Chapter 5 Characterisation of shearing and frictional behaviour during sheet forming

Composite Sheet FormO~g edited by D. Bhattacharyya 9 Elsevier Science B.V. All rights reserved.

Chapter 5

Characterisation of Shearing and Frictional Behaviour during Sheet Forming Adrian M. M U R T A G H Engineering Department, Cambridge University, Trumpington Street, Cambridge CB2 1PZ, UK

Patrick J. M A L L O N Mechanical and Aeronautical Engineering Department, University of Limerick, Limerick City, Ireland

Contents Abstract 163 5.1. Introduction 164 5.2. Transverse fibre flow 170 5.3. Intra-ply shear 173 5.4. Inter-ply slip 177 5.5. Friction during thermoforming 197 References 214

Abstract An introduction is given to the processes involved in the forming of thermoplastic composite and various shearing deformations are explained, including intra-ply shearing and inter-ply slip in unidirectional laminates, and the trellis mechanism in fabric materials. A consolidation rig was used to evaluated the influence of processing parameters such as pressure, temperature and time on the quality of consolidated laminates. Consolidation quality was also found to be heavily dependent on lay-up and whether or not the flow processes were restricted. Unidirectional restricted laminates were difficult to consolidate with best results achieved in low pressure regions. For (0/90) and quasi lay-ups laminate quality improved as the consolidation pressure increased with acceptable parts achieved at pressures of 500 kPa and over. Transverse flow measurements were used to obtain values of transverse flow viscosity for APC-2 material. A consolidation/shearing rig apparatus was used to carry out inter-ply slip experiments and to determine the effects of temperature, pressures and fibre orientation on the inter-layer shear stress, as a function of sliding velocity. The shearing rig is based on a computer-controlled motor-driven leadscrew. Normal pressure, increased 163

164

A.M. Murtagh and P.J. Mallon

beyond a low nominal value ( < 100 kPa), substantially increased shear stress values. Temperature effects showed the influence of the viscous matrix. Laminates in which fibres from adjoining plies lay parallel to one another showed much higher shear stress values compared with cross-ply laminates. Fibre straightening had to be taken into account in fabric specimens before slip could be initiated. The presence of additional PEEK resin layers did not decrease shear stress as expected in APC-2 laminates. Yield stress measurements were also carried out on specified materials and showed a significant level of stress had to be exceeded before true slip occurred, e.g. 1 kPa for a cross-ply laminate, 2.5 kPa for a unidirectional laminate of APC-2. Using the results obtained from the inter-ply slip tests, a model based on modified form of the Herschel-Buckley model for a power law fluid was developed for APC-2 and for 5-H satin Cetex fabric. The values for the multiplier k and exponent n were first related independently to the parameters of temperature, normal pressure and fibre angle, then a factoring technique was used to create a master model for both materials. Yield stress was incorporated in both cases: for the fabric, a modified form of the Oldroyd model was used to account for tow stretching before slip occurs. A combination solid phase/melt model for APC-2 was also generated. Friction testing was carried out on APC-2 and 5-H satin Cetex to establish the respective material's sliding behaviour as it deforms during the typical press forming process. The coefficient of friction was measured using two different test set-ups (twin platen arrangement and friction sled) according to varying parameters of sliding velocity, temperature, normal load, relative fibre angle and presence of release agent. The observed behaviour was found to be hydrodynamic in nature, with a strong adhesive element. A rubber pad/glass fabric composite interface was also investigated under non-isothermal conditions. A fiction law was developed to predict the friction coefficient between tool and part during forming (dependent on the mentioned parameters), again based on a power law fluid model, due to the establishment of a viscous resin layer at the interface. 5.1. Introduction

During the sheet forming of complex 3-D shapes from a flat laminate of fibrereinforced thermoplastic, various forming mechanisms must occur to facilitate part manufacture. For continuous fibre reinforced composites such as carbon-fibrereinforced PEEK (APC-2), there are four such mechanisms: resin percolation of the matrix amongst the fibres, transverse fibre flow, intra-ply shear and inter-ply slip of individual plies across one another. Friction between composite and forming tools may be included as a fifth mechanism. These mechanisms occur irrespective of the actual forming process or the shape of the part. However, different mechanisms may be more or less pronounced, and have a greater or lesser effect, depending on circumstances. For example the presence of flexible diaphragms in diaphragm forming allow more squeeze flow to occur at bends compared with matched-die press forming, and higher forming rates in press forming may cause high intra-ply shearing rates which may lead to buckling in the laminate. Although fundamental research has been conducted to investigate each mechanism thoroughly and in a quantitative

Shear&g and frictional behaviour dur&g sheet forming

165

fashion, the current understanding is not complete. Such an understanding is necessary to allow prediction of possible processing problems and to guarantee a successful manufacturing operation. This chapter outlines the current understanding of the forming mechanisms which are central to the successful and economic manufacture of continuous fibre-reinforced thermoplastics using a sheet forming technique. The first mechanism, resin percolation, is not in effect a shearing deformation of the reinforcing fibres but rather a movement of fluid through the fibres and will not be dealt with in detail in this chapter. The resin flow through the fibre bed for unidirectional materials is shown in fig. 5.1a. In fig. 5.1b, the flow that occurs in fabrics to fill the interstices between layers is shown. Without sufficient resin percolation, complete consolidation of the part will not occur, resulting in defects such as voids. The examination of this mechanism, which is mainly dependent on the viscosity of the matrix, has been dealt with by other researchers [1-4]. Transverse fibre flow is an important mechanism that occurs in unidirectional plies, whereby parallel fibres migrate in a sideways fashion due to a normal pressure differential being applied across the surface of the laminate. The measurement of this mechanism illustrated in fig. 5.2a is generally referred to as the transverse viscosity.

Resin flow through unldlrecflor~l fibre bed

Resin flow through fabric to fill Interstices

(a)

(b)

Fig. 5.1. Resin percolation. -3.

Resin flow

~'~ve~ fibre flow

Normal pressure

'~n'eUing' effect

(a) Fig. 5.2. Transverse fibre flow.

(b)

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A.M. Murtagh and P.J. Mallon

In fig. 5.2b, the barrelling effect that occurs for unidirectional laminates may be observed. This effect is due to the fact that beyond a limiting deformation, fibre movement tends to "lock-up" or stall. Higher pressure towards the centre causes the central region of fibres to deform further. Resin flow may also occur at the ends of the laminate, as the squeezing pressure causes the resin to migrate along the path of least resistance, i.e. in the direction of the fibres. Intra-ply shear is a mechanism that occurs within the plane of each individual ply. For unidirectional fibre-reinforced composites, the shear mechanism entails parallel movement along the length of adjoining fibres. This may occur either through the thickness in the longitudinal (1-3 plane) or along those planes parallel to the surface plane (1-2 plane). The measurement of this mechanism is shown in fig. 5.3a and 5.3b and is normally referred to as longitudinal viscosity. In unidirectional plies where the longitudinal viscosities in the 1-2 plane and the 1-3 plane are equal, the composite is said to be transversely isotropic. Also, the properties of a laminate might infer that a single ply's response is similar to the bulk unidirectional laminate response. This is not true since shear of a laminate in the 1-3 plane results in inter-ply shear between the plies which is a weaker mechanism than intra-ply shear. For fabrics, transverse flow is restricted by the weave. Depending on the thickness of the tows/type of weave, transverse flow is either minimal or practically nonexistent when compared to unidirectional materials. On a smaller scale, if each tow can be approximated to an ellipse/oblong before consolidation, then after consolidation, this ellipse becomes more elongated and its aspect ratio becomes bigger, even though the same number of fibres exists in the tow (fig. 5.4). This elongation is essentially a transverse flow effect, but on a more localised scale than unidirectional materials. This flow is important in filling in the gaps between plies and must occur for good consolidation to occur, even for simple shapes.

3

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Fig. 5.3(a). Intra-ply shear in the 1-3 and 2-3 planes. ~mmmmm m i

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Transverse

Shearing and frictional behaviour during sheet forming

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Fig. 5.4. Transverse flow in fabric tows.

For fabrics, with reinforcement in two or more directions, intra-ply shearing may take the form of the so-called in-plane "trellis effect". This mechanism is seen in all complex curvature shapes made from fabrics and if it does not occur for some reason or is inhibited in some way, then out-of-plane buckling will occur. The trellis mechanism is illustrated in fig. 5.5. When the force acting on the fabric acts at an off-set angle (shearing angle) to the direction of interlacing orthogonal fibre tows as shown, the effect is to cause the angle between the two sets of tows to decrease - - this angle is known as the "trellis" or crossover angle. This angle has a limit beyond which the tows cannot rotate any further and they "lock" in place. This "locking" angle is mainly a function of the weave style and thickness of the tow. Beyond this angle, further deformation may cause out-of-plane buckling in the part. The fourth mechanism that occurs when forming is that of inter-ply slip, whereby the individual ply layers slip across one another when forming a curved shape (fig. 5.6).

Fig. 5.5. Trellis effect in fabrics.

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A.M. Murtagh and P.J. Mallon

Resin k:~yer

Ply slipno buckling

n e d , no slip buckling of Inner plies

Fig. 5.6. Inter-ply slip between the layers.

The reason this must happen is again due to the inextensibility of the reinforcing fibres. For example, when forming a 90 ~ bend, the plies closest to the bend undergo a compressive stress as the laminate deforms, and this can result in out-of-plane buckling. The outermost plies are in tension by not being able to stretch to accommodate the increased arc length on the outer surface. In practical forming operations, each ply behaves as a separate entity with shearing taking place between the plies. Intra-ply shearing of the fibres need not be included in any analysis. The slippage effect between plies is assumed to occur in a resin-rich layer existing between the plies. In this way, the flow can be described as a simple Couette f l o w - viscous flow of a fluid between two parallel plates. This neglects the fact that the thickness of the resin layer can vary considerably from being barely evident (especially for unidirectional lay-ups) to being a few fibre diameters thick in some regions. However, a reasonable assumption is that the average thickness is approximately one fibre diameter (6 lain) which allows an inter-ply slip viscosity to be determined if the shearing velocity is known. Inter-ply rotation occurs when forming double curvature components when the angle between fibre directions in adjacent plies must change during forming to accommodate the part geometry and may be regarded as a complex inter-ply slip deformation. For fabrics, inter-ply slip between plies can also occur. However, the fibres are not so inextensible as with unidirectional materials because of the "crimped" nature of the fabric prepreg and this allows a certain straightening of the fibres when forming over a bend and inter-ply slip only occurs once the fibre tows straighten and become inextensible. Conversely, the crimped nature of the tows also means that buckling is more likely to occur in those plies that experience a compressive stress. One further phenomenon that could be considered a type of flow mechanism is the frictional interaction that must occur between the forming tools and the surface of

Shearing and frictional behaviour during sheet forming

169

the thermoplastic composite. For diaphragm forming, the composite does not make any direct contact with the mould, rather, the sliding mechanism is a combination of the interaction between the composite and diaphragm material on one side, and the diaphragm and mould material on the other side. However, during press forming of thermoplastic composite parts, there is a frictional interaction between the composite and mould surface. For matched-die moulds, it may be censidered an interaction between a fluid-like, molten polymer sliding against a hard, yet smooth, metal surface. When rubber pad forming is the process involved, the friction is between two semi-compliant mediums. The dominating factor here is temperature, further complicated by the fact that most press forming processes are anisothermal in nature and temperature has been shown to have a profound effect on the frictional characteristics of polymer systems. The current understanding of this phenomenon is incomplete and further investigation is necessary to measure the effects of temperature together with other possible variables such as normal pressure, forming speed and surface texture. Figure 5.7 gives an example of a press-formed top-hat section and shows how the frictional forces build up during the forming cycle. In Step 1, the preheated laminate is transferred from the heating station to between the punch and the mould cavity. A blank holder may be used at this point to keep the laminate in position and maintain a slight tensile force on the sheet as it is being formed. As the punch begins to

Blankholder ~ Punch Laminate

/ J

|

Fig. 5.7. Friction during press forming.

