40207_26

26

CONTROL

Fig. 26.1 Schematic of the prepreg lay-up used inautoclave cure (Springer, 1986).

lay-up used in autoclave cure. In a traditionallay-up process, prepregs with different fiberorientation and architecture are placed in cer-tain order forming a near-net-shape compositestructure. During the lay-up operation,whether it is done manually or using a robot,the trapping of air pockets within the structureis unavoidable. Thus a consolidation step afterthe lay-up operation is necessary. Prepregs areusually provided with relatively low fiber vol-ume fraction. With the consolidation step, thefiber volume fraction of the composite productcan be increased and excessive resin can beremoved.

The basic mechanisms involved in a consol-idation process are the fiber deformation andresin flow, which are coupled with thermaleffects and the resin cure reaction. A similarconsolidation process can also be seen in soilmechanics such as the settlement of a founda-tion. However, the deformation behavior offibrous materials is substantially differentfrom that of granular structures and resin flowbehavior is strongly affected by thermal effects

26.1 INTRODUCTION

Consolidation is an important step that occursin almost every process used to make anadvanced composite article. Consolidation isusually brought about by the application ofpressure at a boundary which squeezes air andresin out of the composite thereby changingboth its microstructure and dimensions.Improper consolidation can lead to voids,residual stresses, warping and other unwantedeffects which could ultimately lead to the rejec-tion of the part. A comprehensive discussion ofconsolidation in composites would includemany complex phenomena. Simultaneously,there is heat, momentum and mass transfer,accompanied by the chemical curing reactionof the resin and the deformation and motion offibers.

Consolidation techniques have been used inthe fabrication of both thermoset and thermo-plastic composite parts, but are more crucialsteps in thermoset composite processing. Thetraditional composite manufacturing processfor aerospace industry products usually startswith the B-stage impregnated prepregs con-sisting of fiber preforms and staged resinmatrix. Usually the resin content is relativelyhigh. In order to achieve the required compos-ite material properties which are dominatedby the fibers, consolidation is used as an

important processing step.Figure 26.1 shows a setup of the prepreg

Handbook of Composites. Edited by S. T. Peters. Publishedin 1998 by O1apman & Hall, London. ISBN 0412 540207

Introduction 577

and chemical reaction of the resin. Thus, the study of the consolidation process of fibrous composite materials involves many disciplines.

To effectively control a consolidation process, the selection of the equipment and tooling materials is crucial. Major process parameters for a consolidation process include pressure and temperature and both are functions of time and are usually set as operation cycles. Thus the system setup should be able to effectively control the pressure and temperature profile and transfer heat and pressure to composite parts. Figure 26.2 illustrates the process variables applied during autoclave consolidation and cure.

TEMPER ATU RE PRESSURE

t t i t t Fig. 26.2 Illustration of the process variables (temperature, pressure) applied during autoclave consolidation and cure (Springer, 1986).

One of the commonly used facilities is an autoclave, which is a closed pressure vessel with means for heating and applying pressure and vacuum to its contents. The dimensions of the composite parts are limited by the size of autoclaves. Thus, for large size composite structures, alternative processing techniques have been used, such as vacuum bag molding.

In addition to the equipment, tooling material has direct influence on the composite part surface quality, dimensional accuracy and residual stress. The main considerations for tooling material include strength, stiffness, thermal expansion coefficient, hardness and

surface roughness. Metals are widely used as tooling materials for composite processing. However, their heavy weight and high cost of machining become disadvantages when complex geometry is involved. Composite tooling materials have been used as alternatives in various consolidation processes.

Another tooling component for the consolidation process is the bleeder, which is usually a nonstructural layer of porous cloth or paper which allows the escape or bleed out of excessive gas and resin during the consolidation process. Sometimes the process is called migration. The bleeder cloth or paper is removed after the curing process and is not part of the final composite.

Breather material is used to provide a vacuum path over the surface of the part. Typical materials are glass and mat. They can be stretched over the part contours to ensure an effective vacuum path and sometime also to provide a cushion effect to matched metal tools.

Bagging and sealing are crucial to the quality of the composite parts. General requirements for the bag are: (1) the bag must apply curing pressure uniformly; (2) the bag must not leak under molding conditions; and (3) a good vacuum path must be provided in bagging. Silicone rubber vacuum bags are widely used because of their long service life. Moreover, they are repairable and self-healing with respect to pinholes. The initial cost of fabrication is relatively higher. Nylon is an alternative bag material for up to 193°C (380°F) and is usually discarded after use.

The commonly used form of resin matrix prepreg has a resin content beyond 40% and requires a significant amount of resin bleedout during cure to achieve a cured laminate resin content of 28-32%0. Low resin content prepregs have been developed which can be used without resin bleedout processes. Since there is no bleedout process, less resin and less bleeder material are needed for a consolidation and cure process. However, the removal of entrapped air becomes a more critical aspect of process control.

