Tow placement studies for liquid composite moulding

Composites: Part A 30 (1999) 1105–1121

Tow placement studies for liquid composite moulding C.D. Rudd*, M.R. Turner, A.C. Long, V. Middleton Department of Mechanical Engineering, University of Nottingham, University Park, Nottingham NG7 2RD, UK Received 27 July 1998; accepted 27 January 1999

Abstract The paper describes an investigation of design and manufacturing techniques for preform manufacture based upon tow placement. A kinematic draping algorithm was used to devise flat laydown patterns which were deposited in dry fibre using tow placement with a four-axis robot. The equipment was developed to lay multiple tows and used to test an “undraping” algorithm for the generation of novel architectures. The latter were applied to two simple shapes in attempts to provide more uniform coverage than conventional fabric-forming processes. Models to predict the permeability and elastic modulus of laminates produced by tow placement facility were developed which allow for fibre waviness and were validated by the deliberate manufacture of preforms with sinusoidal fibre paths. Practical applications are illustrated for demonstrator structures from the marine and aircraft industries. q 1999 Published by Elsevier Science Ltd. All rights reserved. Keywords: Tow placement; A. Fibres; E. Preform

1. Introduction The design and manufacture of high quality, cost-effective fibre preforms is a continuing problem for the high-fibre fraction, aligned fibre structural components common to the aerospace industries, although the same principles apply equally to the automotive and marine sectors. Automation of composites manufacturing processes is increasingly viewed as the means by which low cost, high-quality components and structures can be produced. While short fibre technologies pose relatively few automation problems, the handling problems associated with long and continuous fibre composites mean that progress in the latter area has been relatively slow. In aerospace, where pre-preg handling continues to dominate, some aspects of fabrication have been mechanised or automated including the introduction of automated ply cutting based upon CNC manipulators fitted with lasers, water-jets or ultrasonic knives. Pre-pregs or (with some difficulty) dry fabrics may be cut accurately from roll stock to produce kits of net-shaped plies. The same techniques may be combined with ply nesting software so that wastage is minimised and, later, with laser projection guides to reduce positional errors as the materials are stacked in the mould. Similar concepts have been applied to the preparation of fabric plies by Sarhadi and co-workers [1] who linked a laser cutting device with an electrostatic pick-and-place device and a robotic stitching head. The * Corresponding author. Fax: 1 44-0115-951-3784.

robotic tacking produces a handleable preform with automatically positioned plies reducing manual error. Although such systems are non-dedicated, capital costs remain relatively high and the equipment is restricted in the fabrics which can be handled effectively. A further step towards full automation involves the substitution of a tape placement head for the manual layup of pre-preg. This provides a cost effective route for fabricating large parts with relatively simple contours and has been applied for both thermosets and thermoplastics pre-pregs. This has potential to reduce the labour cost associated with manual layup and has been used with some success in the North American aircraft industry for fabricating large composite structures. However, in the majority of applications the continued use of pre-preg involves both a cost penalty and limitation on geometric complexity compared with dry fabric forms. A natural extension of the tape placement approach is to substitute a dry or tacky tow for the pre-preg tape. Tow Placement offers greater versatility and (for dry tows) reduced materials costs since the process is no longer limited by the transverse and shear compliance of a tape and (for tow-preg) the materials handling problem is simplified. Again, a roller is used to provide a compaction pressure and to enable concave surfaces to be followed. Several industrial studies [2–8] have been reported from North American establishments (although many of these relate to tow-preg handling) and a variety of structures for aerospace and defence applications have been produced in this way including engine inlet cowlings, fuselage sections, rotor

1359-835X/99/$ - see front matter q 1999 Published by Elsevier Science Ltd. All rights reserved. PII: S1359-835 X( 99)00 010-X

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Fig. 1. Schematic of tow placement facility.

blade spars and fuel pods. The fibre handling technology used in these applications is generally proprietary and few results have been released concerning the quality of the fibre placement and distribution. However, the consensus view emerges that for large structures, the potential savings in materials costs (since the tailoring process has been eliminated), the tow placement route is an attractive one, despite large capital equipment costs (estimated at US$2 million in 1997 [8]). Given the above background, the present study sets out to address some of the arising problems. The concept involves a flat fibrous laydown which is formed subsequently into a three-dimensional shape using matched dies or diaphragm forming. Tow placement in two, rather than three dimensions simplifies the handling problem and reduces set-up costs. A kinematic draping algorithm [9] was applied to define the net-shape, two-dimensional fibre pattern and this was the subject of subsequent work. The objective was to identify the practical difficulties associated with dry tow placement and to develop actuation techniques and simple control strategies for producing single ply and multi-ply, aligned fibre lay-downs. The practical issues to be addressed included the control of fibre distribution and its effect on laminate properties and the major novelty, the development and manufacture of non-standard fibre architectures for optimal forming and properties. Thus the work involved the commissioning of a CAE-based tow placement facility and the identification of techniques for dry fibre deposition, retention and consolidation. Studies were performed on cycle time and material utilisation and on the processing and performance of the arising laminates. Demonstrator structures, drawn from industry, were used to test the effectiveness of the system.