170

A.M. Murtagh and P.J. Mallon

descend (Step 2), it encounters the laminate and the frictional force begins to develop between metal and composite. As the part is more fully formed (Step 3), the frictional forces reach a maximum, dependent on the normal pressure being exerted between the punch and cavity on the composite sheet. Once the part has fully formed (Step 4), all moving components come to rest and friction has no further part to play, although other flow mechanisms such as resin percolation and transverse flow can occur. This section has served to introduce each particular mechanism as it occurs for sheet forming of continuous fibre composites. The following sections will deal with each particular mechanism in detail. 5.2. Transverse fibre flow

Most research to investigate the transverse fibre flow phenomenon has been carried out using parallel-plate "squeeze flow" apparatus. Using the assumption that the resin adheres to the platen surface, a theoretical analysis relates the transverse intra-ply shear strain to the shear stress and the time scale involved [5]. This analysis also includes the concept of a yield stress value and the transverse flow viscosity may consequently be defined by FH 2 ~7 -

2VW2 L

(5.1)

where F is the applied normal force (N) at time t (s), H is the specimen thickness (m), L is the specimen length along the fibres (m), W is the specimen width across the fibres (m) and V is the transverse flow velocity (m/s). Work performed by Barnes and Cogswell [6] indicate a value of approximately 4,000 Pa s for this transverse flow viscosity. As part of the same study, the limiting transverse flow, i.e. how much the sample spreads upon the application of a force, was measured [6] and defined using a numerical technique as a function of pressure and thickness:

w0

= 1 + 0.4p1/3H

(5.2)

where Wp is the final width, W0 is the initial width, P is the pressure (MPa) and H is the thickness (m) of the sample. A similar experimental approach using a squeeze flow apparatus was taken by Mulholland [7] and a schematic of the test rig is shown in fig. 5.8. In this apparatus, two heated platens are compressed between the jaws of a servo-hydraulic 100-kN press which subjects the lay-up to transverse flow in the lateral direction. A linear variable differential transformer (LVDT) was used to record the squeeze flow as a function of time and from this the transverse flow velocity was calculated. Two sidemounted LVDTs were used to record the thickness variation of the sample as the pressure was applied. As an example, the apparent viscosity of a (0)16 sample of APC-2 as a function of time was calculated using eq. (5.1). In this example, the applied pressure was 410 kPa. Results of this are shown in fig. 5.9, together with the recorded squeeze flow velocity. Initially, the viscosity is infinity (before flow is

Shearing and frictional behaviour dur&g sheet forming

171

Fig. 5.8. Squeeze flow measuring apparatus. 400000

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20

30

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Time (rain) Fig. 5.9. Transverse flow and apparent viscosity variation with time for (0)16 APC-2.

initiated) at point A. The initial flow viscosity around point B is approximately 60,000 Pa s, which slowly increases up to point C to a value of 300,000 Pa s as transverse flow occurs. This point at C represents the transverse flow "locking" effect when slight twisting in the fibres interferes with any further flow. When the consolidation pressure is increased at point C to 1 MPa, the higher pressure causes flow

172

A.M. Murtagh and P.J. Mallon

to re-start and the viscosity decreases once more to 60,000 Pa s at point D. After this, the sample was cooled and transverse flow ceased. The value of 60,000 Pa s observed for the initial flow viscosity is an order of magnitude higher than that recorded by Barnes and Cogswell [6]. However, lower consolidation pressures were used in their work. The effect of lay-up on consolidation quality was also investigated and fig. 5.10 shows three sample C-scans. Sample A ((90)8 lay-up) was constrained in the longitudinal direction but allowed to flow transversely. The B sample, (0,90,0,90)s lay-up, was unrestricted on all sides. Sample C ((0)8 lay-up) was arranged to prevent transverse flow parallel to the fibres, but allowed resin squeeze-out at the ends of the fibres. The C-scan for Sample A again shows the band of poor consolidation quality across the centre as seen previously. Sample B shows good consolidation quality and it was observed that there was minimal transverse flow in this specimen. Sample C shows that the resin squeezed out at the edges of the sample increases the void content substantially, indicated by the white regions along the top and bottom of the sample. Further evaluation of this work involved ultrasonic scanning of samples both with and without transverse flow having occurred. The effect of consolidation pressure can be seen in fig. 5.11 for eight (0)8 APC-2 samples. Scans (a,b,c,d) are for samples that had unrestricted transverse flow. Samples (e,f,g,h) were placed in a picture frame during consolidation and so squeeze flow was minimised. White areas on the scans indicate the presence of voids, and grey areas indicate better consolidation quality. The unrestricted samples show an increase in porosity as the consolidation pressure was increased. Most interesting is the white band that appears across the centre of each sample which grows thicker as the pressure is increased. A possible explanation for this would be the migration of resin from the centre of the laminate towards the edges where the squeeze flow was taking place, resulting in the creation of voids in the central region where the pressure was released. The other four samples (e,f,g,h)

Fig. 5.10. Ultrasonic scans for 8-ply APC-2 laminates.

Shear&g and frictional behaviour during sheet forming

173

Fig. 5.11. Ultrasonic scans for 8-ply APC-2 laminates consolidated at 0.41 MPa.

show the expected pattern, i.e. better consolidation quality as the pressure is increased.

5.3. Intra-ply shear A number of experimental and theoretical studies have been carried out to characterise axial and transverse intra-ply shear in unidirectional continuous fibrereinforced composite materials, specifically APC-2. For small deformations, experiments have been carried out using oscillatory flow techniques to quantify axial and transverse shear viscosities. Conventional rheometers operating in a torsional mode impose both axial and transverse shearing modes in a composite sample as shown in fig. 5.12. Ideal fibre reinforced fluid theory, first developed by Spencer [8], and adapted by Rogers [9], allowed both axial and transverse components to be separated out from experimental results on a number of samples. Work carried out by Groves [10] on APC-2 estimated the value for the axial intra-ply shear viscosity of 7,400 Pa s and for the transverse intra-ply shear viscosity of 6,100 Pa s, a ratio between the two viscosities of approximately 1.2. Cogswell [1] recorded slightly lower values, 6,000 Pa s in the axial and 3,500 Pa s in the transverse mode. At a higher angular velocity of 100 rad/s, Scobbo [11] measured an axial viscosity of 6,000 Pa s and a transverse viscosity of 3,500 Pa s. The average ratio of axial to transverse viscosity is approximately 1.5.

174

A.M. Murtagh and P.J. Mallon

Oscillating

" Axial shear

shear Fig. 5.12. Intra-ply shearing modes in oscillatory torsion.

All the studies performed indicate a highly non-linear response, with a yield stress of about 1,000 Pa. Comparison of the composite viscosity with the measured viscosity for the neat resin indicate a large increase in viscosity for the composite [11]. This would seem in part to be due to fibre twist and interference within the laminate. The temperature-dependence of melt viscosity for the composite material is the same as that for the resin. A 10~ increase in temperature reduces the viscosity by approximately 17% [5]. In fabrics, intra-ply shear in the 1-2 plane, which gives rise to the trellis effect as shown in fig. 5.13, can be investigated in purely kinematic terms to predict the trellis angle for a particular fabric strain. The effects of the resin are negligible on this

Instron

|

1. lnstron frame

5. Extension rod

2. Crosshead

6. Sample clamp

3. Load cell

7. Fabric sample

4. Top cooler

8. Environmental chamber

connection

i

/ \ /"" / \ /"I \ / \ /AN / \. 4

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,, \ J O

:,,.',,I \ .('

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o(t)l Ic;;•

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\ /

..-I

Ix.,, \/I - - - - - - - 1 . ~ ,,. / \ / .,, fl ["4 \ / \ /

Fig. 5.13. Trellis-angle measurement on deforming fabric specimen.

\q

I I0

Shearing and frictional behaviour during sheet forming

157

effect. For a bi-directionally reinforced fabric orientated at 45 ~ to the shearing direction, the trellis angle can be predicted from the simple kinematic expression: cos 0 -- cos 00 exp(,kt)

(5.3)

where 00 is the initial fibre angle (45 ~ in this case), at time t = 0. This shows that movement is symmetric about the X-axis and that the fibre angle decreases as a function of applied strain rate ~ and is independent of the material constitutive law, i.e. is purely kinematic. The true axial strain (not to be confused with the engineering strain) in the sheet is equal to e l l - ln(~0t)) -- )~t

(5.4)

and we can see that the angle change is simply dependent on the axial strain. Experiments were carried out to verify this model. This involved shearing an appropriate fabric specimen (glass-fibre-reinforced polyamide, 7-H satin weave, Vestopreg G101 | at processing temperature inside an environmental chamber [12]. The fabric was gripped using special jaws which were in turn connected to an Instron straining frame, as shown in fig. 5.13. A cross marked in ink at the centre of the specimen was observed as the specimen was strained vertically. The change in angle was recorded and fig. 5.14 is a plot of true strain against predicted trellis angle (from eq. (5.3)) and the observed trellis angle. The experiment shows relatively good agreement with the theory. Differences may be attributed to uneven shearing at the ends of the specimen, which was caused by the presence of the grips which 50 Model 40

D

Specimen

30 Trellis angle (e)

20

10

'Locking'

,=,,,=

angle 0.0

0.1

0.2 True strain

0.3 I; ! 1

Fig. 5.14. Comparison of model and test values for trellis angle.

0.4

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A.M. Murtagh and P.J. Mallon

constrained the sample from reducing in width. The shaded region indicates the spread of values for which the trellis locking angle was observed, i.e. the strain at which out-of-plane buckling was first observed visually on the specimen. Thus for G 101 fabric, the locking angle lies between 25 ~ and 30 ~ This value is in accordance with that observed by other researchers [13,14]. The ideal fibre-reinforced fluid constitutive model for one family of fibres has been extended to include two families of fibres by Johnson [15]. The original model [8] defines stress in terms of strain rate and two viscosity terms/71 and/Tt, the axial and transverse viscosities. For a fabric, a third viscosity term must be included: 1

O'11 -- ~, [4/71 -- (3/71 -- 2/72) sin 2 20 -- ~/73 sin2 40]/sin 4 0

(5.5)

all is the stress in the 1-1 direction, Z is the shear rate and 0 is the trellis angle. The three viscosity terms are denoted by/71,/72 and/73. They may not be related to simple shearing modes in any particular plane in the fabric, unlike/Tz and/Tt which can be related to shear in the 1-3 and 2-3 planes for unidirectional materials. Instead, they may be consider as "mixed-mode" parameters. Using the same experimental set-up that was used to verify the kinematic model for trellis angle prediction, it was possible to obtain an estimate for these three viscosities for the G101 specimen. Three samples were sheared at a constant shear rate at three different temperatures. Tensile stress in the 1-1 direction was calculated as a function of tensile force and the instantaneous cross-sectional area of the sample across the centre. A plot of true strain against tensile stress is shown in fig. 5.15, for 225~ 245~ and 265~ 5.00e+6

-

,,

4.00e+6 al

n_

,,,

,,,

a

Exp 225~

e

Exp 245~

,&

Exp 265~

I

9" ,e,-

. .1

Model225~

2

Model 245~

3.00e+6

t:)

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3

Model 265~

2.00e+6

.e m

m r o I-

1.00e+6

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0.00

-

0.05

0.10 True strain

0.15 ~

Fig. 5.15. G101 fabric stretch test modelled using eq. (5.5).

11

0.20

0.25

Shearing and frictional behaviour during sheet forming

177

Equation (5.5) was used to model the stress response as shown. Note that the initial elastic-viscous stretching of the specimen (around region A) is not modelled as the model assumes steady flow. The value of the three viscosity parameters r/l, 02 and 03 are given in table 5.1. Values for r/1 decrease as temperature increases, while conversely values for r/2 increase with temperature. This would not be expected for "real" viscosity components but may be possible for these "mixed-mode" parameters. The ratio between 01 and 02 decreases from 13.33 to 1.71 as temperature is raised from 225~ to 265~ At lower temperatures, an adequate curve fit was achieved by setting ~2 = 0, which simplifies the model considerably. The optimum value for 773 was found to be approximately zero for all conditions.