578 Consolidation techniques and cure control

26.2 CONSOLIDATION MODELS

As composite applications were expanded rapidly in the late 1970s and early 1980s, studies on the process science of composite materials became very active, especially in the areas of consolidation and cure (Lindt, 1982, 1986; Springer, 1982,1986; Loos and Springer, 1983a,b; Halpin, Kardos and Dudukovic, 1983; Loos and Freeman, 1985; Gutowski et al., 1987a,b; Gutowski and Cai, 1988; Dave, Kardos and Dudukovic, 1987a,b; Tang, Lee and Springer, 1987; Batch and Macosko, 1988; Kim et al., 1988 and 1989; Connor et al., 1993). The purpose of these studies has been to find out the most suitable process parameters and then through scientific process modeling to achieve the optimized composite product quality.

Pioneering work in composite consolidation and cure modeling was led by Springer (Springer, 1982; Loos and Springer, 1983a). Laminated composite structure was considered with bleeder layers placed on top of the composite laminates. When pressure was applied transversely to the laminate plate, excessive resin material which was in a fluid state was squeezed out from the laminate.

Similar consolidation models have also been developed by Gutowski’s group and Kardos’ group (Gutowski, Morigaki and Cai 1987a; Gutowski et al., 1987b; Dave, Kardos and Dudukovic, 1987a,b). In these models, both the fiber material deformation which is highly nonlinear and the outgoing flow of the resin are considered.

As can be seen from the discussion presented later, both approaches are valid within the ranges of parameters considered. Experimental verification results show good agreement with these model predictions (Gutowski et al., 1987b; Kim et al., 1988; Cai and Gutowski, 1989). When resin content is relatively high and fiber-to-fiber contact is insignificant, Springer’s model can be applied. On the other hand, if fiber volume fraction is relatively high and fibers carry a substantial

portion of the load, then Gutowski’s and Kardos’ models are applicable.

In the following discussions, both Springer’s model and Gutowski’s model will be presented. Kardos’ model is equivalent to Gutowski’s model but different process variables are used in the modeling.

26.2.1 RESIN FLOW

The problem of resin flow in composite processing can be treated as flow through fibrous porous media. In general it can be handled by Darcy’s law which states that the flow rate is proportional to the pressure gradient applied and is related to the porous medium permeability and fluid viscosity. The general form of Darcy’s law in a one-dimensional case is:

(26.1)

where q is the average resin flow rate, K is the preform permeability with the units of length squared, p is the resin viscosity and dp/dx is the imposed pressure gradient. In the case of resin flow into the bleeder, the consolidation pressure is established between the advancing flow front and the tool or mold surface.

The main issue involved in using Darcy’s law in the consolidation process is that neither the fiber preform permeability nor the resin viscosity is constant over the process. The preform permeability is a function of the porosity or fiber volume fraction, fiber diameter, fiber orientation and fiber architecture. Among them the fiber volume fraction changes substantially during a consolidation process. Resin viscosity is related to temperature, the cure status and cure time and changes dramatically in the process. Usually at the start of a cure process, resin is in the semi-solid state. With the rise of temperature, it becomes fluid. As the degree of cure increases, it gels and becomes solid.

The permeability of fibrous preforms has been studied both analytically and experimentally (Williams, Morris and Ennis, 1974; Gutowski et al., 1987b; Lam and Kardos, 1988,

Consolidation models 579

1989; Van Den Brekel and De Long, 1989) using the well-known Kozeny-Carman equation. The estimation formula using the fiber structural variables can be written as:

r; (1 - VJ3

4k0 v; K = -~ (26.2)

where rf is the fiber radius, V, is the fiber volume fraction so that (1 - VJ is the porosity and k, is an empirical constant, called the Kozeny constant, which is usually determined experimentally. For different textile architecture and orientation, the value of k, will be different. Reported experimental data show that for an aligned fiber bundle, k,, = 0.5-0.7 for the longi- tudinal flow and k, = 11.0 for the transverse flow. For f 45" cross plies, k, = 2.70. For woven type textile preforms, ko = 5.5. It should be pointed out that many experimental results have been reported and the variation of the Kozeny constant in some cases is significant. Also in the transverse direction, a modified Kozeny-Carman equation has been proposed to account for the stop-flow phenomenon when fiber volume fraction reaches the maximum packing efficiency (Gutowski ef al., 198%). Figure 26.3 shows a comparison of measured axial permeability values for aligned fibers with the Kozeny-Carman equation.

Resin viscosity can be expressed as an empirical function of temperature and degree of cure (Lee, Loos and Springer, 1982). The expression can be written as:

p = p-exp (U/RT+Ka) (26.3)

where p- is a constant, U is the activation energy for viscosity, a is the degree of cure and K is a constant which is independent of temperature.

Experimental study has been performed for the Hercules 3501-6 epoxy resin which is widely used in composite fabrication. Figure 26.4 shows the viscosity measurement as functions of temperature and time. To match the model predictions and experimental data, the constant K is found by fitting a linear least square curve to the p versus a data generated

0 20 ply somple (0.002 in/min.carn oill x 20 ply sample (0.005 in/rnin.com oill

Carman- Kozeny Eq, kxx=0.7

x 5

Fiber Volume Fraction ( V f )

Fig. 26.3 Comparison of measured axial permeability values for aligned AS-4 fibers with Carman-Kozeny equation (Gutowski et al., 198%).

Fig. 26.4 Measured viscosity of 3501-6 resin as a function of time (Lee et al., 1982).

at a constant temperature. Thus the value of K

is found to be 14.1 * 1.2. The values of p, and U are found to bep- = 7.93 x Pa s, U = 9.08 x 104 J mol-'.