2. Tow placement facility The facility was based around a Crispen 2015/4-axis CNC manipulator (SOS Newall Ltd.) and the design, construction and development is described in detail by McGeehin [10] and Turner [11]. The main goal was an automated facility capable of producing two-dimensional net-shaped preforms at acceptable cycle times. Fibre

laydowns were laid flat prior to stacking and forming. Two approaches to tow placement were assessed: airjet propulsion onto a porous bed with suction assistance and roller assisted placement using both suction and tackifiers to assist retention. Both methods are described in detail below. The laydown machine is illustrated in Fig. 1. X and Y movements were achieved with friction drive units. Two units are used for X movements with one each side of the gantry to reduce large moments on the bridge or guides. The X and Y axes drove the bridge at speeds of up to 1.67 m/s with positional accuracy of ^ 1 mm. The Z-axis was operated by a lead screw and the A-axis by a belt drive. The rotational accuracy of the A axis was ^ 0.18. 2.1. Airjet deposition Airjet propulsion relied on a device carried on the gantry of the manipulator which was similar to those used in the weaving industry and provided rapid deposition, comprising two concentric nozzles, the inner feeding the tow and the outer supplying air at elevated pressure for propulsion. This was later replaced since the wake from the air stream tended to displace fibres from the specified paths. As a result, a multi-tow roller placement device evolved as described below. However, because of the modularity of both designs, it was feasible to change from airjet placement to roller placement in a relatively short time. Although it will be shown that roller placement is more accurate and creates a more uniform laydown, airjet placement is advantageous in creating random fibre laydowns at high speed which is well suited to composite structures in non-critical applications such as automotive body panels. Since the tow exit velocity was controlled by the air supply pressure and friction in the payout system, failure to match this precisely with the linear velocity of the payout head resulted in departure from the tow path specified. Thus a separate servomotor device was designed to provide control payout speed. A series of rollers driven by a MacLennan M372TE 7.4 N cm dc servomotor delivered tows to the airjet at a specified and reliable rate. The tension applied to the tow by the airjet created sufficient friction to allow accurate velocity control. The drive motor was connected to a Norwen Electronics NE-P-061 amplifier board and the system was calibrated over a range of input voltages by measuring the fibre length deposited over a 60 s period. The tows were cut to length at the end of each pass during this and subsequent trials using a pneumatic guillotine attached to the base of the airjet. The airjet exit was positioned between 5 and 8 mm from the laydown bed to minimise fibre disturbance and was operated at 0.75 bar (gauge). Where suction assistance was used, a 0.36 kW axial flow fan was operated in a plenum beneath the porous bed. Fibres were retained on the bed using tackifiers so that the preforms could be handled and impregnated without any serious fibre re-orientation. Aligned fibres were fixed

C.D. Rudd et al. / Composites: Part A 30 (1999) 1105–1121

Fig. 2. Multiple tow placement head.

using Fothergill Engineered Fabrics M0924/532 Admesh, a 15.5 gm 22 plain weave polyester fibre scrim with polymer/ latex binder and polyester resin finish which is tacky at room temperature. Random reinforcement was bound using Neoxil 940 bisphenolic polyester powder binder. The powder was sprinkled onto the reinforcement after laydown using a hand held “pepperpot”. Consolidation of the laydowns was required to reduce bulk and here a hand held domestic iron was used for convenience. Clearly, there remains scope to mechanise both of these operations.

constructed from free running roller bearings and actuated pneumatically. Tow velocity was regulated by the vector sum of the velocities of the X and Y axes of the manipulator. Feedback using the tachogenerator and encoder on the pinch roller motor allowed for control of tow straightness. Because of the irregular profile of the majority of applications it was necessary to arrest one or more tows whilst others continued to be laid. To prevent tangling of the inactive tows during this period, a system of brakes was developed which operated by the simple action of an eccentric cam. Again, cutting relied upon five pneumatically actuated guillotines located between the fibre payout system and the compaction roller. Because no airjet was used the problem of tow displacement due to wash was eliminated and the porous manipulator bed was replaced with a solid table constructed from 18-mm fibreboard. Tows were retained on the bed using the scrim binder described above. Although this was tacky at room temperature, its effectiveness increased with temperature and two heaters were incorporated with a 2500 W 400 mm × 600 mm platen heater set into the bed and a 550 W infrared emitter strip on the placement head. Both heaters were controlled independently to heat the binder to approximately 508C. Consolidation of these preforms was done continuously by the action of the compaction roller, with the compliance increased by the local heating. 2.3. System characterisation

2.2. Roller deposition Following difficulty controlling tow displacement by exhaust air a multi-tow roller placement device was commissioned with five independent feeds (Fig. 2). The roller head was modular and attached to the same station as the airjet head. A sprung compaction roller pinched the tow against the laydown bed. The steel roller was free running with bearing end caps and 12 × 0.1 mm air bleed slots along its length to encourage tow separation. A small, curved guide in front of the roller prevented the fibres from buckling as they hit the laydown bed. Again, a metering system was used to control the tow delivery. A 0.2 N m MacLennan M286 dc servomotor with tachogenerator and encoder drove a pinch roller which delivered the tows via a set of stainless steel tubes to the compaction roller. The pinch roller was manufactured from Nylatronw. Delivery was governed by a set of five clutches which pinched the tows against the driven pinch roller. The clutches were Table 1 Tow straightness and elastic properties arising from air-jet and roller placement

RMS Deviation (mm/m) Average fibre misalignment (8) Modulus (GPa) Standard deviation (GPa)

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Roller

Airjet

Quasi-08 fabric

1.26 0.5 27.0 1.41

2.53 2.0 24.9 2.07

n/a n/a 25.0 1.3

The system performance was measured first by comparing the straightness of the as-laid tows using a series of unidirectional preforms from both laydown techniques and a commercial fabric reinforcement. Laydown accuracy was measured in several ways. Single tows of glass fibre were laid at different speeds and length and image analysis used (from still photography) to calculate tow waviness. Complete preforms were then consolidated and processed by RTM. A 50-tonne upstroke press was used for consolidation and clamping, being equipped with a pair of 3.5 kW electrically heated aluminium platens. The plaque mould was sealed with a 4 mm O-ring around the periphery of a 3.6 mm × 518 mm × 538 mm cavity. A thermal break valve controlled the injection of resin (using a pressure pot operating at 5 bar). Injection (for the unidirectional plaques) was via a strip gate along one edge of the cavity with venting along the opposite edge. This minimised drag forces normal to the fibres and reduced the danger of fibre washing. A 1.5 mm deep gallery at the injection side heated the resin prior to entry to the cavity proper and helped to promote rectilinear flow. Pressure transducers and thermocouples were set into the tool and these were used to determine mould fill and resin cure. The specimens were produced at 808C in the aluminium tool using LY 5052 (Ciba Polymers) epoxy resin and post cured for one hour at 1008C in a hot air oven prior to testing. Fibre misalignment was also measured from polished micro-sections of the mouldings.