5.4. Inter-ply slip Inter-ply slip in unidirectional materials occurs due to the fact that the reinforcing fibres in a composite laminate may be considered to be inextensible. When forming single or double curvature shapes, where the stack of plies making up the laminate may be considered as a number of inextensible, yet flexible, plates, then a relative displacement must occur between the layers to accommodate the different path lengths of each ply around the bend. Micrographs taken at the edge of formed parts with initially ground and square edges serve to illustrate this phenomenon [12]. The total slip deformation, d, can be shown to be a function of the bend angle, 0, number of plies, N, and the thickness of each ply, t (see fig. 5.16), and is given by Ot

d - ( N - 1)~-

(5.6)

T A B L E 5.1 P a r a m e t e r s for fabric m o d e l T e m p e r a t u r e (~

O~ (Pa s)

7"]2 (Pa s)

773 (Pa s)

265 ~ 245 ~ 225 ~

1.2 e 7 3.0 e 7 4.0 e 7

7.0 e 6 3.5 e 6 3.0 e 6

0 0 0

t

0

Fig. 5.16. Ply slip a r o u n d b e n d angle 0.

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A.M. Murtagh and P.J. Mallon

Figure 5.17 is taken from a (0,90,0,90,)s APC-2 laminate formed over a 90~ The slip occurs in the resin layer between each ply and intra-ply shear within each ply is negligible. The step-like slip effect can be seen along the edge of the laminate (for example at point A, the interface of a 0 ~ and 90 ~ ply). The only point at which no slip can be seen is at the central axis of the laminate, at the interface between two 90 ~ plies (point B). Other cross-ply laminates show similar degrees of slip fig. 5.18 shows the deformed edges of a quasi-isotropic (0/+ 45/90/-45)s 90~ part. Resin layers can be seen as dark lines between the plies. Some intra-ply shearing is evident on the top 0 ~ and adjoining +45 ~ ply, possibly due to friction against the tool during forming. The slip behaviour of unidirectional (i.e. all plies oriented at 0 ~ laminates contrasts greatly with that for cross-ply parts. In most cases, for pre-consolidated unidirectional parts, the dominant flow behaviour was intra-laminar, with no obvious resin layers existing and with the entire thickness of the laminate behaving in many respects as a single ply. This is shown in fig. 5.19 for a (0)8 90~ part where the

Fig. 5.17. Inter-ply slip for (0,90,0,90)slaminate.

Fig. 5.18. Inter-ply slip for (0/+45/90/-45)s laminate.

Shearing and frictional behaviour during sheet forming

179

Fig. 5.19. (0)8 l a m i n a t e - intra-ply shearing.

flow is even across the thickness, and there is no distinction between plies. One exception to this rule was observed with laminates pre-consolidated at low pressure (100 kPa) before forming. Figure 5.20 shows a distinct step-like behaviour (no apparent resin layers) with each ply seeming to act independently. This may be caused by the low consolidation pressure not causing sufficient fibre/fibre interaction between plies to, in effect, "fuse" the plies together. In fabrics, the inter-ply slip behaviour can differ markedly from that seen in unidirectional materials. The existence of a crimp in the fabric due to the woven nature of the tows means that the assumption of inextensibility does not hold. When a tensile force is placed across the plies during forming, the crimp allows fibres to straighten to a certain degree. The amount of possible stretch varies depending on the weave style [17]. This is shown in fig. 5.21 - - the fibre straightening factor (FSF) is the possible percentage change in length of a ply segment along its length. It is a maximum for the most highly crimped style, i.e. plain weave. To measure this elongation in a ply under tension, specifically for Cetex fabric, rectangular samples were tested in an environmental chamber and their load response to an applied displacement was measured [16]. Figure 5.22 is a plot of

Fig. 5.20. (0)8 laminate - - step-like behaviour.

180

A.M. Murtagh and P.J. Mallon

Fig. 5.21. Possible tow straightening in fabrics.

Cetex fabrics:

-

B

Temperature 320~ Speed 0.5 mm/min

0. =E

I a 9 .

5-H weave Plain weave .Model 5,.H

m. r

I-

m

0

. - ~ -

|

1

....

2 Strain

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3

(%)

Fig. 5.22. Stretch testing of Cetex fabric specimens.

strain in the longitudinal direction against tensile stress, showing the response for both a plain weave and a 5-H satin weave sample. The behaviour shows that there is an initial strain in the samples before the stress begins to increase exponentially, i.e. a "strain-hardening" phenomenon. The plain weave allows for more tow stretching compared to the 5-H satin weave sample. The stress/strain behaviour of the 5-H satin weave sample was modelled using a curve-fitting technique as follows: r~ - 0.03061 e 2"1003(e)

(5.7)

where a = tensile stress and e = strain in specimen. This model is useful in predicting internal tensile stresses in a deforming fabric ply and can be used in conjunction with an understanding of the inter-ply slip behaviour of the material to form part a numerical model for sheet forming of Cetex fabric. As stated previously, inter-ply slip occurs in the thin resin interlayer between plies. In order to obtain a thorough and quantitative understanding of the nature of the shear stresses to be encountered in this layer during forming, a number of analyses

Shear&g and frictional behaviour during sheet form&g

181

by different researchers have been carried out. Cogswell [1] was the first to identify the mechanism for unidirectional tapes, specifically APC-2, and carried out some initial slip characterisation studies. Scherer [18] obtained experimental results from an inter-ply slip apparatus taking into account the effects of shearing velocity, temperature, pressure and lay-up for a carbon-fibre-reinforced polypropylene unidirectional composite. A model obtained using these results was used as input to a finite element analysis on the inter-ply slip process during thermoforming [19]. Xiang Wu [20] also performed pull-out tests on 16-ply uniaxial APC-2 laminates and modelled the forming of 90~ parts as approximating 3-point bending. He calculated the apparent viscosity in the resin-rich interlayer and noted an order of magnitude difference between measured values for a (0)16 laminate and predicted values for the neat resin. Groves [21] observed this viscosity increase in APC-2 using a dynamic spectrometer and one explanation proposed was possible fibre friction and interference between layers. Further investigations by Jones [22] showed that the observed non-Newtonian behaviour of the composite and the assumption of the individual plies behaving as buoyant plates was not due to inertia effects and instead was due to intra-ply shearing. Kaprielian and O'Neill [23] performed pull-out experiments on a steel plate from between stacked plies of APC-2 and were able to correlate results with a model based on the ideal fibre-reinforced fluid constitutive equation [8]. Soll [24] performed tests to optimise the forming conditions when forming simple rightangled bends using a unidirectional material and noted that a tensile force applied across the laminate during forming increased part quality by reducing fibre wrinkling. Tam and Gutowski [25] developed a linear visco-elastic model of the inter-ply slip process, modelling the fibre-rich plies as a linear elastic layer and the resin interlayer as a viscous layer. This model allows prediction of laminate stresses and displacements, and strains in the slip region. Results showed good correlation with isothermal forming experiments. Scobbo [11], using DMA analysis, observed the effects of fibre interaction between adjoining plies with fibre effects beginning to dominate visco-elastic behaviour at higher temperatures. More recently, Morris and Sun [26] have performed inter-ply slip experiments on APC-2 laminates at lower temperatures, approaching and below the melt temperature, Tm, and have modelled this solid-phase forming behaviour. The remaining part of this section will describe and summarise part of another inter-ply slip study performed by the authors [12,16,27]. This work was carried out with the aim of simulating the conditions of slip during forming in a controlled manner and examining the effects of different forming parameters (slip velocity, temperature, normal pressure and fibre orientation) on shear stresses in the interlayer. Experiments were carried out using a consolidation unit [7] together with a custom-built shearing apparatus [28]. The consolidation unit consists of two heatable platens mounted on a 100 kN hydraulic press. The shearing apparatus is made up of a motor driven leadscrew which performs the pull-out motion on the horizontally mounted sample lay-up. A 500 N load cell recorded the shearing load. A schematic of the apparatus is shown in fig. 5.23. Figure 5.24 is an overall view of the shearing apparatus and the consolidation mounted together on the hydraulic press.

182

A.M. Murtagh and P.J. Mallon

Consolidation unit

Exterior ply clamp

DC M otor

Load cell

Shearing apparatus

Fig. 5.23. Consolidation unit/shearing apparatus for inter-ply slip testing.

Fig. 5.24 Test set-up for inter-ply slip experiments.

The sample lay-up comprised of a central pull-out ply, two exterior stationary plies and intermediate plies with various fibre orientations. For the APC-2 sample shown in fig. 5.25, the lay-up is (0, 90, 0)s the central and two exterior plies are oriented at 0 ~ to the pullout direction, and the two inner plies are at 90 ~ A narrow shim of material behind the pull-out ply keeps the gap between top and bottom platen constant to avoid "pinching" of the specimen. All APC-2 and Cetex fabric specimens were first consolidated at a pressure of 1 MPa for 5 minutes prior to testing. A specified normal load (usually 100 kPa) was then applied via the hydraulic press for the duration of the pull-out test. The shearing

Shearing and frictional behaviour during sheet forming

183

Shim

Fig. 5.25. Specimen geometry for inter-ply slip testing.

force and ply displacement were recorded for a particular shearing velocity. The shear stress was calculated simply by dividing the shear force by the sheared area using Fs

r -

(5.8)

w ( L - d)

where r is the shear stress, Fs is the shearing force, L and w are given by the specimen geometry and d is the displacement of the pull-out ply at a particular time t. Figure 5.26 show the results for a particular test, where the shearing velocity has been increased in increments. Each increase in velocity causes a corresponding increase in the associated "steady" shear load, as shown. This allows the shear

0.4

250 u

(0,90, 0)S APC-2 CP 1 MPa, NP 0.1 MPa

200

SPT 375~

0.3 A

A

Z v

150

E

"o

0.2

...o

vE

3

l_

8

2 .-._,,ml

m

L.

0.1 50

Load

]

Velocity 0

100

200 Time (sec.)

Fig. 5.26. Velocity increments during inter-ply slip test.

300

400

0.0

=

,,r

184

A.M. Murtagh and P.J. Mallon

velocity/shear stress relationship to be determined from an average of single sample tests for a particular set of conditions. This technique was used throughout this test programme to obtain all the shear velocity/shear stress plots shown. The effect of processing temperature has a large effect on the shear deformation behaviour of composites. This is due to the resin viscosity, which is strongly dependent on temperature. The reinforcing fibres are unaffected. For APC-2, the melt temperature is 343~ so theoretically a laminate can be formed at any temperature above this. It is recommended that processing should occur in a window between 360~ and 400~ Isothermal conditions, and a constant resin viscosity is readily achievable with diaphragm forming, where the entire laminate is held in a heated chamber. For press forming, temperature may be uniform initially, after being transferred from the heating unit, but decreases rapidly, due to convection cooling in the air. Conduction cooling occurs once contact is made with the tool, which may be around 250~ Indeed, forming may actually occur at temperatures less than 343~ for APC-2 under certain conditions, but in this case, inter-ply shearing may result in a high level of shear stress [26]. Figure 5.27 shows the effect of varying temperature on a (0, 90, 0)s sample. As expected, raising the temperature decreases the shear stresses in the laminate. However, at higher temperatures, problems such as surface oxidation and excessive in-plane fibre wrinkling or "washing" (due to the low viscosity of the resin) may occur. For an anisothermal process such as press forming, a wide variation in temperature is unavoidable. The effect of releasing normal pressure completely during a typical slip test after consolidation is shown in the initial part of fig. 5.28. After reaching a "yield" level of 12

a

,

,,

,,,,

,,=

(0,90,0)S APC-2 CP 1 MPa NP 100 kPa 10

A

J tJ

360~ 3700C

L

l

A v

Ir

380~ 390~ 400~

A

ml|

0.0

|

i

!

0.1 Shear velocity (mmls)

Fig. 5.27. Effect of temperature on inter-ply slip of APC-2.

I

0.2

Shear&g and frictional behaviour during sheet forming

185

2.5

200 a

(0,90,90,0)S APC-2 SPT 385~ CP 1.5 MPa 2.0 150 A

1.5

A

zv

Shear load

_8 loo | _m

Pressure

x

i_

1.0 m

m

50 0.5

0

o Z

0.0 0

200

400

600

800

Time (sec.)

Fig. 5.28. Variation of shearing force with normal pressure.