The degree of cure a and the rate of degree of cure da/dt were determined from the results of 'isothermal scanning experiments.


Efforts were made to describe the daldt versus a data with a modified Arrhenius type equation. The proposed empirical equations are

daldt = ( K , + K,a) (1 -a ) ( B - a) (26.4) a 10.3

da/dt = K3 (1 -a) a > 0.3

(26.5)

When the permeability and the resin viscosity are known, with the imposed applied pressure condition, the rate of the outgoing resin flow can be calculated using the Darcy equation. In general, flow may be multi-direc- tional. Thus 2-D or 3-D flow equations have to be solved. In practice, resin flow in one particular direction may be dominant, and the analysis can be handled as 1-D permeable flow.

26.2.2 FIBER DEFORMATION

The main contribution from Gutowski's model is the description of fiber deformation behavior. Instead of treating fibers as separate layers, a network concept is introduced. In other words, fiber-to-fiber contact is assumed within

where

K , = A, exp (-AE,/RT)

K2 = A, exp (-AE,/RT)

K, = A, exp (-AE,/RT) a fiber assembly, even in the case of aligned fiber bundles. Thus a fiber filament span between the neighboring contact points becomes a small bending beam. During a consolidation process when fibers are pushed closer, more and more fiber-to-fiber contacts

A,, A, and A, are the pre-exponential factors, AE,, AE, and AE, are the activation energies, R is the universal gas constant, and T is the absolute temperature. The constants in the expression are found as:

B = 0.47

A, = 2.101 x lo9 min-'

A, = -2 .014~lO~rnin-~

A, = 1.960 x lo5 min-l

AE, = 8.07 x lo4 J mol-'

AE, = 7.78 x 104 J mol-'

AE, = 5.66 x lo4 J mol-'

As can be seen from the discussion, all the constants involved in the model are determined experimentally through a specified process. Similar treatment can be used for other types of resin systems, and experimental investigation results have been reported, including Hercules HBRF-55 Resin (Bhi et aI., 1987) and Fiberite 976 Resin (Dusi et al., 1987). A similar process model has also been discussed by Roylance (1988).

take place, and the span length reduces. Thus the bending stiffness of these small fiber beams increases rapidly, resulting in nonlinear elastic deformation response. The nonlinear elastic response of a fiber assembly under a compressive load has been also studied in the textile field, and an empirical formula was proposed (van Wyk, 1946).

A proposed fiber deformation model for aligned fiber bundles considers the deformation status variable, the fiber volume fraction Vf , as a function of the consolidation pressure (Gutowski, 1985). The expression is

where V, is the maximum obtainable fiber volume fraction for a given fiber network configuration, and V , is the fiber volume fraction below which the fiber network carries no load. The empirical constant As is obtained from curve fitting on available measurement data. A typical fiber deformation curve for


well aligned graphite fibers is shown in Fig. 26.2.3 CONSOLIDATION MODELS

As discussed above, in Springer’s model, it is assumed that there is no fiber-to-fiber contact.

26.5 with the- co-mparison of measured data points.

In ,Q 700 b

? 500 Data Point-

In v)

al -1 300 e

.c : 200

100 n - 0.4 0.5 0.6 0.7 0.8 0.9 Fiber Volume Fraction ( V f )

Fig. 26.5 Typical fiber deformation curve for well- aligned XA-S and AS4 graphite fibers (Gutowski et al., 198%).

The proposed relationship between the compressive fiber stress of and fiber volume fraction V, provides a tool to estimate the finished consolidation status of the composite products. If the time window for the consolidation is long enough, and excessive resin is completely squeezed out from the structure, the consolidation pressure is then balanced by the fiber stress. However, because of the dramatic change of the resin viscosity and preform permeability during a consolidation process, resin flow may not be complete. Thus, developed consolidation simulation models are needed for the process analysis and improvement.

During the compression of fibrous preforms, structural relaxation has been observed (Gutowski, 198%). Thus the deformation to some extent is not elastic but viscoelastic. This issue has been addressed by using a Maxwell type model (Kim, McCarthy and Fanucci, 1991).

Thus a dynamic fluid pressure exists between the consolidated layers. The consolidation time, which is crucial to the cure process, is related to the permeability of the fibrous preforms, resin viscosity, and the applied consolidation pressure. In Gutowski’s model, the fiber reinforcement and the fluid state resin are considered as a system. Both fiber network deformation and fluid resin flow are solved together. Both models are presented here with a laminated composite structure as the example.

The example for Springer’s model is the laminate consolidation with flow in the laminate transverse direction, or z direction. A bleeder ply is assumed to be placed on top of the composite. Figure 26.6 shows the setup for the model. At any instant of time the liquid velocities in the bleeder Vb and in the composite Vc are given by Darcy’s law. For a constant viscosity liquid, the integrated forms are:

(26.7) K c (Po - P,)

h C 9, = 7

where p , and pb are the pressures at the composite-bleeder interface and in the bleeder respectively, po is the consolidation pressure and is related to the applied force or pressure,

is the instantaneous thickness of the liquid in the bleeder, and hc is the thickness of the resin starved layer, or the thickness of the layers through which resin flow takes place, and K, and Kb are the permeability of the composite layer and bleeder respectively. If the compacted composite layer thickness is h,, then

hc = nh, (26.9)

where n is the number of layers or plies already compacted.