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Fig. 3. Design and manufacturing concept based upon two dimensional tow placement.

Table 1 compares the results from the analysis and tensile testing of plaques produced from airjet and roller placed tows with commercial fabric (Tech Textiles ELP-b 567) at approximately 30% fibre by volume. The average deviation of the 08 fibres suggested that laydown speed and length had little influence on straightness over the test range, but significantly, the airjet placed tows had almost twice the deviations of those placed by roller. This was attributed to displacement by the wake of the airjet and these results are supported by the misalignment angles for the rollerlaid tows following microscopy. Mechanical testing produced consistent results which are also summarised in Table 1. The tensile modulus of the laminates produced using roller placement was superior to that arising from both the airjet process and control specimens produced from non-crimp fabric. However, it should not necessarily be inferred from these results that the roller placement resulted in superior alignment to the warp knitting process since the fabric samples contained approximately 5% transverse yarns to improve handling stability.

2.4. Control strategies The manipulator was controlled by a separate, hard-wired unit. In order to place tows to a predetermined scheme, a CNC part program was required. Three approaches were used for generating part programs: 1. Two-dimensional CAD drawings were post-processed along with a preform specification (ply orientations, tow spacings) to generate a part program for tow placement. The laydown geometry (in IGES format) was loaded into a proprietary post-processor and the existing outline was edited as necessary (to add a trimming allowance for example) before defining the preform specification. The part program was output as an ASCII file and transferred directly to the memory of the manipulator control unit via a standard RS232 connection. The process and validation is illustrated in Fig. 3 and described in more detail below. 2. Kinematic draping [12] was used to generate a template

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Fig. 4. Kinematic drape model output for hemisphere quadrant.

or outline which provided the basis for the net-shape part programs. The drape software describes the deformed geometry of the nominal fabric or laydown in threedimensional space. A mapping algorithm was then applied to flatten the elements or “patches” in the surface model to produce the flat template. This geometry was then post processed as before to generate the NC part program which enabled the net-shape laydown to be produced. 3. Surface models were created and a uniformly spaced fibre pattern applied using CAD. Kinematic draping, as above, was then used to map the fibres onto a horizontal surface, thus producing an “optimised” laydown that when formed would contain the desired fibre paths in three-dimensional space. This method is described in more detail below.

3. Drape modelling During the forming stage, the fibre architecture will

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change to account for the different geometry and surface area. For simple components (i.e. for single curvature components or where the amount of draw is small), this change will be trivial. However if a two-dimensional reinforcement is formed to a complex shape (i.e. one of large draw or sudden change in geometry) the fibre re-orientation can be substantial. The dominant form of fibre deformation in directional reinforcements is generally assumed to be by inter-fibre shear although some fibre slip (which is usually defined as a change in inter-tow spacing) may also occur. During shear, the inter-fibre angle (for example the angle between the fibres of the two principal directions of a biaxial reinforcement) will change. In severe cases this interfibre angle can reduce until the reinforcement locks and will no longer deform by shear. Further deformation will result in fibre buckling (reinforcement wrinkling) thus producing a defective preform. The change in fibre orientation from that of the two-dimensional reinforcement to the three-dimensional preform will affect subsequent processes including resin flow, structural properties and formability. Such effects have been reported in previous studies by several authors (e.g. [13]) and a common approach to the anticipation of such effects relies upon kinematic drape modelling, where, briefly, the reinforcement is idealised as a pin-jointed structure. The application of drape modelling in this work was, initially, to predict the blank shape required to drape to the final geometry without trimming and secondly, to generate a laydown pattern which, after post-forming, would take up a predetermined architecture (in this case, that of uniform fibre spacing and porosity). This latter case is explained in detail in subsequent sections. While the application of draping algorithms has been explored in detail for the common forms of textile reinforcement the structure arising from tow placement is unusual in that no interlocking of the tows occurs and no stitching yarn is present. Therefore it was required to verify that the reinforcement behaviour could

Fig. 5. Comparison of kinematic drape predictions and measurement for tow placed, hemispherical preforms.

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preform types, the samples using the polyester net followed the estimates more closely and were subject to less scatter. This was attributed to the restraint offered by the continuous polyester filaments which inhibited the tow slippage which is thought to occur at the normal limits of preform compliance. 3.1. “Undrape” modelling Fig. 6. Surface model for hemisphere quadrant showing “optimal” tow paths (uniform tow spacing).

be approximated kinematically. This was tested using a hemisphere since the continuously changing surface and high level of draw produces significant fibre movements and an interesting technical challenge. A surface model of the hemisphere was produced using 400 triangular patches (clearly, symmetry enabled one quadrant only be modelled). Constrained fibre paths were positioned along the quadrant axes and used to generate the result shown in Fig. 4 for a tow spacing of 5 mm. Maximum fibre re-orientation occurs at the extremity of the quadrant where inter-fibre angle is reduced from 90 to 428. To verify this result a series of hybrid carbon/E-glass 0/ 908 flat preforms were manufactured. One set were bound with thermoplastic polyester net and another with polyester powder binder. Alternate tows of carbon and E-glass were laid to provide contrast and facilitate measurement of the included angle between intersecting fibres before and after post-forming. The laydowns were post-formed and moulded in a single operation in a heated (708C) matched aluminum mould. The angles at the tow intersections were determined along the 458 axis by measuring the major and minor axes of the cells formed by four intersecting tows and are referred to a geodesic from the pole to the equator. Fig. 5 shows the inter-tow angles generated by the kinematic drape model together with the experimental data. The results show that the fibre distribution arising from both types of preform is similar to that predicted by the drape model. Of the two

Fig. 7. Lay-flat model with “optimal” tow paths.