100 N, shear load relaxes to a reduced level of 25 N. However, once pressure begins to increase, to about 100 kPa, shearing forces recover to a higher level than the original maximum (120 N). At a constant pull-out velocity, as pressure is increased further, the shearing force continues to increase, but the magnitude of further increases is much smaller than the initial "jump" when pressure was first reapplied. The initial jump increased stress levels by 500% - - at a pressure of 2 MPa (20 times the original step increase), shear stress has only increased by a further 50%. One possible reason for this is that under no normal force, the plies are free to disengage from intimate contact and can slide semi-freely against one another. However, as soon as a nominal normal pressure is re-applied, the plies come together once again, intimate contact is re-established and viscous slip in the layer between plies resumes. Further increase in pressure does not greatly affect this viscous slip. Theoretically, for an incompressible fluid, viscous forces should be independent of normal pressure, in the absence of a pressure gradient along the direction of flow. In practice, pressure increases do increase the shear stresses, possibly due to increased fibre contact, but not significantly. This effect need not be crucial in terms of diaphragm forming, where there is always a normal pressure present in the form of a vacuum applied between the diaphragms. As the pressure rises and the part begins to form, a hydrostatic pressure across the area of the laminate would cause an increase in shear stress. However, most forming would occur at low pressures, as there would be no reaction force against the back side of the bottom diaphragm until it made contact with the tool, when inter-ply slip would have ceased. For press forming, the situation is more

186

A.M. Murtagh and P.J. Mallon

complex. Initially, the laminate is heated to forming temperature and is under no normal pressure, except for the parts of the surface directly beneath those parts of the forming die which first come into contact with the top ply. At a constant pull-out velocity, as pressure is increased further, the shearing force continues to increase, but the magnitude of further increases is much smaller than the initial "jump" when pressure was first reapplied. Thus, some regions may undergo inter-ply slip, under no normal pressure, until the forming die comes into direct contact. The pressure distribution across the laminate is much more difficult to predict than with diaphragm forming. Further tests were carried out to establish the shearing velocity/shear stress relationship as normal pressure was varied from 20 to 400 kPa. This involved preconsolidating (0, 90, 0)s APC-2 specimens at 1 MPa. Pull-out tests were then carried out on each specimen under a particular normal pressure, and the shear stress level was recorded as shearing velocity increased. Temperature during this programme of experiments was kept constant at 385~ Results are shown in fig. 5.29. As expected, shear stress rises as normal pressure is increased. Figure 5.30 shows a plot of normal pressure plotted against shear stress at two typical shearing velocities, 0.075 and 0.22 mm/s. This plot shows further evidence that significant increases in shear stress occur at lower values of pressure, especially so for the higher velocity. Once pressure increases beyond 100 kPa, the increase in shear stress is less pronounced as the pressure rises further. To examine the effect of varying fibre orientation, a series of experiments was carried out under standard conditions of pressure and temperature, for different lay30ee

(0, 90, 0)S APC-2 SPT 385~ CP 1 MPa

Normal

pressure (kPa) " " 20 eL ,=~

:':9- - - . , . e - - - -

v

w

f~

I

20 40

-'-

100

i

200

t-.

tll @

,=

----o---10

0

I

0.0

i

0.1

,I

I

0.2

9

|

I

0.3

0.4

Shearing velocity (mmls) Fig. 5.29. Effect of normal pressure on inter-ply slip of APC-2.

9

0.5

400

187

Shearing and frictional behaviour during sheet forming

20

A

cO

A.

Shear stress at velocity"

v

T= lo

.-e-

(4

-e=- 0.22mm/s

t,=

0.075mm/s

@

0

100

200

300

400

Normal pressure (kPa)

Fig. 5.30. Significant increase in shear stress at low normal pressure. up orientations. As it was not possible to change the orientation of the pull-out and exterior plies to any angle other than 0 ~ relative to the pull-out direction, the fibre angle of the free ply between the exterior and pull-out ply was varied. Thus the layup for a particular experiment may be expressed as (0, 0, 0)s, where 0 is any angle between 0 ~ and 90 ~ Figure 5.31 shows a plot of shear velocity versus shear stress for a number of lay-ups. Lay-ups where the fibres in the free ply lay at some angle other than 0 ~ to the pull-out direction gave similar results, showing that the resin layer in each case was fully existent and of similar proportions for all orientations. Some 30 (0, 0,0)S APC-2

01010

SPT 385~ =

013010

n_ 20

-=

0/60/0

m

=

0/90/0

CP 1 MPa NP 0.1 MPa

014510 A

(II

r, L_

Q

,c

10 "

,

0.0

1

i

i

illll

i

|

iii

0.1

i

i|1

i

illl

0.2

Shear velocity (mmls)

Fig. 5.31. Effect of lay-up variation on inter-ply slip of APC-2.

i

i

0.3

i

188

A.M. Murtagh and P.J. Mallon

fibre rotation was observed, especially in the 30 ~ and 45 ~ lay-ups. Fibres in the free ply tended to re-orientate themselves to become aligned with the fibres in the pull-out ply and the more grossly displaced specimens showed most evidence of this. In the (0, 0, 0)s lay-ups, shear stresses were much higher than that for angled layups: in this case, it might well be assumed that a distinct resin layer was not formed during consolidation, and shearing actually occurred through the thickness of the middle ply. Micrographs taken through a section of a (0)8 laminate indicate the absence of any distinct resin layer [12]. Further evidence of this inhibited form of shearing in (0, 0, 0)s lay-ups is shown in fig. 5.32, which illustrates the instability in shear stress as shearing occurred at a steady velocity. This is probably due to fibre interference and entanglement between plies. Observations on tested specimens showed that this was indeed the case, with some fibres being grossly distorted and buckled. As already mentioned, the effects of fibre interaction have also been observed by other researchers [11]. If we now consider a further analysis of the different slippage behaviour between the parallel-plied (0, 0, 0)s and other cross-plied lay-ups ((0, 90, 0)s (0, 45, 0)s, etc.), we can relate the different shearing rates that occur to determine an inter-ply slip viscosity. If we assume the cross-plied lay-ups to behave in the same fashion, i.e. a 6 pm resin layer being sheared between the plies, the associated shearing rate in this layer can be calculated and from this the viscosity can be found. For the parallelplied lay-up, shearing occurs throughout the thickness of the ply between the exterior 500

-

0.5 o

(0,0,0)S APC-2 SPT 385"C C P 1 MPa NP 0.1 M P a

z 9-" Io t~ o

400

0.4

300

0.3 Unstable

1

e= G

200

0.2

# IlL

l

100

,=.,.

i....

r @ ,c

I

0.1

--D- Shear load --

Velocity

0

0.0 0

200

400

o @ > ==

m

600

800

T i m e (sec.)

Fig. 5.32. Stress instability during (0, 0, I))s pullout.

1000

1200

r

Shear&g and frictional behaviour dur&g sheet form&g

189

and pull-out ply so the sheared layer thickness is equal to the ply thickness (125 ~tm). The viscosity is this case must be much higher due to the lower shear rates and this is shown in fig. 5.33. The viscosity of the parallel-plied lay-up ( > 20,000 Pa s) is almost two orders of magnitude greater than the viscosity of the cross-plied lay-up ( < 1,000 Pa s). The viscosity seen in the cross-plied lay-up is similar to the viscosity of neat PEEK resin as measured by Cogswell [5]. The slip viscosity for the (0, 0, 0)s lay-up may be related to the longitudinal intra-ply shearing mechanism, as mentioned in the previous section. Inter-ply slip of fabrics is affected by the tow straightening effect, as mentioned previously. Pull-out experiments on fabric samples were further complicated by having two directions of reinforcement tensile force could be applied to fibre tows lying in the pull-out direction, but the transverse tows at 90 ~ to the direction of pull-out tended to remain behind in the sample whenever deformations over a few millimetres occurred. This effect was not so pronounced in unconsolidated specimens and shear results shown here are for samples tested at lower pressures than would be expected during full consolidation. Figure 5.34 illustrates the difference observed in shear behaviour between unidirectional and fabric materials during a typical pull-out test. Both results shown are for a carbon fibre/PEI material (the fabric is 5-H satin weave), tested under similar conditions of heat and pressure. The 0~ pull-out ply in the (0, 90, 0)s allows direct transmission of the traction load to the sample and the measured shear load quickly rises to a steady level. For the fabric material, the shear load

100000

-

10000 A I0

/

F" : -

Parallel-plied layup

Cross-plied layup

v

looo

s

0 O

lOO

10 1000

i

a

i

J

t

a

i

iJ ........

I

I

10000 Shear stress (Pa)

Fig. 5.33. Inter-ply shearing viscosity for parallel-plied, cross-pSed lay-ups.

I

I

I

I I

190

A.M. Murtagh and P.J. Mallon

120

i-

t 7 b~176176 F

lOO

......... ,,

, ,,

,

Material:

CF/PEI

,

SPT 320 ~ NP 100 kPa

8O Z

Unidirectional" Transistion 55N

ID

o

A

60 I

,

_ L LI __.111 .

i. , ..... i

slip

r a=

r

(0,90, 0)~ Fabric"

40 Fibre straightening

20

([o/9o1, [o.9oi ),

0 0

50

100

150

Time (sec.)

Fig. 5.34. Direct comparison of unidirectional/fabric materials.

increases in a different f a s h i o n - initially, the load applied to the ply causes the fibre tows to straighten and this continues up to point A. Here, the nature of the load increase changes from an increasing to a decreasing rate. This point, at 55 N, can be regarded as the required load to cause inter-laminar shearing along the full length of the sample at the interface between two plies, rather than cause the tows to straighten any further. The stationary level of shear stress reached thereafter is composed of two parts the tensile force in the stretched fibres plus the viscous traction force required to shear the plies. The transition load is a function of the total amount of fibre stretching, in both the area of the sample under normal pressure and the pull-out ply length between the sample and ply clamp on the shearing apparatus. Any analysis of inter-ply slip in fabrics must take account of initial fibre stretching. Shear velocity/shear stress plots may be generated in a similar fashion to unidirectional materials by determining the steady shear load levels corresponding to various shear velocities. The effect of processing temperature on inter-ply slip of fabrics may be regarded as similar to that for unidirectional materials once fibre stretching has been taken into account. Figure 5.35 shows a plot of shear velocity versus shear stress for a Cetex 5-H satin fabric sample, sheared at temperatures between 300~ and 340~ Normal pressure remained constant at 100 kPa for each test. As with APC-2, an increase in temperature of 40~ reduces the level of shear stress occurring in inter-ply slip substantially, due a decrease in viscosity of the PEI matrix material. Localised inplane wrinkling of surface fibres that can occur at increased processing temperatures

Shearing and frictional behaviour during sheet forming

191

50

NP 100 kPa 3 plies 5-H Cetex 40

A t~

30

W

t= i_ tll Q

20 t

tn

300"C

10

,

_=

,

r. 0

I

0.0

|

!

0.1

I

320oc 3400C

,

0.2

I

0.3

Shearing velocity (ram/s)

Fig. 5.35. Effect of processing temperature on Cetex fabric slip. with unidirectional fibre-reinforced composites is not such a large problem with fabrics as the woven nature of the tows constrains any excessive movement of fibres. For variation of normal pressure, tests were performed between applied pressures of 20 kPa and 400 kPa. Figure 5.36 shows the obtained results; load cell limitations reduced the attainable shearing velocity to approximately 0.25 mm/s at higher normal pressures. As with APC-2, a significant increase in shear stress is seen as the pressure is increased. The magnitude of the recorded stresses are much higher compared with the unidirectional material. Even allowing for matrix and viscosity differences, there is a much greater resistance to inter-ply slip for the Cetex material, again probably due to the nature of the interaction between adjoining plies, with much more interference to sliding being caused by the uneven surface of the woven plies. Initial test results showed that inter-ply slip of laminates was not initiated until a certain yield point had been reached. Even at low shearing velocities, pull-out forces required to shear the laminate were substantial. In order to determine this value more accurately, the test set-up was changed. Rather than use the leadscrew to provide the pull-out force, a pulley system using dead-weights was installed at the front of the shearing frame. To measure the very small displacements at loads approaching the yield point, two miniature LVDTs were mounted behind the ply clamp (see fig. 5.37). This allows very small movements of the central ply to be recorded. Care was taken to exclude any movement of the ply clamp and any displacement necessary to take up slack in the pull-out ply. The yield load was defined

192

A.M. Murtagh and P.J. Mallon

50

40

Normal

pressure (kPa) A

lg a,.

30

20

w m @

"-='-

20

--e--

50

=

100

--o-

200 400

10 SPT 320~

3 plies 5-H Cetex 0

9 0.0

~ ,, 0.1

,

I 0.2

i

I 0.3

=

I

0.4

"

0.5

Shearing velocity (mm/s)

Fig. 5.36. Effect of normal pressure on Cetex fabric slip.

