Friction in textile thermoplastic composites forming

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

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.

Fn

F_f , U

pull out length [mm]

F o rc e [ N ]

pull out length [mm]

F o rc e [ N ]

Figure 1. Pull-through experiment (left) and typical force-displacement graph (right).

transverse bundle matrix longitudinal bundle

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 3 6 . h p h U x η x x   ∂ ∂ ₌ ∂   ∂ _ ∂ _ ∂ (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).

( )

( )

( )

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,

( )

_∫

( )

∫

∫

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],

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.

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 = 2 9 0 ºC 0 0.2 0.4 0.6 0.8 1 1.2 1.4 20 100 500 c o e ff . o f fr ic ti o n 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 20 100 500 c o e ff . o f fr ic ti o n to o l/ p ly 0 0.2 0.4 0.6 0.8 1 1.2 1.4 20 100 500

pul l-out ve lo city

c o e ff . o f fr ic ti o n 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 20 100 500 c o e ff . o f fr ic ti o n T = 3 1 0 ºC p ly /p ly v s t o o l/ p ly 0 0.2 0.4 0.6 0.8 1 1.2 1.4 20 100 500 pull-o ut ve locity [m m /m in ] c o e ff . o f fr ic ti o n 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 20 100 500

pull-out velocity [mm/min]

c o e ff . o f fr ic ti o n G/PPS ply/ ply ⊥ G/PPS tool/ply ⊥ C/PP S tool/ply // C/PP S ply /ply // 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 20 100 500

pull-out velocity [mm/min]

c o e ff . o f fr ic ti o n G/PPS ply/ ply ⊥ G/PPS tool/ply ⊥ C/PP S tool/ply // C/PP S ply /ply //

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 ⊥

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

0.01 0.1 1 10

1.E-06 1.E-05 1.E-04 1.E-03

He [m] c o e ff . o f fr ic ti o n [ -] 0.01 0.1 1 10

1.E-06 1.E-05 1.E-04 1.E-03

He [m] c o e ff . o f fr ic ti o n [ -] // ⊥ η U/p[m] η U/p[m] 0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2 20 100 500

pull-out velocity [mm/min]

c o e ff . o f fr ic ti o n [ -] 0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2 20 100 500

pull-out velocity [mm/min]

c o e ff . o f fr ic ti o n [ -] exp sim

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 ⊥

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:

( )

( )

( )

( )

( )

(

)

(

)

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.

v a ri a ti o n o f h [µ m ] -5 0 5 0 0.5 1 time [s] c o e ff . o f fr ic ti o n [ -] -5 0 5 0 5 10 15 time [s] v a ri a ti o n o f h0 [µ m ] -5 0 5 0 0.1 0.2 0.3 time [s] v a ri a ti o n o f α [ m ra d ]

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.

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.

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