Since tow placement permits, potentially, some of the detrimental effects of forming such as fibre buckling and large variations in superficial density to be reduced a further method of producing preforms (method C. above) was developed by modifying the kinematic model to produce “optimised” preforms. This involved the creation of a surface model for the geometry in question to which a series of vectors (representing the required fibre pattern) was applied using CAD. Kinematic draping (undraping in this case) was then used to map the fibres onto a horizontal surface, thus producing an “optimised” geometry. This geometry represented the coordinates of the fibres in two dimensions that, when mapped onto the former, would represent the desired fibre paths in three-dimensional space. A simple example of this is where it is required to produce a preform which, after moulding, would yield a uniform fibre content (in matched die processing) or a uniform thickness (in diaphragm forming). The resulting “optimised” laydown geometry was then processed to provide a CNC part program which was used to produce the flat preform by tow placement. The feasibility of this procedure was first investigated for the small hemisphere described above, again using a hybrid E-glass/carbon arrangement. The purpose here was to reduce the effects of the fibre re-orientation on variations in preform superficial density. A surface model of one quarter of the geometry including tow paths at a uniform perpendicular spacing was developed (Fig. 6). This was intended to ensure uniform superficial density (thus uniform fibre volume fraction) and reduce the possibility of fibre buckling and the severity of property variations within the structure. The surface model was input to the kinematic undraping algorithm which mapped the fibres back to the “optimised” two-dimensional geometry (Fig. 7). It is clear that the necessary pattern is constructed with curved rather than straight fibres to account for the tow displacements during forming. The resulting “optimised” laydown geometry was then converted to a part program which was used to produce the flat preform. While the process is, in principle, aimed at net-shape manufacture the final laydown was produced oversize to assist in retaining the tows in their as-laid positions since short fibres are subject to spring-back when laid on a curved path. Several flat preforms were produced to test the uniform Vf hypothesis in carbon/E-glass and were formed to the hemispherical geometry and moulded as described above. Control samples were also produced using a more conventional 0/908 scheme.

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each quadrant. Again due to symmetry, each quadrant should be similar and so for each hemisphere, only three values were required. The results for eight mouldings are shown in Fig. 10. Each 458 point represents 16 samples and the 0/908 and 22.5/67.58 points each represent 32 samples. The results appear to confirm that the “optimised” fibre architecture produces a more uniform fibre volume fraction around the equator which is the region where such variations would be most extreme using conventional reinforcements. The normalised superficial densities range from 0.98 at the 458 axis to 1.01 at the 0/908 axes. The corresponding variation for the conventional 0/908 reinforcement is from 1.09 to 0.92. This uniformity has potential for improving mould filling characteristics due to a more even permeability and porosity and should also reduce variability in mechanical performance in applications where such nonuniformities cannot be ironed out by staggering the plies. 4. Fibre straightness Fig. 8. Control hemisphere produced using conventional 0/90 (tow placed) lay-up.

Figs. 8 and 9 show finished mouldings produced using a conventional 0/908 reinforcement and the arrangement “optimised” for uniform fibre volume fraction. It can be seen that the conventional reinforcement (Fig. 8) exhibits significant fibre bunching on the 458 axis, suggesting a higher preform superficial density in that region. The optimised structure (Fig. 9) exhibits less bunching in this region. The preforms were impregnated with polyester resin by cold press moulding and sectioned along the 0 and 908 axes to produce quadrants for burn-off specimens. The cutting plan is included in Fig. 10. Due to symmetry, the two outer samples (0 and 908) and the two intermediate samples (22.5 and 67.58) in each quadrant were grouped together thus leaving three samples (0/908, 22.5/67.58 and 458) for

Fig. 9. “Optimised” hemisphere designed for uniform tow spacing.

Out-of-plane waviness is inherent in all woven fabrics due to crimp. In-plane waviness tends to be associated more with the effects of handling or forming operations but was also evident in preforms produced using the tow placement system when several plies were laid in one pass. Waviness reduces the axial properties of an otherwise unidirectional preform. Numerous studies exist [14–22] relating to the effects of waviness on composites manufactured from woven fabric, although waviness is also problematic in other kinds of composites such as in thick filament wound structures [23–30] where buckling of fibres results from the pressure exerted by the over-wrapped layers. 4.1. Tensile modulus Much of the work cited above applies to a relatively short range of waviness and fibre architectures. There are two potential applications for waviness modelling in the present context. The first is to anticipate the degradation of properties due to the (relatively slight) waviness induced due to operating errors and the second is where waviness might be introduced deliberately to, for example, enhance formability on a local basis. Two models were investigated to estimate the effects of waviness on the resulting laminate tensile moduli. Both models assume waviness to be approximated by a sine wave where the waviness factor (a/l ) was measured as the ratio of amplitude (a) of the sine wave to the wavelength (l ). These are described in detail within Appendix A but, briefly, include a simple model based on a modified rule of mixtures and a classical laminate theory derived model. Both models, which account for fibre waviness within a nominally unidirectional composite and predict the effective tensile modulus, were tested for a range of waviness factors and materials using the tow placement facility.

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Fig. 10. Comparison of conventional and “optimised” hemispheres.

It is well understood that the mechanical performance of fibre reinforced composites is highly dependant upon fibre orientation. High strength and stiffness composites can be produced by aligning the fibre axes with the load direction. However, if the fibres are not aligned with the principal loading axis, the composite will have reduced strength and stiffness in that direction. A similar effect will be noticed if the fibres are not straight but include some waviness. Modulus estimates using both the modified rule of mixtures and classical laminate theory were produced for waviness factors from 0 to 1. The model laminates in each case consisted of E-glass/epoxy and carbon/epoxy both at a fibre volume fraction of 30%. The estimates for both materials are shown in Figs. 11 and 12. The results are normalised to 08 laminate

properties. The constituent properties used in the calculations are given in Table 2. 4.2. Specimen manufacture To validate the models, glass and carbon laydowns were produced using the tow placement facility at nominal volume fractions of 30%, although these values actually varied between 25–34% (glass) and 25–30% (carbon). One tow was laid with a predetermined waviness factor under CNC control. The next tow was laid over the top of the first but offset along the longitudinal axis between 5 and 20 mm depending on the wavelength of the specimens. This procedure was repeated until an even coverage and the required fibre volume fraction were achieved. The

Fig. 11. Effects of fibre waviness on normalised laminate modulus for glass/epoxy laminates.