.

0 LVDT A

3

...

Pulley

Load cell

o I

LVDT B Dead

weights

Fig. 5.37. Positioning of LVDTs for yield measurements.

as the force required to cause an irrecoverable displacement of the pull-out ply. For unidirectional materials, this point was relatively easy to determine as pull-out force is transmitted directly through the straight fibres to the specimen, For fabrics, however, initial fibre straightening meant it was difficult to separate fabric stretching from true slippage displacement in the lay-up. Inter-ply shear tests were carried out under various conditions of temperature, pressure and lay-up to see their effect on the yield point. Figure 5.38 shows a typical yield plot of applied dead-weight loading and recorded displacement against time for a steel foil sheet being sheared from between two plies of APC-2 (i.e. friction of

Shear&g and frictional behaviour dur&g sheet forming

i

'F

40

8

30

6

193

-

20

(g

2

t~ 100 0

0 500

1000

1500

2000

Time(sec.) Fig. 5.38. Yielding situation for APC-2/steel foil.

composite against a smooth surface see section 5.5). This result shows the ideal s i t u a t i o n - no displacement is seen until the critical load is applied, when a definite yielding of the ply is observed, and seen to increase steadily as the load is kept constant. This sudden yield occurs only when there is no interaction between fibres from adjoining layers. Further increases in applied load causes the displacement to occur at a faster rate. In reality, most samples with two plies interacting did not show this behaviour. Instead, initial yielding showed a slight displacement (as in the initial part of fig. 5.39), followed by an almost complete cessation Of movement. This behaviour represents an elastic/plastic effect and cannot be regarded as a true yield in terms of continuous inter-ply slip. Instead, the yield point can be defined as the point at which irrecoverable, steady shear flow commences. A typical response in yielding is shown in fig. 5.39, for a (0, 0)s APC-2 sample. Here, as would be expected, application of an increasing dead-weight force causes movement of the pull-out ply, but without continuous sliding occurring. At point A, when 45 N of load has been applied, displacement of the pull-out ply has reached approximately 0.2 mm. Region B shows the response when the load is taken off; the ply displacement recovers back to practically a zero value, indicating an elastic effect. This effect may be due in some measure to the fact that as loading occurs, fibres from adjacent plies interact and cause an elastic resistance to shearing, and "spring-back" when the load is removed. Although this behaviour was observed in all lay-ups of APC-2, it was most evident in the (0, 0)s sample. This behaviour in the fibre direction may be related to a similar elastic "spring-back" effect observed in the through-

194

A.M. Murtagh and P.J. Mallon

100 ' " l Note: Noise in the load cell signal [ causes each dead-weight load to

-I 2.0

I

,,--.~

80 I flickerto some degree

n

60 ~- (0,O)sAPC-2 [ SPT 385"C '~k~ 40 P I a CPr 1MPaNP100

I

I

0

A /-d

Load

ne

/

ii

]

1.5

/!

/I

fl 1"0 e=eue 0.5

i~

20

o0

250

500

750

1000

io.o

1250

Time (see.) Fig. 5.39. Yielding of a (0, 0)s sample.

thickness direction, in other research by Muzzi [29], where the fibre bundles are assumed to behave like coiled springs. Applying a compressive force to a bundle of stiff, slightly wavy fibres causes them to compress elastically, and when the compressive load is removed they can recover. In shear, adjoining fibre layers m a y impress u p o n one another under pressure and slippage of one layer across the other causes a slight elastic effect where some fibres m a y " s n a g " temporarily. Once shearing is halted, elastic recovery of any " s n a g g e d " fibres m a y occur. Increasing the load to the level shown at C caused complete yielding of the specimen. Table 5.2 summarises the results for various yield stress measurements carried out on unidirectional materials under different conditions. F o r APC-2, the yield stress TABLE 5.2 Yield stress values for APC-2 Material

Lay-up

Conditions

APC-2

(0, 90, 0)s

365-405~ NP 50,100 kPa NP 400 kPa 365-405~ NP 50,100 kPa NP 400 kPa

(0, 0)s

Yield stress (kPa) 1.1• 1.2-t-0.2 2.4• 2.6+0.2

Shearing and frictional behaviour during sheet forming

195

more than doubles (1.2 to 2.6 kPa) when the fibres are in an aligned state, i.e. (0, 0) s, compared with a (0,90) arrangement, or for any other lay-up where adjacent plies lie at an angle 0 to one another. To model the inter-ply slip behaviour of APC-2 and Cetex fabric, a modified form of the Herschel-Buckley power model was used as shown (v = shear velocity): (5.9)

Z" = "gyield + k(v) n

This allows a value for a yield s t r e s s (Z'yield) t o be inputted and then a power-law relation can be used to describe the viscous flow in the resin interlayer between plies. The values of the parameters Z'yield, k and n are dependent on the various process conditions (temperature, normal pressure, fibre orientation). Using the experimental data obtained, a curve-fitting technique was used to determine relationships between the process conditions and between ~'yield, k and n. For example, the effect of normal pressure on shear stress for APC-2 can be described by r(P, V) = (0.95 + 1.28 e-3(P)) + (-28.639 +

31.143(logP))(V~176176176 (5.10)

i.e.

"~yield - - 0.95 + 1.28 e-3(P) kPa

k = (-28.639 + 31.143(log P) n = 0.1635 + 0.3079(log P) valid for: velocity 0 < V < 0.5 mm/s, and pressure 20 kPa < P < 400 kPa. The curve-fits for the effect of normal pressure are shown in fig. 5.40. In order to combine these three different models into one master equation to calculate shear stress for any arbitrary combination of temperature, T, relative fibre angle, 0, and normal pressure P, in terms of shearing velocity, V, a factoring technique was applied. Using this method, a standard set of values was used to normalise all other points, which results in three factors fr, fp and fo, which, when multiplied by the original standard value, give the shear stress for any set of parameters. The standard conditions chosen for APC-2 were as follows: 9 Temperature 385~ 9 Normal pressure 100 kPa 9 Fibre orientation 90 ~ This results in the standard model: rs = 1.0 + 28.7083 V 0"8152

(5.11)

Then we can write

Apc-2(v, T, P, O) =iT "Up- Jb-

(5.12)

or "~APC_2(V, T, P, 0) =

r(v, T) r(V, e) r(v, 0) rs

rs

rs

(5.13)

A.M. Murtagh and P.J. Mallon

196 35

3O

20 kPa model ---2

40 kPa model

25

-"-= 100 kPa model

tl Q" 20

"-'-8 400 kPa model

__.4 200 kPa model

A

L..

[]

20 kPa exp

0

40 kPa exp

&

100 kPa exp

X

200kPaexp

4-

400 kPa exp

o r

10

0 0.0

0.1

0.2

0.3

0.4

0.5

Shear velocity (mints) Fig. 5.40. Model versus experimental values - - pressure variation.

A similar method was applied to modelling the behaviour of Cetex fabric, the exception being that the possibility of tow stretching has to be taken into account. From fig. 5.34, it is known that the transition between tow stretching and inter-ply slip occurs at a shear stress of 5.5 kPa. By equating shear stress in the interlayer to tensile stress in the ply at this point, we can write: ~'A s - - f i a t

(5.14)

where As is the sheared area and A t is the cross-sectional area of the ply under tension. Assuming constant width, we can than deduce that the tensile stress in a ply, that causes a shear stress of 5.5 kPa, can be represented by a-~

rLf

= 1.667 e7Lf

(5.15)

where r = 5.5 kPa, and tp = ply thickness = 0.33 mm Thus the effective "span" length, Lf, of the structure being formed, i.e. the length of the ply undergoing shear, is critical in determining the tensile stress. If the flange length Lr is sufficient that the total required amount of inter-ply slip (depending on geometry conditions) can be accommodated as pure fabric stretching, then no slippage between the plies need occur. Thus in a numerical simulation of press forming, a check should first be made to calculate the required amount of inter-ply slip from eq. (5.6). Using this value, and by calculating the required amount of longitudinal strain, the associated tensile stress in the deforming ply can be found from eq. (5.7).

Shearingandfrictional behaviourduringsheetforming

197

If the resulting tensile stress is sufficient to cause a shear stress of greater than 5.5 kPa in the interlayer, then inter-ply slip must occur and the power-law part of the model must be introduced. Similarly to APC-2, the effect of temperature and normal pressure on shear stress can be modelled using a curve-fitting/normalisation technique. The effect of fibre angle was ignored for the fabric. The standard set of conditions chosen for Cetex are as follows: 9 Temperature 320~ 9 N o r m a l pressure 100 kPa This results in the standard model: rs = 5.5 + 55.246(V) 0"4902

(5.16)

The master model thus obtained may be described by

r(v, 7") r(v, e)

"~Cetex(V, T, P) = ~

rs

~~'s rs

(5.17)

For both materials, the inter-ply slip model to predict inter-layer shear stresses may be condensed into an alternative general form as

T--(~TY(i)-~-kigni) "~s

(5.18)

The value of rs, yield stress ry(i) and of the power-law parameters k and n are given in tables 5.3 and 5.4. for APC-2 and Cetex respectively.

5.5. Friction during thermoforming In sheet forming of thermoplastic composite sheet, friction must occur due to the motion of the composite against the contacting surface which transmits the forming force. In the case of diaphragm forming, friction occurs between the composite and diaphragm, and between the diaphragm and tool surface. The study of this type of

TABLE 5.3 Inter-ply slip power-law model parameters for APC-2 Material: APC-2 rs = 1.0 + 28.708V~ i

ry

k

n

Conditions

1 2 3

1.0 0.95+(1.28 e-3)(P) 1.0 2.4

28.708 28.639+ 31.143 (log P) 28.708 79.14

-2.01 +(7.33 e-3)(T) 0.1635 + 0.3079 (log P) 0.8152 0.4471

360~
Velocity: 0 ~
198

A.M. Murtagh and P.J. Mallon

TABLE 5.4 Inter-ply slip power-law model parameters for Cetex fabric Material: Cetex 5-H satin fabric rs = 5.5 + 55.246 V 0"4902 i

ry

k

n

Conditions

1

5.5

-752.71 + 5.8641(T) -(1.043 e-2)(T)

-16.06 + 0.1019(T) -(1.5689 e-4)(T) 2

300~< T~<340~

2

5.5

11.52 + 0.5081 (P) -(7.086 e-4)(e) 2

0.5079(2.139 e-4)(e)

20 ~

3

5.5

55.246

0.4909

Velocity: 0 ~< V ~<0.5 mm/s

friction requires an understanding of the characteristics of the diaphragm and has been dealt with by Monaghan [30]. This research showed high coefficients of friction between the diaphragm (Upilex) and the steel tool, between values of 0.7 and 0.98 depending on temperature. This indicates the difficulty in allowing slippage of the diaphragm sheet over the mould surface. Diaphragm rupture occurred at some points where high shearing tractions were encountered. For friction of Upilex against APC-2 at melt temperatures, a Herschel-Buckley power-law model provided a good fit to the friction data, indicating the presence of a resin layer existing between composite and diaphragm film during forming. In press forming, matters are both more simplified and more complicated. Friction occurs between two mediums, the composite and tool surface. The anisothermal nature of press forming complicates matters as temperature has been shown to have a large effect on the friction of polymeric materials. In examining the friction that occurs during press forming, a number of parameters must be investigated, namely: 9 Interface temperature 9 Normal pressure 9 Fibre orientation 9 Mould surface/presence of a release agent As expected, these parameters are similar to those influencing other forming processes, for example as shown in section 5.4 for the inter-ply slip mechanism. The main difference is that friction is a surface characteristic of the material, and the other various shearing deformations occur internally within the composite. Two main types of friction have been identified for surfaces in contact: Coulomb friction and hydrodynamic friction [31]. Coulomb friction occurs between "dry" surfaces and in general, the frictional force is proportional to the applied normal force and independent of the sliding velocity. Hydrodynamic friction is a form of lubrication whereby a thin film of fluid exists between the two surfaces in question and viscous shearing of the film can occur in this region. In this case, sliding velocity may affect the frictional force. For polymeric composite friction, especially at high