L Resin Flow

Fig. 26.6 Illustration of the consolidation model proposed by Springer (1982).

The equation of continuity gives the rate of change of volume of the composite as:

-~ d(hA) = Aq, = Aq, (26.10) dt

where A is the surface area of the composite laminate, and h is the total thickness of the composite laminate. The second equation expresses the fact that at any instant of time, the flow out of the composite is equal to the flow into the bleeder. The pressure po is related to the applied force as:

f = + pa

(26.11)

where F is the applied force and pa is the atmospheric pressure. By combining these equations, the consolidation equation becomes:

Therefore for each individual layer, the consolidation time can be calculated. The total consolidation time is the summation for these

layers. The final status of the composite is dependent on the compaction of each individual layer.

As a comparison, Gutowski’s consolidation model combines the flow of resin through porous media and the fiber deformation behavior. Similar treatment has been presented in studies of other fields including soil mechanics (Biot, 1941,1955,1956; Gibson and Hussey, 1967). In general, consolidation occurs in only one direction, but flow may take place in all three directions. Thus an element is deformable in the z direction. A new variable 6 is used to represent the deformation, and 6 = z + w where w is the local displacement of the fiber network. The laminate setup for the model is illustrated in Fig. 26.7. If the initial fiber volume fraction for the composite is V, and the fiber volume fraction at any instant is Vf, the fiber continuity condition states

vo=-v, 36 a Z

(26.13)

Resin flow continuity condition requires:


(26.14)

With the application of Darcy's law, a consolidation equation using the fluid pressure p , and fiber volume fraction Vf as variables can be written as

This equation gives a relationship between the spatial and time-varying nature of the pressure in the resin and the fiber volume fraction of the composite. The equilibrium statement

Here it is assumed that the inertial effects in the process are small. Therefore the applied pressure is balanced by a combination of the average resin pressure and the fiber stress. In other words, any load which is carried by the fibers is then unavailable for pressurizing the resin.

Since both the permeability and the fiber stress are expressed as functions of fiber volume fraction V , with the given initial and boundary conditions, the variables V, and p , as a function of time and location can be solved. In general numerical calculation procedures have to be developed for solving the partial differential equations. In some simplified cases, analytical solutions are possible.

Example problem 1: One-dimensional flow in compression molding

for the consolidation is: A simplified example of composite consolidation is the compression molding of a flat rectangular laminate. The composite part is pressed between two solid dies. Therefore only in-plane flow is possible. In other words, flow components are in the x and y directions only. If the initial fiber volume fraction is uniform, the equation of the resin flow and fiber deformation becomes:

J2Pr + K - a2pr + -~ P av, = 0 (26.17)

(26.16) A

K X ~ y ay2 v, at ho Here it is also assumed that there is no signifi-

cant pressure gradient in the z direction, and the viscosity p does not vary spatially.

In some cases, K Z / a 2 >> Ky/b2 where a and b are the dimensions of the laminate in x and y directions respectively. The compression molding results in primarily one-dimensional flow in the x direction. Then the equation can be solved analytically. With the assumed boundary conditions of p , = 0 at x = M and ap, /ax = 0 at x = 0, the result is a parabolic pressure distribution as

a€

Fig. 26.7 Illustration of the consolidation model proposed by Gutowski et al. (1987a).


The solution for the fiber volume fraction Vf as a function of time is:

Po = Of(Vf) + 3 K _ V , T pa' dvf

Example problem 2 Compression molding with two-dimensional flow

Here the case of compression molding of a rectangular laminate with an isotropic in-plane (26.19)

* L

This expression shows how the applied pressure p, is carried by the fiber stress G~ and the average pressure in the resin. The load sharing in a composite is directly analogous to how the load is shared in a parallel spring and damper set. For example, initially if Vf is less than V,, then there is no deformation in the spring (fibers) and the entire load is carried by the resin. On the other hand, at long times and finite viscosity, if the rate of change of Vf is close to zero, then the pressure in the damper (resin) goes to zero and the total load must be carried by the fibers. Figure 26.8 shows an

permeability is considered. In other words, Kx = Ky = K. This may correspond to a quasi- isotropic lay-up. The flow equation becomes Poisson's equation, which can be solved by the separation of variables technique. The solution for the pressure distribution in a laminate with zero pressure at the boundaries is:

example of the one-dimensional flow in compression molding with the comparison of computer simulation results.

With the applied load balance condition, the final result is:

600 - PR , Theory

0 PR, Measured '3 400 - 0 -

Time ( m i d

Fig. 26.8 Example of one-dimensional flow in compression molding and computer simulation results (Gutowski, et al., 198%).

It can be seen that the result is analogous to the previous case except for a geometry effect term which is shown in the bracket.