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Fig. 12. Effects of fibre waviness on normalised laminate modulus for carbon/epoxy laminates.

40 mm × 250 mm specimens were processed by RTM in a heated (808C) aluminium tool using Ciba Geigy 5052 epoxy resin. The samples were post cured for 1 h at 1008C in a hot air oven prior to testing. 4.3. Results Following post cure, the wavy specimens were tested in tension using an Instron 1195 screw driven testing machine. Fig. 11 shows the results of the tests for the glass/epoxy laminates with the predicted results from the two models. Six samples were produced for each value of fibre waviness from 0 to 0.4 and the mean values are plotted. Standard deviation of the samples was 5.6% on average and can be accounted for by the variation in fibre volume fraction in each sample set. The Y-axis shows the modulus normalised to the 08 value (30.1 GPa). Reasonable agreement with the estimated values is shown over the range of values tested. The classical laminate model gives an excellent agreement to the results for the first four points (which is the main region of interest, waviness ˆ 0 to 0.05) with deviations from the experimental results of less than 3%. For a/l . 0.1 the classical laminate model over-predicts the effects of waviness although it still follows the results closely. The modified rule of mixtures results follow a similar trend to the previous set but under-estimate the effects of waviness by up to 20%. Fig. 12 shows the results for the carbon/epoxy laminates. Again, six samples were produced for each fibre waviness

factor (from 0 to 0.3). The plotted data are the average of these six samples normalised to the 08 modulus (80.2 GPa). The standard deviations were reduced compared to the glass/epoxy samples at 3.3%. The rule-of-mixtures approach provides a surprisingly good approximation of the experimental data. However, the classical laminate model under-estimates the results over the entire range by up to 14%. Possible reasons for this difference could be that the tow placement head, although programmed to lay the specimens in a sine wave pattern, may induce tension during laydown, resulting in the fibre deviating from the design path. A further explanation can be attributed to the constituent properties used in the predictions which were taken from literature [31,32]. As the classical laminate model is highly dependent on the shear modulus of the unidirectional laminate, more reliable values would obviously improve the accuracy of the predictions. 4.4. Application to woven samples Although the above models predict the effects of waviness on the mechanical performance of composites produced from tow placement produced laydowns with good agreement, the measured waviness induced in laydowns was very small (typically less than 0.5%) which has little effect on laminate stiffness. However similar models can also be applied to other composites with induced waviness such as woven fabrics, although the waviness here is out-of-plane, in contrast to the in-plane waviness

Table 2 Constituent properties for wavy fibre models (from manufacturers data unless otherwise stated)

Epoxy (Ciba Geigy LY5052) E-Glass (PPG 1062) Carbon (Tenax 5131) Glass/epoxy Carbon/epoxy a

E1 (GPa)

E2 (GPa)

G12 (GPa)

n 12

3.15 84 260 27.4 [R] a 82.4 [R] a

3.15 84 20 6.62 from [32] 7.0 from [32]

0.98 32.8 from [32] 104 from [32] 1.74 from [32] 5.3 from [31]

0.38 0.22 0.2 0.33 0.33

[R] ˆ Estimation using the rule of mixtures.

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waviness and standard deviation for the glass/epoxy laminates was calculated to be 0.033 ^ 15% and 0.03 ^ 56% for the specimens made using carbon fabric. Predictions of the effects of the measured waviness on the tensile modulus of the samples were developed from the modified rule of mixtures and the out-of-plane classical laminate models. Constituent properties were derived from manufacturers data or from rule of mixtures and Halpin–Tsai estimations and are listed in Table 3. Validation of the models for the woven materials was performed by testing the samples described above using an Instron 1195 screw driven test machine. Tensile tests were performed on all samples and the results are shown in Table 4 which represents the average for six specimens in each case. Fig. 13. Micrograph of 600 gsm plain weave laminate used to establish waviness parameters.

described previously. The modified rule of mixtures model above should be equally applicable and thus is unchanged. However, the classical laminate model is modified slightly to account for the change in architecture. The main difference between the two classical laminate models is the change in boundary conditions due to a shear coupling associated with in-plane waviness. The full derivation of the model is provided in Appendix A. To verify the two models for woven fabrics, test specimens were produced from woven E-glass and carbon fabrics. The glass fibre samples were produced from Flemings W/R 600 plain weave fabric and Ciba Geigy 5052 epoxy resin at a fibre volume fraction of 57% (^2.8%). Six carbon/epoxy samples were manufactured from Hexcel CGG108 five harness satin weave and the same epoxy resin at a fibre volume fraction of 46% (^2.5%). Both types of specimen were 40 mm wide and 250 mm long and were produced by RTM in a heated (808C) aluminium tool and post cured as before. Waviness was determined by preparing a 20 mm × 20 mm specimen taken normal to the fabric plane. The sample was ground with successively fine abrasive paper and then polished using 3 mm alumina paste. The samples were scanned at 10 × magnification (Fig. 13) using a closed circuit digital camera mounted on an optical microscope. Fibre waviness was determined by image analysis. Waviness was measured for ten fibres chosen randomly within the samples. Average