Shear&g and frictional behaviour during sheet form&g

199

temperatures, hydrodynamic friction is dominant, due to the presence of a surface resin layer. However, a certain degree of Coulomb friction may occur wherever fibres come into direct contact with the hard surface. In terms of explaining the actual mechanisms of friction on a microscopic scale, two theories have been presented in the literature [32] - - the deformation theory and the adhesion theory. The deformation theory involves considering asperities from the hard (i.e. normally the tool) surface "ploughing" into the softer (polymeric matrix) material and the frictional force being related to the net loss in energy due to hysteresis as the material is strained elastically (see fig. 5.41). This net loss of energy can be related to the bulk visco-elastic properties of the polymer for a particular temperature, contact pressure and rate of deformation, rather than some special surface condition. With adhesion, it is thought that bonds are being continuously made, broken and re-made between contacting asperities from adjoining surfaces. The work needed to break these molecular bonds results in friction. The amount of adhesion, caused by mutually attractive van der Waals forces between the bodies, is dependent on the area of true contact which in turn is increased by van der Waals forces around the contacting regions. This effect is small when the normal load is large and increases in importance as the normal load is reduced. Strong adhesion is to be expected between "rubbery" materials, i.e. polymers above their softening temperature and other bodies. With rougher interfaces, adhesion is much reduced. Bartenev [33] treated rubber as a viscous fluid and introduced the concept of decrease in area of true contact (decrease in adhesion) as the speed increased, or as the temperature decreased. Bahadur and Ludema [34] lent further credence to the adhesion theory with this work relating sliding friction of polymers to visco-elastic properties and the strong connection between area of contact, adhesional shear strength and friction force. Previous work on friction of polymeric and fibre-reinforced materials have shown the influence of the parameters of temperature, normal load, sliding velocity and fibre orientation. Tanaka and Yamada [35] investigated the friction of polymeric materials, including PI, PES, PPS and PEEK, sliding against a smooth steel disc at various temperatures. Low coefficients of friction were observed at low temperatures,

Fig. 5.41. Ploughing of asperities in soft polymeric material.

200

A.M. Murtagh and P.J. Mallon

which increased rapidly as the temperature rose due to an increase in the deformation component at higher temperatures. At even higher temperatures, approaching or at the polymer melt temperature, a thick, transferred polymer layer was produced at the interface. This allowed direct comparison of frictional force with the shear strength of the material assuming that shearing occurs below the actual interface, in the polymeric material. Grosch [36] was able to describe the velocity and temperature effects on friction of various polymers, using a single master curve generated by the Williams-Landel-Ferry (WLF) transform to correlate velocity effects in terms of a universal temperature function, which is related to the glass transition temperature of the material. Friction tests carried out over a wide range of normal loads show that the coefficient of friction does not remain constant but increases as the load is reduced [37]. Thus, a simple linear relationship between frictional force and applied normal load does not exist. The type of relationship which can be used is of the form: F--

otW n

(5.19)

where c~ is a constant and n is some value less than 1. This is of similar form to the power-law model used to describe the shear velocity/shear stress relationship for inter-ply slip (see section 5.4). Over an extensive load range n is not a constant. If the adhesive mode of friction is assumed to be dominant, then the frictional behaviour of a material can be assumed to be dependent on the applied load/area of true contact relationship. This area of true contact has been shown to be difficult to measure in the past. It should not be considered to be the same as the sheared area undergoing inter-ply slip within the composite. The effect of surface fibre orientation on the friction of composites at elevated temperature has not been fully investigated. At room temperature, some researchers [38] have shown outer ply orientation to have little effect on friction. Friction coefficients were much more dependent on the resin surface characteristics. At higher temperatures, however, fibres may come to the surface and establish contact with the metal surface. In wear experiments using short glass-fibre-filled PPS [35], it was shown that at elevated temperatures, the fibres became exposed on the surface and were seen to be supporting the contact load. For hydrodynamic friction to occur during forming, a thin resin layer must be assumed to exist between the fibres and tool surface. Mould surface finish and the presence of surface coatings also need to be investigated to see how they may affect the frictional behaviour of the deforming composite. In most press forming operations, the mould surface in normally ground and polished. This is because regardless of any advantages that may result from being able to vary the surface roughness in certain circumstances [39] in the majority of cases the final surface finish is of paramount importance and for this reason, the mould surface must be as smooth as possible. Table 5.5 shows the results of surface roughness measurements taken from a typical mould surface used in thermoforming, a steel foil used during friction testing, and for comparison, measurements for APC2 prepreg in the longitudinal and transverse directions.

Shearing and frictional behaviour during sheet forming

201

TABLE 5.5 Surface roughness (Ra) measurements (in micrometres) Reading

Mould surface

Steel foil

APC-2 (long.)

APC-2 (trans.)

Steel foil (roughened)

1 2 3 4 5

0.6 0.7 0.8 0.7 0.7

0.22 0.18 0.25 0.22 0.22

1.3 1.2 1.25 1.5 1.2

6.0 6.5 5.8 6.0 6.1

0.9 0.85 0.8 0.82 0.8

Average

0.7

0.22

1.3

6.1

0.83

A 1-gm diamond paste was used to slightly roughen the surface of the foil and this consequently showed a similar value of surface roughness to the mould surface material. During actual friction testing [40], it was observed that there was little or no discernible difference in friction between the unroughened and roughened foil material, suggesting that any roughness differences at a level of less than 1 gm have little effect on friction of thermoforming composites. As expected, the surface roughness of the APC-2 prepreg is greater across the fibres than along them. The roughness value is similar in size to the diameter of a carbon fibre, which might be expected. Pre-consolidated laminates have a much smoother surface finish due to the presence of a surface resin layer. In many moulding operations, a surface release agent is applied to the surface of the mould which allows easy removal of any parts and can also improve the surface finish. Most of the agents work on the principle of depositing a solvent-soluble substance, e.g. a silicone, as a thin film a few molecules thick on the surface of the mould which acts as a barrier to prevent sticking and adhesion of the polymer to the moulding surface. In most cases, the release agent is applied using an aerosol spray or simply wiped on and allowed to dry. Multiple coatings may be applied. In terms of friction, release agents may also assist in reducing the degree of adhesion during sliding and so reduce the coefficient of friction. Two different experimental set-ups were used by Murtagh [40] to investigate the frictional properties of APC-2 under forming conditions. One method consisted of drawing a central specimen (composite or tool material) from between two sheets of the other material, all three being initially subject to a normal load between two heatable platens located between the jaws of a hydraulic press. Horizontal motion of the central specimen was achieved using the shearing apparatus already used for investigating the inter-ply slip mechanism explained previously in section 5.4. Using this set-up, isothermal conditions of between 25~ and 400~ were easily achievable and normal forces of between 0.1 kN and 100 kN could be applied. A shim placed at the rear of the pull-out sample maintained the alignment of the platens. Figure 5.42 shows a schematic of this apparatus. The other method used involved the development of a "friction sled', based on an ASTM standard [41], for obtaining frictional coefficients of plastic film and sheeting. This more accurately simulates the actual friction that occurs during press forming

202

A.M. Murtagh and P.J. Mallon

Fig. 5.42. Twin platen arrangement for friction testing.

as it mimics the anisothermal conditions and allows the concept of the two materials coming together suddenly, with the mould material being rapidly slid across the composite. It consists of an aluminium block with two embedded cartridge heaters, to which various test materials, in the form of plates, can be attached to the bottom surface (see fig. 5.43). The composite material is located on the surface of the consolidation apparatus already mentioned and is heated by conduction. Insulation is placed over the surface of the composite during initial heat-up and around the sled during testing to avoid excessive heat-loss. Weights may be applied on top of the sled to vary the normal load and the load range is lower than used with the twin platen arrangement. Again, the shearing apparatus is used to apply the sliding motion to the sled via a steel wire with a speed range of between 0.0125 mm/s and 4 mm/s. Friction force was measured using a 500 N load cell mounted at the front of the shearing apparatus and connected to the steel wire.

Fig. 5.43. Schematic of friction sled.

Shearing and frictional behaviour during sheet forming

203

Initial tests involved shearing a 50-mm wide sheet of 0~ APC-2 from between two sheets of steel foil, coated with Frekote release agent. The sample was placed between the platens of the consolidation apparatus and heated to testing temperature with 0.2 kN normal pressure being applied during heat-up. The results of this sample being sheared are shown in fig. 5.44. It was observed that due to the presence of this substantial normal load during heat-up that an adhesive bond tended to develop and this inhibited sliding from occurring until the shear strength of the bond had been exceeded at approximately 270 N shearing force. Following the breaking of this bond, the sliding force decreased to a steady level dependent on the shearing velocity characteristic of the assumed presence of a viscous resin layer between the composite and metal surface. This behaviour was typical of most tests carried out at high temperatures. In order to prevent the adhesive bond from being developed during heat-up, the test set-up was modified so that a slight gap was maintained between the top platen and the sample for most of the heating time. A brass plate, 6 mm in thickness, was placed over the sample which allowed heat-up via conduction through the bottom platen and when testing temperature had been reached, only then was the top platen brought into contact with test pressure being applied from the outset. Other research [37] has noted the effect of time of loading on friction due to the visco-elastic nature of the polymer material. In actual press forming, the laminate is not in contact with 350

300

.......

i

0.20

0.15 250 A A

z

200 v

_o

0.10

_8

L_

m

E E

150

t._

,00l

..2

t~

0.05

Shear load ] 50

Velocity

0

I

0

5

,

,,,

I

,

10

I

1

I

15

Displacement (ram) Fig. 5.44. Pull-out of 0 ~ APC-2 from between two sheets of steel foil.

0.00 20

204

A.M. Murtagh and P.J. Mallon

the mould until forming occurs so time of loading does not influence the frictional characteristics and it is important to eliminate this effect in friction testing. To examine the effects of surface temperature on the frictional characteristics of the composite/tool deforming system, the set-up using the twin platen/steel foil arrangement was used. In this case, isothermal conditions could be ensured throughout the sample, not just at the composite/tool interface. The temperature can be varied by adjusting the platen set-points and the appropriate effects observed. The friction sled set-up, although more accurately mimicking actual press forming conditions, as mentioned previously, is essentially a non-isothermal process, even in the event of both sled and composite having the same set-point temperature. The heat losses to the environment compared with the alternative platen set-up are quite large. In reality, the tool and composite are normally at different temperatures during actual press f o r m i n g - in the processing of APC-2 for example, the tool temperature may be as much as 150~ below the surface temperature of the composite just prior to forming. This allows more rapid cooling, and faster processing. Obviously, large thermal gradients are seen throughout the composite as it begins to deform after coming into contact with the mould and the frictional behaviour would be expected to vary accordingly. It has been not feasible to date to be able to cope with such anisothermal conditions. Instead, the effect at various isothermal temperature conditions has been observed and analysed. Figure 5.45 shows the influence of temperature on the friction of a 0 ~ ply of APC-2 as it was drawn from between two sheets of release-agent treated steel foil. The set 1.4 Normal load 0.2 kN 1.2

1.0

~,. ~ "6",~

0.8

9o ~

0.6

~

0.4

--o-

405 ~

--e~

385 ~ 3650

---e-:

0.0

I 0.0

0.5

1.0 Velocity (mmls)

Fig. 5.45. Effect of temperature on friction of 0~ APC-2.