Example problem 3: Bleeder ply molding

This has been presented with the Springer's model. In this case, a porous bleeder ply is placed on top of the composite, and flow is principally in the z direction. With the introduction of a new variable, the void ratio e = (1 - Vf)/V, one may obtain the nonlinear one-dimensional consolidation equation. An equation similar to this was first derived by Gibson et al. (1967) for the consolidation of saturated clays. The expression is:

- de = (e, + 1)2- a ( - Kz 'Of ") (26.22) at 3.z p ( l + e ) ' e az


The void ratio e or the fiber volume fraction V, is a function of both time and location. An equivalent equation using variables Vf and p , can be written as

With similar pressure equilibrium conditions, the distribution and time history of Vf or e can be solved numerically. Figure 26.9 shows an example of the bleeder ply molding measurement setup, and the comparison of the computer simulation results with the measured data.

It is interesting to see that, with Gutowski’s consolidation model, the final status of the composite in terms of the average fiber volume fraction can be estimated from the proposed fiber deformation model if the consolidation process is complete. The consolidation time for a particular setup can be solved through numerical simulation.

As can be seen from the analysis, the total consolidation time for a composite structure is strongly dependent on the dimension in the resin flow direction. For laminated composite structures, usually the dimensions in x and y directions (directions within the laminate structure) are much larger than that in the z direction (direction transverse to the laminate plane). For example, many aerospace structural parts range from a few inches to several feet in x or y direction, but only have a thickness of a fraction of an inch in the z direction. Thus the bleeder ply molding process is pre- ferred and is widely used in many part fabrication processes. However, for the so- called thick composites, for example with lay-up of 64 or 96 plies, the consolidation time required increases dramatically in the bleeder ply molding cases. With the selected cure cycle for thin composites, complete consolidation may not be achieved for thick composites. Thus the final fiber volume fraction of the thick composite tends to be relatively lower. This has been observed in experiments involving thick

DT

TRANSDUCER

/Applied Pressure

Modified Cormon-Kozemy

600

o doto 500 - -theory

- .- 400- v) a v

300 - 3

\o \O

L a IO0

0 P

0 10 20 30 40 50 60 Time (rnin)

Fig. 26.9 Example of bleeder molding and computer simulation results (Gutowski et al., 198%).

composites (Kim, Jun and Lee, 1989). It can also be seen from the comparison of

the two models that with relatively low fiber volume fraction, fibers carry almost no load. Thus the consolidation process is dominated by the resin flow through the fiber network. Then the difference between the two models is very minor. Springer assumes the consolidation is done layer by layer, while Gutowski treats the fiber network as a whole system. However, in both cases the top layers are consolidated first. When the fiber volume fraction becomes high, then the predictions from the


two models show significant different results. On the other hand, the numerical schemes of the two models are different. Springer’s model requires only the solutions of a series algebraic equations, while in Gutowski’s model nonlinear partial differential equations have to be solved. A comparison study has been presented by Smith and Poursartip (1993).

Specifically, a selected cure cycle must

the temperature inside the material does not exceed a preset value at any time during the cure;

2. at the end of the cure the resin content is uniform and has the desired value;

3. the material is cured uniformly and com-

ensure that:

26.3 CURE CONTROL

Fiber reinforced thermosetting resin composites manufactured in autoclaves are made by forming the uncured fiber-resin mixture into the desired shape and then curing the material. Curing requires the application of heat and pressure. Heat is used to facilitate and control the chemical reactions of the resin, and pressure is used to consolidate the composite, squeeze out the excess resin, and minimize the void content. A cure cycle usually means the magnitude, duration, and profile of the temperature and pressure applied during a curing process. Selection of the cure cycle directly affects the quality of the finished composite product, such as fiber content, fiber distribution, and void percentage.

pletely;

void content;

mal and mechanical properties;

4. the cured composite has the lowest possible

5. the cured composite has the desired ther-

6. the curing is achieved in the shortest time.

Figure 26.10 shows schematically the overall cure process model structure. In an early study, Loos and Springer (1983a) proposed a thermo- chemical model. Heat transfer from the environment to the composite material deter- mines the temperature distribution, the degree of cure of the resin, and the resin viscosity within the composite structure. The temperature inside the composite can be calculated using the law of conservation of energy. By neglecting the energy transfer by convection, the energy equation can be expressed as:

Viscosity b Flow

/ Reaction kinetics

\ Heat transfer Residual stress

Fig. 26.10 Schematic of overall cure process model (Dave et al., 1990).

Cure control 587

I I 1 1 I I

3501 - 6 -

0 5 rn coI/sec - 1 -

s\

I , , , , ~ , , l , l l l l

(26.24) dH - k - + p - i ~ ( zE) dt

Fig. 26.11 Rate of heat generation and rate of degree of cure of the

where p and Cv are the density and specific heat of the composite, kx, k and k, are the thermal conductivities, and ?is temperature. In the case of relatively thin composite structure, conduction heat transfer is mainly in the z direction. Thus terms in the x and y directions can be dropped. The rate of heat generation dH/dt is defined as:

(26.25)

where H, is the total heat of reaction depending on the resin type. The rate of the cure reaction is a function of temperature and the cure status, and can be expressed symbolically as:

(26.26) da dt - = f(T, 4

The degree of cure is then determined as:

a = I:($) dt (26.27)

It is assumed that for an uncured material, a = 0, and for a completely cured material, a approaches unity. As discussed earlier the

TEMPERATURE ( K l

resin viscosity, the degree of cure a and the rate of the cure da/dt can be characterized using a modified Arrhenius type equation, with relevant constants in the model determined experimentally (Lee, Loos and Springer, 1982; Bhi et al., 1987; Dusi et al., 1987; Roylance, 1988). Figure 26.11 show an example of the rate of heat generation and rate of degree of cure of the 3501-6 resin system as functions of time and temperature.