4.5. Glass/epoxy results Using the experimental results, manufacturers data and the well known Halpin–Tsai equations for ply properties as shown in Table 3, the predicted modulus for a 0/908 E-glass/ epoxy composite at 57% fibre volume fraction using the rule of mixtures is 29.8 GPa. Using a modified rule of mixtures approach with an included fibre waviness ratio of 0.033 to represent the crimp, the predicted modulus reduced by 10% to 26.9 GPa. Classical laminate theory predicted a similar reduction of 12% to 26.3 GPa. The experimental moduli for the average of the six specimens was 27.5 GPa (a reduction of 8% on the rule of mixtures value). The standard deviation was 1.1 GPa. Thus both models provided a useful estimate (within 4%) of the effects of fibre waviness on modulus for these fabrics. Although the experimental results were as high as 28.7 GPa (a 3.7% decrease on the rule of mixtures value) and as low as 26.5 GPa (11% decrease), this can be attributed mainly to the variation in fibre volume fraction (55–59%) in the test laminates. 4.6. Carbon/epoxy results Using the constituent properties in Table 3, the estimated modulus for a 0/908 carbon/epoxy composite at 46% fibre volume fraction was 62.4 GPa (rule of mixtures). Classical laminate theory with a waviness factor of 0.03, predicted a reduction in modulus of 13% to 54 GPa. A reduction in modulus of 8% to 57 GPa was predicted by the modified rule of mixtures. The mean experimental value was 53 GPa

Table 3 Constituent properties for woven fabric models

Epoxy (Ciba Geigy LY5052) E-Glass (Flemings W/R 600) Carbon (Hexcel CGG108) Glass/Epoxy Carbon/epoxy a

E1 (GPa)

E2 (Gpa)

G12 (GPa)

n 12

3.15 75 230 29.8 [R] a 62.4 [R] a

3.15 75 20 29.8 [R] a 62.4 [R] a

0.98 29.25 from [32] 100 from [32] 3.2 from [32] 5.3 from [31]

0.38 0.22 0.2 0.33 0.34

[R] ˆ Estimated using the rule of mixtures.

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Table 4 Comparison of predictions with experimental results

E1 Ex Ex Ex

(Rule of mixtures) (Modified rule of mixtures) (Classical laminate) (Experimental)

Glass/Epoxy (GPa) [S.D.]

Carbon/Epoxy (GPa) [S.D.]

29.8 (100%) 26.9 (90%) 26.3 (88%) 27.5 (92%) [1.08]

62.4 (100%) 57.2 (92%) 54.1 (87%) 53.4 (86%) [3.86]

(a reduction of 14% on the rule of mixtures predicted value). The classical laminate theory and modified rule of mixtures models both under-estimated the severity of the effects of fibre waviness by 1 and 6%, respectively. The modified rule of mixtures model is less accurate because it is a very simple geometry based model and does not take account of laminate properties such as shear modulus and Poisson’s ratio. Other potential sources of error include the apparently large deviations (up to 56%) from the mean measured waviness factor (0.03) or departures from the idealised sinusoidal fibre architecture. 4.7. Permeability Fibre waviness influences processing properties, notably the in-plane permeability which characterises the ease of impregnation for conventional in-plane flow, in addition to the mechanical performance. Permeabilities are important in practice for screening candidate reinforcement materials and are useful in process simulations to assist mould design. The derivation of an expression for the axial permeability of a laminate containing sinusoidal fibres is given in

Appendix B. This relationship was applied using numerical integration to estimate the in-plane permeabilities for nominally unidirectional preforms with induced fibre waviness.

4.8. Specimen manufacture To test the above model, a set of nominally unidirectional glass fibre preforms with induced fibre waviness were manufactured on the tow placement facility. The preforms were produced at waviness factors from 0 to 0.1. This represents a reduced range compared to the mechanical property tests but was imposed by the need to produce wider specimens to suit the permeability test facility. The laydowns were produced by offsetting successive tows along the Yaxis (mechanical property specimens were offset along the X-axis) and the arising fibre tensions resulted in disturbance of the as-laid architecture at smaller radii of curvature, thus limiting the degree of waviness to 0.1. All preforms were produced to a fibre volume fraction of 43% (^1%) by altering the spacing between consecutive fibres.

Fig. 14. Effects of fibre waviness on in-plane permeabilities.

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Table 5 Tensile properties of marine transom flange from fabric and tow placement Reinforcement

Fibre volume fraction (%)

Tensile modulus (GPa)

Standard deviation (GPa)

UTS (MPa)

Standard deviation (MPa)

Quasi-isotropic fabric 1 random mat Tow placement

33.09

11.85

0.68

162

11.4

32.58

13.33

0.74

171

16.9

4.9. Results

5. Technology demonstrators

The preforms were tested for permeability using an existing bench-top permeability rig that has been used successfully to determine the in-plane permeability of various preforms using pressure monitoring at several positions within a circular cavity. Preform stacks are loaded into the cavity and SAE30 mineral oil is injected at a constant flow rate (1.67 ml/s in this case). Pressure and temperature histories are recorded via a PC-based data acquisition system and are downloaded into a spreadsheet that is used to calculate permeability using Darcy’s law. The detailed experimental stages and analytical procedures are described in detail elsewhere [33]. Up to six preforms were tested at each value of waviness. The results are shown in Fig. 14 and include the estimated values based on the principal permeabilities of a 08 fibre bed at 43%Vf. The experimental data represent four fibre waviness factors. Both the averaged longitudinal and transverse permeabilities follow the trend of the estimated values although the experimental results generally show greater isotropy. Also, the deviation of the experimental values in each waviness sample was quite large ranging from 1.3 to 53%. This was probably an effect of the variations in preform superficial density within each sample, since permeability is well known to be a strong function of porosity. The apparent differences in isotropy between the predicted and experimental values may be attributable to small errors in alignment of the preforms with respect to the principal axes of the mould cavity.

