/ply friction, as more wedges are present. Possibly the edges sticking out of the tool artificially increase the measured friction.   Further, the experimental results indicate that the tool/ply friction is indeed in the hydrodynamic regime, as can be seen in the Stribeck plots presenting the coefficient of friction versus the Hersey number He = ηU/p. As an example, Figure 6 shows a constant positive slope for glass/PPS in the two principal pulling directions. The simple 1½D hydrodynamic lubrication model generally shows good agreement with the experimental results, as illustrated in Figure 7.  The G/PPS results in the fibre direction show the largest deviations. This is attributed to the fairly flat dominant contact region for this 8HS fabric, which is represented as an elliptical geometry in the friction model (compare to Figure 2).  Figure 8 presents exemplary measurement data. The coefficient of friction is fairly reproducible. The variation of h0 shows an increase of the film thickness in comparison to the state of rest. Since we do not know the film thickness in the state of rest it is not readily comparable with the simulations. From the measurements it is also observed that the self-aligning block tilts when the friction motion is initiated, and the 0.010.11101.E-06 1.E-05 1.E-04 1.E-03He [m]coeff. of friction [-]0.010.11101.E-06 1.E-05 1.E-04 1.E-03He [m]coeff. of friction [-]//⊥η U/p [m]η U/p [m] 00.20.40.60.811.21.41.61.8220 100 500pull-out velocity [mm/min]coeff. of friction [-]00.20.40.60.811.21.41.61.8220 100 500pull-out velocity [mm/min]coeff. of friction [-]expsim Figure 6. Stribeck plots for tool/ply G/PPS (all conditions), fibres predominantly in (top) and transverse to (bottom) the pulling direction. Figure 7. Experimental (top) and simulated (bottom) friction coefficients at 290ºC and 10kPa. G/PPS tool/ply //G/PPS tool/ply ⊥C/PPS tool/ply //C/PPS tool/ply ⊥C. Binetruy, F. Boussu (Eds.) - Recent Advances in Textile Composites, 2010 p.271-279 spacing between the blocks increases in the pulling direction. A similar phenomenon is known from the Michell/Kingsbury tilting-pad [ 13], which is schematically depicted in Figure 9. The tilting-pad features a normal force FN, which is applied at an offset e from the center of the pad, causing the pad to tilt. The degree of the tilt is self-aligning, since the resulting pressure distribution p(x), described by Equation (1), forms a counter moment on the pad. The relation between the tilt and the offset is known from theory: ( ) ( ) ( )( ) ( )( )( )2 2ln 3 4 ln 3 ln12 ln ln 2 2 1n n n n n nn n n n nε− + + +=+ + − − (4) where ε = e/L and n = h1/h2, as defined in Figure 9. In the case of our exemplary data, the offset would be 1.5mm. We expect that these small misalignments of the blocks, which are difficult to control, will cause a change in the pressure distribution which may once again lead to incorrect measurements. This requires further investigation.    CONCLUSION  Tool/ply and ply/ply friction in textile composite forming of PPS laminates is generally in the hydrodynamic regime. Large variations are encountered in the friction coefficients as well as the film thicknesses, depending on the local conditions (temperature, pressure and velocity). A 1½D captures most of the experimental trends with good accuracy. However, the model results are sensitive to inaccuracies in the geometric representation of the textile structure, whereas the experimental results are sensitive to slight misalignments in the set-up, especially at low film thicknesses.    variation of h [ µm]-5 0 500.511.5time [s]coeff. of friction [-]-5 0 505101520time [s]variation of h0 [ µm]-5 0 500.10.20.30.4time [s]variation of α [mrad] Figure 8. Measurement results for tool/ply G/PPS at 290°C and 10kPa, fibres predominantly transverse to the pulling direction. The velocity U changes instantaneously from 0 to 100mm/min at time t=0. Variation in time of resp. friction coefficient (left), film thickness h0 (center) and tilting angle α (right).  Three measurements are indicated with different symbols, the average by a solid line.  Figure 9. Schematic pressure distribution underneath a tilting-pad. C. Binetruy, F. Boussu (Eds.) - Recent Advances in Textile Composites, 2010 p.271-279 ACKNOWLEDGEMENTS  Paul Jannink performed most of the experiments. This work was supported by Agentschap NL under the Strategic Research Programme, Ten Cate Advanced Composites and Stork Fokker. Additional work was carried out within the Thermoplastic Composite Research Centre with  support of the Region Twente and the Gelderland & Overijssel team by means of the GO programme EFRO 2007-2013. This support is gratefully acknowledged by the authors.    REFERENCES  1. Thije, R.H.W. ten. 2007. Forming simulation of laminated composite forming processes. PhD thesis, University of Twente. URL http://doc.utwente.nl/57908/.  2. Thije, R.H.W. ten, R. Akkerman, M. Ubbink and L. van der Meer. 2010. “A lubrication approach to friction in thermoplastic composites forming processes,” Composites A (submitted). 3. Murtagh, A. M., M. R. Monaghan, and P. J. Mallon. 1993. “Investigation of the interply slip process in continuous fibre thermoplastic composites,” in Proc 9th ICCM Conf. 4. Gorczyca J. 2004. A Study of the Frictional Behaviour of a Plain-Weave Fabric During the Thermoforming Process. PhD thesis, University of Massachusetts Lowell (USA). 5. Lin, H., P. Harrison, K. van de Haar, J. Wang, A.C. Long, R. Akkerman and M.J. Clifford. 2006. “Investigation of Tool-Ply Friction of Textile Composites”, 8th International Conference on Textile Composites (TEXCOMP), 16th-18th October, Nottingham, UK, 2006. 6. Phung T., R. Paton, and A.P. Mouritz. 2003. “Characterisation of the interply shearing resistance of carbonepoxy unidirectional tape and fabric prepregs,” in Proc 6th Int ESAFORM Conf, pages 867–870. Nuova Ipsa Editore, Palermo, Italy, 2003. ISBN 88-7676-211-6. 7. Liu L., J. Chen, J.L. Gorczyca and J.A. Sherwood. 2004. “Modeling of friction and shear in thermostamping of composites - Part ii,” J. Compos. Mater., 38:1931–1947. 8. Lebrun, G., M.N. Bureau, and J. Denault. 2004. “Thermoforming-stamping of continuous glass fiber/polypropylene composites: Interlaminar and toollaminate shear properties,” J Thermoplast Compos, 17:137–165. 9. Harrison P., H. Lin, M. Ubbink, R. Akkerman, K. van de Haar and A.C. Long. 2007. “Characterising and modelling tool-ply friction of viscous textile composites,” in Proc 16th ICCM Conf, 2007. 10. Ubbink, M., “Tool ply friction of woven fabric composites”, Master’s thesis, University of Twente, 2006. 11. Akkerman, R., M.P. Ubbink, M.B. de Rooij, and R.H.W. ten Thije. 2007. ”Tool-ply friction in composite forming”, 10th International ESAFORM Conference on Materials Forming, 18th-20th April, Zaragoza, Spain,  2007. 12. Thije R. H. W. ten, and R. Akkerman. 2009. “Design of an experimental setup to measure tool-ply and ply-ply friction in thermoplastic laminates,” International Journal of Material Forming, 2:197–200, 2009. 13. Van Beek, A.. 2001. Tribologie–Levensduur en prestatie (in Dutch). Delft University of Technology-Tribology Dept.  

Friction in textile thermoplastic composites forming

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Friction in textile thermoplastic composites forming

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