It is noted that the densityp, specific heat Cy, heat of reaction Hr, and thermal conductivity k are all dependent on the instantaneous and local resin and fiber contents of each ply, and

05 4 5 0 K 1 01

c3501-6 400K 1

o f \ , , Iojo\J 0 0 02 04 0 6 0 02 04 06

DEGREE OF CURE,a


can be handled using rule of mixtures (Loos and Springer, 1983c) or proposed approximate formulas (Springer and Tsai, 1967).

The solution to these equations can be obtained once the initial and boundary conditions are specified. The initial conditions require that the temperature and degree of cure inside the composite be given before the start of the cure. The boundary condition requires that the temperatures on composite surfaces in contact with the tool be known as a function of time during cure. Therefore the boundary condition is related to the specified cure cycle and the equipment setup.

The objective for the cure control scheme is to achieve the desired composite quality. Some of the main targets are reasonable temperature distribution, complete consolidation, minimum thermal stress and minimum void content.

With a developed numerical scheme, the temperature distribution inside the laminate is calculated as a function of position and time. A good cure scheme should realize the two main targets: (a) the temperature is reasonably uniform inside the material and (b) the temperature does not exceed a preselected maximum at any time.

For a given cure temperature and cure pressure, the time window for the consolidation is then specified. From the consolidation models, the compaction status of the consolidated composite can be obtained. In Springer’s model, the result is the total number of compacted plies, while in Gutowski’s model the result is the V, distribution across the layers. If the consolidation cannot be completed with the selected cure cycle, proper modifications are then made. The compaction issue becomes crucial to the cure process of the thick composite structure. A multiple stage heating process may be designed to defer the cure reaction of the resin and thus prolong the con-

pressure early in the cure cycle and the initial resin moisture are crucial considerations in producing void-free laminates (Kardos et al., 1983, 1988). Since the driving force for diffusion rises with temperature, in order to prevent the potential for pure water void growth by moisture diffusion in a laminate at all times and temperatures during the curing cycle, the resin pressure at any point within the curing laminate must be higher than the minimum resin pressure required, which is a function of the relative humidity and temperature (Dave et al., 1990). Figure 26.12 shows a void stability map for pure water void formation in epoxy matrices. A similar pressure requirement also holds for small air/water voids after an initial growth period. It has also been observed that the void content is reduced

(1 ATM I 101 kPI)

(RH), = 1ooO/o (RH), = 50%

300 400 500 1, K solidation time window.

Voids within the composite material are harmful to its mechanical Performance. Fig. 26.12 Void stability map for pure water void Experimental study shows that the resin formation in epoxy matrices (Dave et al., 1990).

Efects of tooling and part shape 589

significantly when the applied pressure is suf- ficiently high to collapse the vapor bubble before the gel point is reached. Therefore, after the time-temperature cycle is determined, it is possible to obtain a profile of the minimum pressure versus cure time. The boundary pressure is then maintained greater than the minimum pressure throughout the cure cycle.

During the cooling stage after the cure of the composites, residual thermal stress is related to the difference between the cure temperature and ambient temperature, and the thermal expansion behavior of the composite material. For a laminated structure, calculation of the thermal stress has been discussed and formulated by Tsai and Hahn (1980). Since the material shows viscoelastic behavior, stress relaxation has been observed over time. A post-cure process is usually applied to the structure to relieve the induced thermal stress.

For large complex-shaped composite structures, non-autoclave curing methods are used. Compared with traditional autoclave curing methods, the component size restrictions are eliminated, energy consumption is reduced, and capital equipment cost can be cut down. The non-autoclave processes use an oven, inte- grally reinforced tools, and presses. Major issues related to non-autoclave curing are the effective compaction of the composite plies, and the elimination of the trapped interlami- nar or intraply air.

26.4 EFFECTS OF TOOLING AND PART SHAPE

Properly designed tools that produce acceptable parts on a reproducible basis are a must when fabricating composite structures. The tool design requires the consideration of as many factors as are studied in the design of the part itself. The main requirement for the tools is to maintain proper geometric dimensional stability and surface profile during the compression and thermal cycling processes. On the other hand, the tool must also be

heated to a specified temperature at a specified rate under controlled conditions in the autoclave.

Tooling materials may be metal (steel, nickel, nickel alloys, and aluminum), graphite-epoxy and elastomer, depending on different composite part shape, size, volume of production and curing method. Selection of the tooling material often reflects a compro- mise among these considerations. Thermal behavior of the tooling material is also crucial in the design and fabrication. Table 26.1 lists the coefficient of thermal expansion of different composite and tooling materials. The values for the composites are dependent on the ply orientation and fiber volume fraction, and typical values are shown there.