Four demonstrator parts were studied in order to provide a representative range of industrial problems. 5.1. Marine transom flange The marine transom flange was a load-bearing structure used in the manufacture by RTM of rigid inflatable boats. The flange supported an inflatable tube running around the perimeter of the boat. The flange preform was produced conventionally from two skins of random reinforcement (Vetrotex Unifilo U816) with a quadri-axial stitch bonded core (Tech Textiles EQX 2336). For comparison purposes, flat plaque mouldings were produced from conventional fabrics and using the tow placement facility before moulding as described above, post-curing and mechanical testing. Table 5 shows the results of tensile tests (twelve for each composite). It can be seen that the strength and modulus of the tow placement laminates provide an adequate match for the fabric equivalents (with small improvements overall) despite the earlier concerns over fibre waviness. One factor here may be the reduced fraction of non-structural stitching yarn which is generally present at up to 5% by mass in fabrics produced by warp-knitting and related techniques. This was not taken into account in the fibre fraction calculation and since the presence of non-structural binder yarns in the tow placed preforms was much lower (c. 1%), the effective reinforcement fraction is likely to have been significantly higher. 5.2. Generic power bulge

Fig. 15. Generic power bulge produced using conventional 0/90 (tow placed) lay-up.

The power bulge involved complex curvature and a shallow draw and was used to validate the “undrape” software described earlier. A drape analysis was performed on a surface model of the bulge and showed high levels of shear at each corner. This local deformation represented an increase in superficial density of up to 43% which would otherwise cause substantial local property changes and some processing difficulties. As a solution to the problem of fibre bunching, a surface model was produced with uniform fibre spacing. This model was used as an input to the “undrape” software described earlier to produce an “optimised” laydown which when formed would produce a preform with reasonably uniform fibre volume fraction throughout. Several such laydowns were produced using the tow placement facility. The preforms were moulded in

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plan in Fig. 17. The results show the normalised superficial density for each burn-off specimen. The normalised results from the “optimised” structures range from 0.90 to 1.08. The corresponding values for the conventional fabric range from 0.86 to 1.18 confirming that the “optimised” fibre architecture produces a more uniform fibre volume fraction around the edge of the component which is the region where such variations would be most extreme. 5.3. General aviation propeller blade

Fig. 16. “Optimised” power bulge designed for uniform tow spacing.

a matched GRP tool using unsaturated polyester resin. For comparison, several conventionally oriented (0/908) laydowns were also produced on the facility and moulded. Figs. 15 and 16 show the mouldings manufactured from the conventional style material and the “optimised” laydown. Carbon tow was added to the preforms to increase contrast and ease measurements. Superficial inspection suggests that the conventional preforms exhibit greater deformation at the corners. In order to quantify the uniformity of the preforms, an Eglass laydown was produced for each fibre architecture. Each laydown was preformed and moulded. Burn-off specimens were cut from the base of the mouldings to the cutting

The blade is constructed conventionally from 90 layers of tailored fabric in orientations of 08, ^ 458 with a randomly oriented glass fibre core. The tow placement facility allowed a modified layup plan to be used. 2400 tex glass fibre was used to manufacture a 30-layer preform whilst maintaining the original fibre architecture specifications. The preforms were made using the multi-tow roller placement device. Fig. 18 compares the time taken to produce each layer of the preform with data for a similar preform from a previous study by McGeehin [10]. Each ply took longer to lay than during the previous study due to the cutting operations which must be performed whilst the head is stationary. The use of a sub-routine within the control system, allowing cutting to be performed whilst the machine is moving would obviously overcome this problem. Although the placement times for each layer were increased, the total preparation time for the preform was reduced by 14% (22 min) due to

Fig. 17. Superficial density variations for conventional (lower) and “optimised” (upper) power bulge mouldings.

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Fig. 19. Compressor outlet guide vane in carbon/epoxy.

5.4. Compressor outlet guide vane (OGV) The conventional OGV (Fig. 19) was produced from 21 laminates, each comprising five sub-laminates of 6k carbon fibre. The tow placement facility reduced the number of plies required from 105 to 63 by using 12k carbon tow and adjusting fibre spacing to provide the correct proportion of longitudinal and transverse reinforcement. Two complete carbon fibre OGV preforms (Fig. 20) were manufactured by tow placement and moulded in a matched die, heated steel tool and subject to proprietary vibration tests for comparison to a carbon/epoxy OGV manufactured from conventional stitch bonded carbon fabric. Microscopic examination of the RTM vanes indicated that the porosity associated with the compression moulded parts was reduced. However, vibration testing showed that the part with the towplaced preform reduced the “first bow” frequency by 10% compared with the compression moulded version. This was attributed largely to a lower overall fibre fraction and to the possibility of increased fibre waviness, a lower fibre fraction on the leading edge and the absence of an erosion shield. Despite these limitations, the tow-placed

Fig. 18. Preform preparation times for 30-layer propeller blade preform using single tow (upper) and multi-tow (lower) placement devices.

the on-line consolidation associated with roller placement which eliminates the need for separate processing operations. The total preparation time of 2 h and 13 min was a 32 min (20%) reduction over the time taken to produce a preform from tailored fabric. Also material wastage was reduced from 15% to less than 2%.

Fig. 20. 0/90 preform element for compressor outlet guide vane.

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economic case for tow placement would seem to be attractive.

Table 6 Relative raw materials prices [34] Description

Carbon tow 3k Carbon tow 12k Carbon fabric (5 harness weave) 290 g/m 2 Equivalent carbon fabric pre-preg Admesh polyester netting LY 5052 resin RTM6 resin

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Relative cost (per kg unless stated otherwise) 1 0.5 2.16 US$60/m 2 US$3/m 2 0.35 1.12

part out-performed a stitched fabric preform which suffered 18% reduction in the first bow frequency. Measurement of the torsional vibration frequencies produced similar results for all of the vanes tested, the tow-placed part having marginally higher values (2%) than the two alternatives.