1.5

,

t 2.0

345 ~ 25 ~

Shearing and frictional behaviour during sheet forming

205

point of both platens was varied for each test and a normal load of 0.2 kN was applied during shearing. The construction of the graph is similar to the method used for inter-ply s l i p - the behaviour is essentially viscous and various constant friction loads were measured as velocity increased. Rather than use shear stress as the measure of the ordinate, the coefficient of friction (It) was selected as it is the more traditional method of presenting friction results. Also, to present any result as a shear stress value would mean having to assume a value for the sheared area, which is to be considered invalid in any friction analysis. For the test performed at room temperature (25~ the frictional coefficient did not vary significantly with velocity and remains at a low value of approximately 0.16, similar to values obtained by Tanaka [35] at low temperatures. This non-variation with velocity change is indicative of simple Coulomb friction with a lubricating resin layer not existing at this point. As the temperature is increased to a value approaching the melt temperature, however, the friction behaviour begins to change significantly. The coefficient of friction increases and becomes more dependent on velocity. Tanaka [35] has shown an increase in friction for P E E K at lower temperatures than this (at the materials softening temperature, approximately at the glass transition temperature of 143~ The difference may be associated with different methods of testing (a steel sphere indenter compared with a flat plate interface) and the presence of a release agent in these tests. Above the melt temperature, in the temperature regions 365 to 405~ the friction coefficient rises rapidly from a common value of approximately 0.4 at low velocities up to a maximum of 1.2 at a velocity of 2 mm/s. This would seem to indicate a strong adhesive component as the P E E K matrix becomes molten and allows the real contact area (on a microscopic level) to increase substantially as temperature increases. It might have been expected that once a resin layer had become established at the interface that the friction force would be due totally to a shearing of the viscous fluid layer, and that increasing the temperature would cause shearing forces, i.e. frictional force, to reduce as the fluid viscosity decreased. This effect was observed previously with inter-ply slip behaviour. In this case, the opposite was observed, thus indicating that a strong adhesive bond develops between composite and tool surface as localised surface deformation occurs. Other factors such as a decrease in the resin layer thickness between the composite and steel foil as the temperature increases would also cause an increase in shearing force. Degradation of the release agent at higher temperatures may also be a factor. The influence of normal load was investigated in the range 0.1-0.75 kN. The lower limit was a function of the equipment available and the upper limit was considered a maximum that would occur on a laminate during any press forming operation, at least during the forming stage. Again, the results presented are for a 0 ~ ply of APC-2 sheared from between two steel foil sheets, coated with two layers of Frekote FRPNC release agent. Temperature was kept constant at 385~ for this series of tests. As shown in fig. 5.46, the coefficient of friction decreases as normal load is increased, a trend observed by other research [37]. The coefficient of friction varies from approximately 0.1 at low speeds (less than 0.25 mm/s) and a normal load of 750 N, to 1.4 at high speed (2 mm/s) and a low normal load (0.1 kN). It is to be observed that as the shear velocity increases, the coefficient of friction tends to level off.

206

A.M. Murtagh and P.J. Mallon

2.0

, NormalIo~ad:2NN 00100 1.5

--o-

Temperatu385~ re

500 N

_

n

A

Z

O

: ._u

1.0

0

0 "0

[~

0.5 .

,

0.0 0.0

0.5 1

9

.

.

.

.

1.0 I

_

.

.

.

L+

.

- -

1.5 .1_

a

2.0

Sliding velocity (mmls)

Fig. 5.46. Effect of normal load on friction of 0~ APC-2.

Shear thinning (a viscosity decrease at higher rates) would explain a certain decrease in frictional/shearing force. However, another explanation for this phenomenon may be that as the speed and displacement of the sample is increased, the lubricating resin layer does not remain at a constant thickness, but increases as there is a build-up of transferred polymer layer to the steel foil surface. This is also evidenced by visual inspection of the foil after pull-out, where a substantial polymer layer had adhered to the foil. Once the part has formed, any transferred polymer is reconsolidated into the surface of the part. Figure 5.47 shows further evidence of this. Again, a typical steel foil specimen was sheared from between two 0 ~ APC-2 plies at a constant velocity of 2 mm/s. The frictional load increased to a maximum of 215 N, then levelled off and began to decrease. If displacement is terminated for a period of time (at point A), the shear load relaxes similar to visco-elastic relaxation behaviour. When shearing was recommenced (at point B), the force recovers to a value at point C, approximately 40 N greater than the point at which movement had ceased. This would be the expected behaviour if the resin layer had become thinner again before sliding had recommenced. A sample of steel foil with the resin residue still intact on the surface was taken from a specimen which had been sheared from between two 0 ~ sheets of APC-2 [12]. The specimen had been sheared at two distinct velocities (0.25 and 2 ram/s) and two bands of resin were evident on the steel foil. Three identically sized samples were taken: (i) one from an unsheared part of the steel foil, with no resin, (ii) a sample

Shear&g and frictional behaviour dur&g sheet form&g 300

.

.

.

.

.

.

.

.

200

~9

207

4

100

:,=

0

0

5

10

15

20

25

30

Time (sec.)

Fig. 5.47. Frictional force reduction at high velocities/resin layer build-up.

from the part of the foil that had been sheared at 0.25 mm/s, and (iii) a sample that had been sheared at 2 mm/s. By measuring the mass increase due to resin adhesion, it was possible to calculate the average resin layer thickness on the foil samples, assuming a resin density of 1.3 g/cm 3 [42]. Table 5.6 shows that as expected, an initial average resin layer thickness of approximately 10 ~tm on each side of the foil exists at a shearing velocity of 0.25 mm/s. As the T A B L E 5.6 Resin thickness measurements of steel foil Sample

Mass (g)

Resin mass (g)

Resin volume

(cm 3)

Resin (~tm)

(i) (ii) (iii)

0.6032 0.60482 0.60671 .

Resin density (1.318 g/cm) 3 Sample area 4.931 x 1.219 cm 2

1.62 e -3 3.51 e -3

1.229 e -3 2.663 e -3

10.22 22.12

thickness/2

208

A.M. Murtagh and P.J. Mallon

speed is increased to 2 mm/s, the resin layer thickness more than doubles in thickness to 22 ~m and this would have the effect of reducing the frictional forces. These resin layer thicknesses compare with a value of approximately 6 ~tm assumed to exist between individual plies in the inter-ply slip process. In the actual press forming process, this resin layer increase-in-thickness during forming may be of practical assistance, as it reduces frictional forces and any tendency that may exist for sticking on the forming surface. Even though there may be a local build-up of resin on the surface of those areas that move across the mould, it is reasonable to assume that during pressure application and consolidation of the part that any transferred polymer is re-percolated into the composite and accommodated amongst the fibres. To further investigate the presence of a surface resin layer, two micrographs were taken through a section of a tested specimen, one at 90 ~ and one at 0 ~ to the fibre direction, cooled under pressure to maintain an adhesive bond between the composite and steel foil, with no release agent [12]. Figure 5.48 shows that a thin layer of resin does exist at the boundary of the composite (APC-2 orientated at 90 ~ to the plane of observation) and steel foil/mould surface. Here, the resin layer seems to be of relatively constant thickness, compared with the interface in an actual laminate, which indicates the lack of fibre/fibre interference with frictional flow against the smooth foil surface. Shown inset is a section from a specimen with fibres running at 0 ~ At point A, a fibre can be seen to approach the steel surface, showing that fibre/ tool contact may be possible. It should again be emphasised here that the thickness shown in this micrograph is for a specimen which has interacted with the steel foil,

Fig. 5.48. Micrograph of tested specimen showing resin interlayer, fibres at 90 ~ (inset: fibres at 0~

Shearing and frictional behaviour during sheet forming

209

then cooled to room temperature. It does not conclusively illustrate the resin layer that is developed during processing. The effect of surface fibre orientation must also be taken into account in examining frictional characteristics of thermoplastic composites. Despite the fact that a thin resin layer is assumed to exist between the composite and tool, the presence and relative direction to flow of the fibres may affect the flow properties of the resin layer and indeed, the tool may come into contact with fibres in certain locations where resin "squeeze-out" has occurred. In terms of testing, the most suitable set-up to use in investigating the effect of fibre orientation was the friction sled [40]. With the twinplaten set-up, it proved too difficult to mount any specimen at the appropriate angle or to apply a traction force to any ply at an angle other than 0 ~ Care had to be taken that isothermal conditions existed for the sled tests. Set-point temperature of the sled and bottom platen was 385~ and 400~ respectively for all tests using APC-2. The higher temperature of the bottom platen was to allow for conduction through to the surface of the composite and to partly compensate for heat loss. Figure 5.49 shows a plot of sliding velocity versus coefficient of friction for a variety of different surface fibre orientations, relative to the direction of sliding. The highest resistance to sliding occurred when the fibres lay parallel to the sled. As the fibre angle changed, friction decreases, down to a minimum for a sample with fibres lying at 90 ~ or perpendicular to the sliding direction. Figure 5.49 shows that, even for an angular change from 0 ~ to 10~ the frictional force is significantly reduced. Therefore, it appears that although a resin layer is assumed to exist between the fibres and mould surface, the fibre orientation does have an effect on the 1.0

0.8 Fibre angle

0 _

~"

O.6

0~

T-

---4--

10 ~

--0--

90 ~

45 ~ "~ 'i5

0.4

tJ

0.2

0.0

,

0.0

I

0.5

,

....

9

,

1.0

.... I

1.5

Sliding velocity (mmls)

Fig. 5.49. Effect of surface layer fibre orientation on friction of APC-2.

,

.I

2.0

210

A.M. Murtagh and P.J. Mallon

frictional behaviour. The apparent higher resistance to movement for the 0~ tated fibres compared with 90~ is analogous to what was observed with the inter-ply slip behaviour of APC-2 (see section 5.4). However, fibre/fibre interactions between different layers cannot happen with a flat mould surface-composite interface. Therefore, as proposed by Groves [21], the flow behaviour and viscosity of the resin material is influenced by the fibre orientation ~ the longitudinal viscosity may be an order of magnitude above the transverse viscosity for measurements carried out on APC-2 using a torsional rheometer. In terms of the frictional behaviour of APC-2 against a mould surface above the melt temperature of the composite, the orientation of the fibres affects how replacement resin material is allowed percolate to the interface region, as the resin layer is being continuously sheared along the mould surface. Flow over 90 ~ fibres is potentially easier than resin flow along the fibre direction. Figure 5.50a and fig. 5.50b shows a possible mechanism for this; the flow of resin is less restricted when the fibres are perpendicular, or at some angle other than 0 ~ Various release agents were also investigated to determine their effect on the frictional behaviour of unidirectional APC-2 on stainless steel. It should be noted here that not all these agents (Frekote FRP-NC, Wurtz PAT 807-B, Wurtz PAT 808 and ChemTrend E274) were regarded as suitable for giving optimum results at such high temperatures (approaching 400~ Also, surface finish may be a more important quality in selecting the appropriate agent as opposed to reducing frictional forces to a minimum. For all these tests, two coats were applied to an acetonecleaned steel foil sample and allowed to dry. For the E274 agent, the foil had to be pre-heated to 400~ before the agent was applied using an aerosol. However, in practice, this agent has the advantage of not requiring as many "touch-ups" during repeated use compared with other agents. Normal load was 0.2 kN and temperature 385~ for all tested samples. As expected, fig. 5.51 shows that the highest coefficient of friction was recorded for an untreated foil sample. Frekote FRP-NC and E274 gave similar results. PAT-808

Fig. 5.50. Resin flow depending on fibre orientation.

Shearing and frictional behaviour during sheet forming 1.5

A Z

211

--

10

a-.O m 4..0

q.. O I:

05

No release agent Frekote NC

Temperature 385~ Normal load 0 2 kN ,

0.0

00

I

05

'

PAT 807B

A

PAT 808

",

I

,

E274 i

1.0

[

15

i

I

20

Sliding velocity (mmls) Fig. 5.51. Effect of surface release agent on friction of APC-2. (due to the fact that it dries to leave a thicker coat compared with the other agents) gave the best results. However, when compared with PAT-807B (slightly higher coefficient of friction), PAT-808 resulted in poor surface finish q u a l i t y - the surface had a matt rather than a glossy finish and tended to be of uneven quality, in typical, press-formed APC-2 parts [43]. Friction tests have been carried out on Cetex 5-H fabric (carbon-fibre-reinforced PEI) material to measure the material's frictional behaviour as a function of temperature, normal load and other process parameters [12]. Again, the twin-platen test set-up was used, with a sample of the composite being sheared from between two sheets of stainless steel foil, pre-treated with two coats of Frekote release agent. No significant normal load was applied to the sample prior to t e s t i n g - heat-up to test temperature was done with a slight gap between the platens. Normal load was applied only for two minutes prior to commencement of any test. Fibre straightening in the fabric sample was accounted for by allowing a known yield load to be developed in each sample before sliding would occur. This value was measured and defined to be 55 N for each test carried out. The significance of this was to exclude any internal stresses in the test sample, and to concentrate on surface phenomena only. Figure 5.52 shows results for a series of tests carried out at various test temperatures between 300~ and 340~ for Cetex. This shows a reduction in measured frictional forces as the temperature is increased. This is exactly the opposite trend as observed with APC-2 (fig. 5.45), where friction increased with temperature.