Table 26.1 Coefficient of thermal expansion (CTE) for various materials (Borstell and Turner, 1987)

Material CTE ( I P / K )

Structural composite material Boron-poxy Aramid+poxy Graphiteepoxy Fiberglass-epoxy

Tooling material Graphiteepoxy Cast ceramic Tool steel Iron (electroformed) Nickel (electroformed) High-temperature cast epoxy Aluminum Silicone rubber

3.6-10.8

1.8-9.0 -2.0-5.8

7.2-9.0

4.1-9.0 0.81 11.3 11.9 12.6 19.8 23.2

81-360

26.4.1 TOOLING FOR AUTOCLAVE MOLDING

The traditional autoclave molding process uses a vacuum bag to impose a pressure difference on the composite lay-up. A typical bagging system consists of the following steps (Schwartz, 1983).


1. Cover the lay-up with a perforated parting film or separator. Then lay up a layer or layers of bleeder material. The requirement of the bleeder layers should be such as to ensure adequate bleeding of air and excess resin out of the part.

2. Place a strip of jute (vent material) just beyond the edge of the lay-up and put bag- sealing compound along the outside perimeter.

3. Cover the lay-up, jute, and sealing compound with a flexible-film diaphragm and seal the diaphragm to the mold with the seal compound.

4. Connect the vacuum lines and slowly apply the vacuum pressure while working the wrinkles and excess air out of the lay-up, bleeder material, and vacuum bag.

5. Check system for vacuum leaks. 6. Keep the part under vacuum while it is

waiting to be cured in the oven or autoclave.

To prevent surface irregularities on the bag side (untooled surface) of the parts, a caul plate may be used. The sole purpose of caul plates is to improve the visual appearance of the parts. They do not control part thickness. A flexible caul plate with a thermally stable rubber such as silicone or a fluoroelastomer is often used to accommodate the surface geometry. Figure 26.13 shows examples of autoclave tooling setups with caul plates.

The three issues related to the tooling design (Borstell and Turner, 1987) are thermal expansion correction, coordinating the location of partial plies and use of caul plates.

Because of the low coefficients of thermal expansion of composites when compared with metal tooling materials, thermal strain or stress must be considered for a curing process. In the autoclave, the temperature at which the resin solidifies is the gel temperature. At that specific temperature, the part is the same size as the thermally expanded mold. At a temperature above the gel temperature, the tool expands more than the partially cured part

Graphite-epoxy laminate Angle caul plate

Mold form -/ Caul plate stop

{Resin reservoir cdp\

Mold half-

- .Mold half

I \ C a o

Fig. 26.13 Example of autoclave tooling (Borstell and Turner, 1987).

introducing a thermal strain. As the part and tool cool down from the gel temperature, the tool usually shrinks more than the part. As an alternative, graphiteepoxy molds are used in some applications. Although some data has been published, not all composite materials have been measured. One empirical method is to cure a representative panel on a plate of the specified tooling material using the specified cure cycle. Corrections can be estimated by comparing the difference between the mold and part dimensions. Another recommended empirical correction method is to correct steel or nickel tools by making the tool 0.999 of the

Effects of tooling and part shape 591

engineering dimension, and to correct aluminum tools by 0.998. For example, a 2540 mm (100 in) dimension is tooled to be 2537 mm (99.9 in) for the steel tool. These corrections are needed to ensure an acceptable fit of mating composite parts.

Most parts contain partial plies to accommodate local areas of increased stress. Several techniques are used to control the location of partial plies, including polyester film templates, slotted templates, and rails and banking surfaces. These tools serve as supple- mental guidance to position the partial plies in the lay-up process.

Typical cases of applying a caul plate are to control the edge of a panel or the flanges of channels. The design of the metal caul plates must take into account the fact that the matrix resin melts in the autoclave to a very low viscosity. The caul plate performs by pushing excess resin sideways. Thus the rigid metal caul plates must have high rigidity so that they do not deflect under autoclave pressure at curing temperature. The thickness of the caul plates can be calculated by use of the equations for unsupported bending beam analysis. The deflection of the caul plate can be estimated using the balance condition of resin pressure and applied force (Gutowski and Cai, 1988). The caul plate deflections should be limited to half the tolerance permitted in the part.

26.4.2 ELECTROFORMED NICKEL TOOLING

An electroformed nickel tool consists of a 4.6-6.4 mm (0.18-0.25 in) thick electrode- posited mold surface that is supported by a simple steel substructure. The mold surface is produced by the electroplating process (Sheldon, 1987).

The electroformed tooling concept offers numerous advantages. The size of the mold is restricted only by the size of the electroforming tank. The cost of producing duplicated tools is low. The mold surface is very smooth and scratch resistant. The coefficient of thermal expansion is approximately 40% less than

aluminum. During autoclave curing of composite parts, the thermal uniformity is excellent with rapid heat-up and cool-down rates. It is easy to handle and transport because of its light weight. It also offers out- standing durability because the mold surface resists cutting or impact damage and is not thermally degraded. When damaged, it is easy to repair by welding, soldering, silver-soldering, or selective plating. It can provide complex contours without expensive machining. With most resin systems, it shows good release properties.