6. Economic evaluation An economic comparison of tow placement with conventional preform manufacture and pre-preg lay-up for a typical carbon/epoxy composite was done to compare the material costs for carbon/epoxy pre-pregs with similar materials used in RTM including carbon fabric and carbon tow. Table 6 compares the cost for individual materials, while Table 7 describes the material cost to manufacture representative composites. Table 7 shows that the pre-pregging operation invokes a cost penalty of around 20% compared with the equivalent (conventional) RTM materials. A further 20% saving can then be made if tow placement can be substituted for high cost, aerospace fabrics. On this basis then, an RTM composite relying on tow placement can be produced at between 28 and 60% of the cost of a pre-preg, depending upon the specification of the tow and epoxy resin. It should also be pointed out that these costs are based, conservatively, on 100% materials utilisation. Given likely industrial figures of 70, 80 and 95% for pre-preg, RTM using fabrics and RTM using tow placed preforms respectively, the Table 7 Normalised costs of pre-preg, fabric and tow-placed composites [34] (assuming 100% materials utilisation) Description

Relative cost/kg @50% Vf

Carbon/RTM6 pre-preg Fabric/RTM6 Fabric/LY5052 3k Tow/RTM6 3k Tow/LY5052 12k Tow/RTM6 12k Tow/LY5052

1 0.82 0.69 0.59 0.47 0.41 0.28

7. Discussion and conclusions The results presented have demonstrated the technical feasibility of producing flat preforms for subsequent postforming using tow placement, assisted by a roller device and a polymer-based tackifier for tow retention. One of the main practical difficulties involves maintaining fibre straightness since the tension which can be applied to the tow is limited by the adhesion of the tackifier. Clearly, a mechanical holding device would improve the situation (as would several additional hardware refinements) although this would add to both the capital costs and the complexity of the operation. Nevertheless, reductions in preform preparation times and material waste were demonstrated compared to conventional fabric-based preforming for representative industrial parts. Simple materials costs comparisons suggest reductions of 20 and 40% compared with fabrics and pre-pregs, respectively which appear attractive to commercial operators. The laminates manufactured here using (roller assisted) tow placement approached the static performance achieved using conventional fabric reinforcements. Some waviness was evident and the effects on both elastic properties and in-plane permeabilities were anticipated using a simple geometric model. In neither case were the properties downgraded significantly compared to equivalent fabric-based laminates. Comparison of the fibre distribution in a hemispherical preform produced by tow placement showed that the forming process could be approximated using kinematic draping routines. This enabled a new mapping method to be used to design a two-dimensional tow pattern which would form to produce an approximately uniform superficial density in three dimensions. This approach may have useful applications in the future for components of complex curvature.

Acknowledgements This work was carried out under the DTI/EPSRC LINK Structural Composites Programme. The authors wish to thank the following project partners for their collaboration: Ciba Polymers, Hexcel Composites, Dowty Aerospace, Island Plastics, MSC, Newall Aerospace, PPG, Rolls Royce plc. The major contribution of Geoff Tomlinson to the development, maintenance and running of the CNC manipulator cannot be over-emphasised. Appendix A. Axial modulus of a laminate with sinusoidal fibre distribution A.1. Modified rule of mixtures model The simplest model uses the load bearing efficiency of the

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reinforcement based on the well known orientation efficiency factor within the modified rule of mixtures. E1 ˆ hEf Vf 1 Em …1 2 Vf †:

…A1†

The efficiency factor is based on the orientation of the fibres with respect to the direction of loading. The efficiency of a fibre path following a sine wave is described by: 1 Z 2p hˆ cos3 u dx …A2† S 0 where the length of the fibre over one cycle is  Z2p 21 21 cos u dx Sˆ 0

…A3†

and the angle subtended by the fibre with the longitudinal axis is     2p 2px u ˆ arctan acos : …A4† l l The laminate efficiency factor for a sinusoidal fibre distribution was calculated by numerical integration of Eqs. (A2) and (A3) using the trapezium rule. A.2. Classical laminate model for in-plane waviness A second model was applied, using the same fibre geometries, based on classical laminate theory [25]. This includes the effects of shear modulus and Poisson’s ratio in the calculation. Here the laminate modulus for a 08 composite with off-axis loading at an angle, u , to the fibres is given by 1 S2 ˆ S11 2 16 Ex S66

…A5†

where S11 ; S16 ; S66 are components of the modified compliance matrix for off-axis loading. For a sinusoidal fibre, u varies along the wavelength of the fibre. At any point along the wavelength, u is derived from Eq. (A4). The modulus was derived by integrating Eq. (A5) over one wavelength using numerical integration:   Z2p S11 S66 2 S216 dx: …A6† Ex ˆ S66 0 A.3. Classical laminate model for out-of-plane waviness The classical laminate model can also be applied for outof-plane waviness as found in woven fabrics. Here the laminate modulus for a 08 composite with off-axis loading at an angle, u , to the fibres is given by 1 ˆ S11 : Ex

(A7) over one wavelength using numerical integration similar to the method described above.

…A7†

Again for a sinusoidal fibre, u varies along the wavelength of the fibre. At any point along the wavelength, u is derived from Eq. (A4). The modulus was derived by integrating Eq.

Appendix B. Axial permeability of a laminate with sinusoidal fibre distribution Permeability is defined by Darcy’s law: Q ˆ 2K

A DP : m L

…B1†

For a unidirectional reinforcement, the principal permeabilities are described by K1 and K2. A transformation of the matrix of principal permeabilities gives the permeability at an angle, u to the fibre axis: Kx ˆ K1 cos2 u 1 K2 sin2 u 2

…K2 2 K1 †2 sin2 ucos2 u : K1 sin2 u 1 K2 cos2 u

…B2†

For a laminate containing sinusoidal fibres, the permeability in the longitudinal fibre direction is obtained from the following equation: Zp dx p ˆ …B3† Kavg 0 Kx where Eq. (B2) is used to calculate Kx using the fibre orientation given by Eq. (A4). The estimated permeability of nominally unidirectional preforms with induced fibre waviness was obtained from Eq. (B3) using numerical integration.

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