212

A.M. ~urtagh and P.J. Mallon

1.0

0.8

A

Z

0.6

-

c

._o u o..

o

0.4

C

300~

~

o

,m

o

0.2

0.0

t

0.0

... t,

310oc

..... =

320~

o----

,

340~

Normal load 2 5 0 N ,

I

0.5

,

. . . . .

,

. . . .

I

,

,

,

1.0

I

1.5

,

~

,,

2.0

Sliding velocity (mmls)

Fig. 5.52. Effect of temperature on the friction of Cetex. Intuitively, one would expect friction to decrease if the viscosity in the intervening resin layer between the surfaces decreased with an increase in temperature [40]. The difference between the unidirectional material (APC-2) and the fabric may be explained by the possible development of the resin layer at the interface for APC2 or Cetex. In the unidirectional material, the volume fraction of the fibres is high, and compared with the under/over pattern of the fabric weave, offers a relatively "flat" surface to be offered to the mating steel surface. Consequently, the resin layer exists as a thin layer between the fibres and increasing the temperature may cause more fibre/steel surface contact if the resin were to flow back into the composite. For the fabric material, the amount of maximum fibre/steel contact is limited to the contact points between the highest points in the weave and the mould surface. At the same time, the amount of resin present at, or close to, the interface is higher than for unidirectional materials due to "pools" of resin lying in the interstices or low points on the surface of the weave. An increase in the temperature of the resin would lower the viscosity and reduce shearing forces in the interlayer. The adhesional properties of PEEK are known to be high, whereas PEI may not develop the same bond strength at such high temperatures. At the same time, the lower temperatures involved with the Cetex material may mean that the Frekote release coating may not degrade as rapidly as with typical APC-2 processing temperatures (,~400~ which would cause an increase in friction force. Using Frekote at such high temperatures is approaching the limit of the coating's effectiveness. Further tests were carried out to examine normal loading effects on the friction of Cetex. Similarly to APC-2 and other polymeric materials, the coefficient of friction

Shearing and frictional behaviour during sheet forming

213

decreases as normal load in increased. The lower bound for testing, 100 N, resulted in the highest level of friction. Figure 5.53 shows a plot of sliding velocity versus coefficient of friction for a variety of normal loading conditions, between 100 N and 500 N for Cetex. As with APC-2 (see fig. 5.46), a reduction in the coefficient of friction for Cetex is seen at higher sliding velocities, again possibly due to the development of a thicker resin interlayer as sliding progressed. Given the obtained experimental data, it should prove feasible to develop a predictive model for friction dependent on forming conditions. This has been achieved for APC-2 and Cetex 5-H satin fabric material [12]. A similar method was used to that for inter-ply slip since the power-law model form is again valid due to the viscous effects of the resin layer between tool and composite, which is analogous to the presence of a resin layer between plies for inter-ply slip. The relationship between friction coefficient, sliding velocity and the forming parameters (temperature, normal load and surface fibre orientation) can be described by the following:

air'i)

(5.20)

#s

~-i=l

/Zs

tz defines the coefficient of friction, V is the velocity in mm/s and/Zs is the coefficient of friction under standard conditions of temperature, normal load and fibre direction. The value of lZs and of the power-law parameters a and b are given in table 5.7 and table 5.8 for APC-2 and Cetex respectively.

1.4

Temperature3200C

i

......

1.2 ~"

1.0

cr

o

g

0.8

o

0.6

~

"O

~

100N

o.4

o

250N 0.2

500N

0.0 0.0

0.5

1.0

Sliding velocity (m mls) Fig. 5.53. Effect of normal load on the friction of Cetex.

1.5

2.0

214

A.M. Murtagh and P.J. Mallon

TABLE 5.7 Friction power-law model parameters for APC-2 Material: APC-2 /zs = 0.9276V~ i

a

b

Conditions

1

-34.136 + 0.172(T) -(2.1 e-4)(T) 2

(3.36 e-3)(T) - 0.874

345 ~
2

3.52- 1.127log(N)

0.686-0.116 log(N)

100~
3

0.9276 -- 0.0743(0)0.2606

0.4198 + (2.827 e-4)(0)

0~<0~<90~

Velocity: 0 ~
TABLE 5.8 Friction power-law model parameters for Cetex fabric Material: Cetex 5-H satin fabric /Zs = 0.868 V 0"4 i

a

b

Conditions

1

33.656 - 0.1917(T) +(2.79 e-4)(T) 2

-13.192 +(8.7525 e-2)(T) -(1.3375 e-4)(T) 2

300~< T~<340~

2

1.6687 - (4.026 e-3)(N) +3.293 e-6(N) 2

0.2348 + (1.229 e-3)(N) -2.2733 e-6(N) 2

100 ~
3

0.868

0.4

Velocity: 0 ~
References [1] Cogswell F.N, "The Processing Science of Continuous Fibre Reinforced Thermoplastic Composites", Intl. Polymer Processing, 1, 4, pp. 157-165, 1987. [2] Gutowski T.G., "A Resin Flow/Fiber Deformation Model for Composites", SAMPE Quarterly, 16, pp. 58-64, 1985. [3] Wheeler A.J., Ph.D. Thesis, University of Wales, Aberystwyth, 1990. [4] Lam R.C., Kardos J.L., "The Permeability of Aligned and Cross-Plied Fiber Beds during Processing of Continuous Fiber Composites", American Society for Composites, 3rd Annual Tech. Conf., pp. 3-11, 1988. [5] Cogswell F.N., from Thermoplastic Aromatic Polymer Composites, Butterworth-Heinemann, Oxford, 1992. [6] Barnes J.A., Cogswell F.N., "Transverse Flow Processes in Continuous Fibre Reinforced Thermoplastic Composites", Composites, 20, 1, pp. 38-42, 1989. [7] Mulholland A.J., Monaghan M.R., Mallon. P.J. "Characterisation of Consolidation Flow Processes in Continuous Fibre Reinforced Thermoplastic Composites", 13th Intl. Conf., European Chapter, SAMPE, Hamburg, 1992. [8] Spencer A.J.M., from Deformations of Fibre-reinforced Materials, Clarendon Press, Oxford, 1972.

Shearing and frictional behaviour during sheet forming

215

[9] Rogers T.G., "Rheological Characterisation of Anisotropic Materials", Composites, 20, 1, p. 21, 1989. [10] Groves D.J., Bellamy A.M., Stocks D.M., "Anisotropic Rheology of Continuous Fibre Thermoplastic Composites", Composites Manufacturing, 2, 2, 1992. [l l] Scobbo J.J., Nakajima N., "Dynamic Mechanical Analysis of Molten Analysis of Molten Thermoplastic/Continuous Graphite Fiber Composites in Simple Shear Deformation", 21st Intl. SAMPE Tech. Conf., pp. 730-743, Anaheim, California, 1989. [12] Murtagh A.M., "Characterisation of Shearing and Frictional Behaviour in Sheetforming of Thermoplastic Composites", Ph.D. Thesis, University of Limerick, May 1995. [13] Bergsma O.K., "Computer Simulation of 3-D Forming Processes of Fabric Reinforced Plastics", ICCM-9, Madrid, 1993. [14] Van West B.P.,"A Simulation of the Draping and a Model of the Consolidation of Comingled Fabrics", CCM Report 90-07, Center for Composite Materials, University of Delaware, 1990. [15] Johnson A.F., "Rheological Model for the Forming of Fabric Reinforced Thermoplastic Sheets", Composites Manufacturing, 6, 3-4, pp. 153-160, 1995. [16] Murtagh A.M., Mallon P.J., "Shear Characterisation of Unidirectional and Fabric-Reinforced Thermoplastic Composites for Pressforming Applications", ICCM-10, Whistler, Vancouver, 1995. [17] Blanlot R., Billoet J.L., Gachon H., "Study of Non-Polymerised Prepreg Fabrics in 'Off-Axes' Tests", ICCM-9, Madrid, 1993. [18] Scherer R., Friedrich K., "Experimental Background for Finite Element Analysis of the Interply-Slip Process During Thermoforming of Thermoplastic Composites", ECCM 1994, Stuttgart, 1990. [19] Scherer R., Zahlan N., Friedrich K., "Modelling the Interply-Slip Process During Thermoforming of Thermoplastic Composites Using Finite Element Analysis", Proc. CADCOMP 90, Brussels, 1990. [20] Xiang Wu, "Thermoforming of Thermoplastic C o m p o s i t e s - Interply Shear Flow Analysis", 21st International SAMPE Tech. Conf., p. 915, Anaheim, California, 1989. [21] Groves D.J., "A Characterisation of Shear Flow in Continuous Fibre Thermoplastic Laminates", Composites, 20, 1, p. 28, 1989. [22] Jones R.S., Oakley D., "An Interpretation of Rheological Data in Terms of Model Systems", Composites, 21, 5, p. 415, 1990. [23] Kaprielian P.V., O'Neill J.M., "Shearing Flow of Highly Anisotropic Composite Laminates", Composites, 20, 1, p. 43, 1989. [24] Soil W.E., "Behaviour of Advanced Thermoplastic Composite Parts", M.S. Thesis, Dept. of M.E., MIT, 1987. [25] Tam A.S., Gutowski T.G., "Ply-Slip During the Forming of Thermoplastic Composite Parts", J. of Composite Materials, 23, p. 587, 1989. [26] Morris S.R., Sun C.T., "An Investigation of Interply Slip Behaviour in AS4/PEEK at Forming Temperatures", Composites Manufacturing, 5, 4, p. 217, 1994. [27] Murtagh A.M., Monaghan M.R., Mallon P.J., "Investigation of the Interply Slip Process in Continuous Fibre Thermoplastic Composites", ICCM-9, Madrid, 1993. [28] Murtagh A.M., Monaghan M.R., Mallon P.J., "Development of a Shear Deformation Apparatus to Characterise the Interply Slip Mechanism of Advanced Thermoplastic Composites", IMF-8, University College Dublin, 1992. [29] Muzzi J., Norpoth L., Varughese B., "Characterisation of Thermoplastic Composites for Processing", SAMPE Journal, 25, 1, p. 23, 1989. [30] Monaghan M.R., Mallon P.J., "Study of Polymeric Diaphragm Behaviour in Autoclave Processing of Thermoplastic Composites", 14th Intl. Conf., European Chapter, SAMPE, Birmingham, 1993. [31] Barone M.R., Caulk D.A., "A Model for the Flow of a Chopped Fiber Reinforced Polymer Compound in Compression Moulding", J. of Applied Mechanics, 53, p. 361, 1986. [32] Tabor D., "Friction, Adhesion and Boundary Lubrication of Polymers", ACS International Symposium on Polymer Wear and its Control, Plenary Lecture, Los Angeles, 1974. [33] Bartenev G.M., El'kin E.I., "Friction Properties of High Elastic Material", Wear, 8, 8, 1965. [34] Bahadur S., Ludema K.C., "The Viscoelastic Nature of the Sliding Friction of Polyethylene, Polypropylene and Copolymers", Wear, 18, p. 109, 1971.

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[35] Tanaka K, Yamada Y., "Effect of Temperature on the Friction and Wear of Some Heat-Resistant Polymers", from Polymer Wear and its Control, American Chemical Society, Washington DC, 1985. [36] Grosch K.A., "The Relation between the Friction and Visco-elastic Properties of Rubber", Proc. Roy. Soc., A274, p. 21, London, 1963. [37] Bowden F.P., Tabor D., "The Friction and Deformation of Polymeric Materials", from The Friction and Lubrication of Solids, Part XIII, Clarendon Press, Oxford, 1964. [38] Herrington P.D., Sabbaghian M., "Factors Affecting the Friction Coefficents between Metallic Washers and Composite Surfaces", Composites, 22, 6, 1991. [39] Throne J.L., from Thermoforming, Hanser Publishers, Munich, 1987. [40] Murtagh A.M., Lennon J.J., Mallon P.J., "Surface Friction Effects Related to Pressforming of Continuous Fibre Thermoplastic Composites", Composites Manufacturing, 6, 3-4, p. 169-176, 1995. [41] ASTM Standard 1894-90,"Standard Test Method for Static and Kinetic Coefficents of Friction of Plastic Film and Sheeting", 1990. [42] ICI Thermoplastic Composite Handbook, 1992. [43] Maher E.J., "An Investigation of Isothermal Pressforming with Thermoplastic Composite Materials", M.Eng. Thesis, University of Limerick, Ireland, 1994.