Figure 26.14 shows the procedures of making an electroformed nickel tool. As in some other types of tooling, constructing a model of the part surface is the first step in creating an electroformed mold. The models are the same net dimensions as the required nickel mold. Compensation may be required when the coefficient of thermal expansion of the composite part differs greatly from that of the nickel mold. Models are made from plaster, epoxy- faced plaster, fiberglass, fiberglass-epoxy, wood or other materials. From the model a reverse mandrel 'splash' is generally fabricated from epoxy-faced fiberglass or plaster. The mandrel to be used in electroforming is then copied from the 'splash', although the model can be used as the mandrel if it is pre- pared correctly. The comers of the mandrel should be designed to have radii in excess of 0.76 mm (0.030 in) to avoid thin spots in the deposit. Draft and taper should be designed into the mandrel to facilitate its removal from the electroform. Sharp corners or narrow, deep grooves should be avoided if possible. The mandrel can be fabricated from epoxy-faced fiberglass, rubber, or other materials. The surface of the mandrel is made conductive by proper coatings. The back of the mandrel must be reinforced to keep the mandrel from dis- torting during the electroforming process.

Electroforming is the process of producing an article by electrodeposition of a metal onto a conductive mandrel surface. An anode sus- pended in an aqueous electrolyte is connected


Model Splash

---t - Fiberglass plating mandrel Mold electroformed

I

i I

Plated mold and tool Mold and structure Plating mandrel upport structure joined removed

Fig. 26.14 Example of electroformed nickel tooling (Sheldon, 1987).

to the positive pole of a DC electric source, and the mandrel (cathode) is connected to its neg- ative pole. The flow of electricity or electrons results in the oxidation of a nickel anode to nickel ions and the reduction of nickel ions to nickel metal at the cathode (mandrel). The typical rate of growth is approximately 0.013-0.025 mm (0.0005-0.001 in) per hour. When the electroform is removed from the mandrel, its surface is a mirror image of the surface of the mandrel. A natural physical characteristic of electrodeposition is that electric current will tend to localize the deposit on all edges and corners, causing an uneven thickness on the electroform. However, there are a variety of techniques to offset this effect.

After the desired mold thickness is obtained, the mold is removed from the tank, cleaned and the steel back-up structure is attached. The nickel mold is then polished to the required finish, and ready for use.

26.4.3 GRAPHITE-EPOXY TOOLING

parts. These include low coefficient of thermal expansion, ease of preparation, low density, and thermal stability (Harmon, 1987). Their disadvantage is that they are less durable than metal tools.

Composite tool making starts with a master model, usually built with plaster or hard- wood. The master models require proper drying, sealing, and coating with mold release. Then lay-up can be done directly on the plaster or wood master. Liquid gel coats are required to obtain a high fidelity surface on tools cured by the vacuum bag process which does not generate enough pressure to ensure a void-free surface, but may not be required on tools cured by the autoclave process which does provide sufficient positive pressure. Prepregs with light weight fabrics are used directly against the tool surface, while prepregs with heavier fabrics are used to build up the thickness. During the lay-up, care should be taken to work each ply into all radii and corners and to remove all entrapped air. Debulking is applied after the lay-up, either with a vacuum

Composite tools have definite advantages over metal molds for large or highly contoured

bag setup or with -assistance of an autoclave for a pressure debulk, to consolidate the plies

Eflects of tooling and part shape 593

and remove all entrapped air. The curing process is done with a vacuum bagging system or with an autoclave. With the tool still on the model, the support structure, either a solid laminate or an ’egg-crate’ panel is attached to the tool by means of locally applied fabrics, room-temperature curing, and high-temperature resistant resins. Once the support structure is cured to the laminate shell, it is removed from the master. Care should be taken to avoid damaging either the tool or the master. Figure 26.15 illustrates the graphite-epoxy tooling making process.

Composite tools are being used successfully throughout the aerospace industry to produce parts that are structurally reliable, reproducible, and dimensionally accurate.

Teflon separator film Breather cloth

/ Vacuum bag

- ‘PFP master

Fig. 26.15 Example of graphite-epoxy tooling (Harmon, 1987).

26.4.4 ELASTOMERIC TOOLING

Elastomeric tooling or rubber tooling can be used to generate molding pressure or to act as a pressure intensifier. In thermal expansion molding, elastomeric tooling is constrained within a rigid frame to generate consolidation pressure by thermal expansion during the curing cycle (Foston and Adams, 1987).

In thermal expansion molding, two basic methods are employed: the trapped or fixed- volume rubber method and the variable-volume rubber method. Figure 26.16 shows the setup for both methods. The fixed- volume method exploits the large difference between the coefficient of thermal expansion of the elastomer and that of metals. The elastomer is confined within a closed metal tool

Rubber tool sized to fill the cavih, in the pan

,Pan

Floating-plate .--- pressure control

Rubber tool projects above the

by forcing the floating plate to the bag.

’ pan 30 excess pressure is vented

,Outer box M w Fig. 26.16 Example of elastomeric tooling (Foston and Adams, 1987) (a) fixed volume method; (b) variable volume method.


cavity. When heated, it expands into the cavity, exerting the pressure required to compact a composite laminate. The variable-volume method offers more flexibility and control than the fixed-volume method because a precisely calculated volume of rubber is not normally required. In most applications, the rubber is simply 'set back' to allow for the bulk factor of the molding material during assembly of the tooling details. A floating plate is used for the pressure control.

Thermal expansion molding with elastomeric tooling has been successfully used on commercial aircraft parts such as rudders and spoilers (Schneider and Carroll, 1987). This reduces the number of detail parts fabricated and the need for bonding and mechanical fas- tening on assembly, thereby effecting significant reductions in production time and cost.

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