residual stresses in carbon fibre-thermoplastic matrix laminates

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1988 22: 401Journal of Composite MaterialsG. Jeronimidis and A.T. Parkyn

Residual Stresses in Carbon Fibre-Thermoplastic Matrix Laminates

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401

Residual Stresses in Carbon

Fibre-Thermoplastic MatrixLaminates

G. JERONIMIDIS AND A. T. PARKYN

Department of EngineeringUniversity of Reading

P.O. Box 225

Reading RG6 2AYUnited Kingdom(Received May 3, 1987)

(Revised August 13, 1987)

ABSTRACT

Residual stresses in composite laminates depend on thermoelastic properties of thematerial and processing temperatures. Their distribution in the various laminae is a func-tion of stacking sequence and ply orientation. In this work residual stresses in APC-2cross-ply laminates have been investigated. Predictions based on classical laminate theoryare compared to measured levels of residual stress obtained from a number of experimen-tal techniques. The analysis of the results shows that accurate predictions can be made pro-vided that the changes in thermoelastic properties of the materials with temperature aretaken into account.

INTRODUCTION

ESIDUAL THERMAL STRESSES in composite laminated structures are an

Runavoidable consequence of the difference in coefficients of thermal expan-sion between fibres and polymeric matrices and of the cure or moulding tempera-tures of these materials. Their undesirable effects are twofold: distortions offinished components when cooled and removed from moulds (dimensional stabil-ity) and locked-in stresses. Tensile residual stresses in the matrix are particularlyimportant because they may represent a significant fraction of the tensile strengthof the polymer and can lead to premature failure. Prediction and measurement ofresidual stresses are therefore important in relation to production, design andperformance of composite components.

Residual stresses in thermosetting resins have received much attention in thepast [1-6]. In recent years new advanced composite materials based on high tem-perature thermoplastic resins have been developed [7-8]. Their moulding and

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402

consolidation temperatures are much higher than the curing temperatures ofthermosetting resins, such as epoxies, and hence greater temperature excursionsarise during fabrication. They also have well defined glass transition tem-peratures with corresponding significant changes in coefficient of thermal expan-sion and elastic properties.

In this paper we report the results obtained during the investigation of residualstresses in cross-ply laminates made of APC-2 (carbon fibre-PEEK) prepregssupplied by ICI plc. The polymeric matrix PEEK is semicrystalline and thedegree of crystallinity can be altered by changing the cooling rate of the com-posites after consolidation [9].

Several indirect methods for the measurement of residual stresses are dis-cussed ; direct methods, such as embedded strain gauges, have not been usedbecause the elevated processing temperature of APC-2 (about 400°C) would haverequired expensive high temperature gauges and complicated techniques for

embedding. Predictions based on classical lamination theory are compared withexperimental results: very good agreement has been obtained by taking into ac-count the changes in coefficients of thermal expansion and elastic properties ofthe matrix as a function of temperature. Most critical are the changes from aboveto below Tg and it has been found that about 75 % of the residual stress level inthe laminates builds up between Tg and room temperature. A very good agree-ment has also been found between residual stress levels and first ply failure ofsymmetric [02/902]s laminates and curvature of non-symmetric cross-plies.

PREPARATION AND CHARACTERISATION OF MATERIALS

Laminate Manufacture

APC-2 is provided in the form of a pre-impregnated tape of PEEK matrix rein-forced by continuous, unidirectional Hercules AS4 carbon fibres. Laminates aremade by stacking layers of tape in the desired orientations, melting the matrix andapplying a consolidation pressure. As far as possible, the manufacturers recom-mended cure cycle was followed:

Heating to 380°C under a contact pressure of .5 MPaHolding for five minutes at this temperature at a pressure of 1.4 MPaCooling under a consolidation pressure of 2 MPa

Heating and pressing took place in either a matched metal mould or in a pictureframe between soft aluminum sheets. In this work the third step of the recom-mended cure cycle was altered; the manufacturers, ICI plc., use a two presssystem in which transfer to a cold press allows cooling rates up to 1000°C/min(standard cooling rate being 40°C/min). Only a single press was available for thiswork, therefore cooling was achieved simply by turning off the heat and lettingthe mould cool under pressure, the measured cooling rate being approximately3 °C/min. Crystallinity is affected by cooling rate typically increasing from 25 %at 1000 ° C/min to 50% at 3 ° C/min [9]. According to the manufacturer’s data [ICIFiberite Data Sheet No. 2 (1986)] the slow cooling rates used in the experiments



403

will result in some reduction of composite toughness. Thus the samples preparedby slow cooling represent a lower bound in respect of the properties which areobtained.

Measurement of Elastic PropertiesCalculation of the residual stress requires the knowledge of the elastic proper-

ties and coefficients of thermal expansion of the laminae. The PEEK matrix ofAPC-2 has a glass transition at 143°C [ICI Literature Ref. VK2/0586 (1986)] andtherefore these properties need to be measured above and below Tg.At room temperature the elastic moduli in the fibre and transverse directions

(EP and Et) and the principal Poisson’s ratio (v,,) have been measured in tensionusing an Instron 1026 test machine and strain gauges bonded onto unidirectionaltensile specimens cut from a laminate using a diamond saw. The specimendimensions used were 175 x 12 x 1 mm for the longitudinal modulus andPoisson’s ratio and 100 x 12 x 3 mm for the transverse modulus. Dummygauges were used in each case for temperature compensation. The shear modulus(G,,) was measured in four point bending on a unidirectional square sample ofdimensions 48 x 48 x 1 mm [10] and in tension using strain gauges on a175 x 12 x 2 mm [ =i:45] specimen. The values obtained from all these testswere within 4 % of the values quoted by ICI plc. for APC-2 and so publishedvalues were used in the calculations of residual stresses.

Figure 1. Dynamic-mechanical measurement of the temperature dependence of the prin-cipal elastic moduli Ee and E~ of APC-2.



404

11:&dquo;’1-’1:’1 dlUI~ t k-I

Figure 2. Temperature dependence of the transverse coefficient of thermal expansion ofAPC-2.

High temperature measurements of E, and E,, below and above Tg were ob-tained using a Polymer Laboratories Dynamic Mechanical Thermal Analyser ata frequency of 3Hz between 20 and 300°C. The longitudinal modulus remainseffectively constant over the temperature range while transverse modulus dropsrapidly beyond Tg from 8.9 GPa at 150°C to about 1 GPa at 300°C. It was not

possible to take measurements above 300°C (Figure 1).

Coefficients of Thermal ExpansionThe coefficients of thermal expansion in the fibre and transverse directions

were measured using high temperature strain gauges WK-06-250BG-350 bonded

Table 1. Measured thermoelastic properties of APC-2.

*Assumed value.Note’ The shear modulus above Tg was not measured since it does not affect residual stresses of cross-ply laminates.



405

with a high temperature adhesive M-Bond 610. Measurements were limited toabout 240°C, the adhesive having been cured at 280°C. Thermal strains in thegauges were compensated using a half bridge configuration where one arm wasthe gauge on the APC-2 sample, the other being a strain gauge bonded with thesame adhesive on a block of EN-58B stainless steel with a known coefficient ofthermal expansion (cx = 17.7 ~6/°C). The APC-2 specimen and the stainlesssteel plate were placed in a small insulated enclosure and the temperaturechanges were monitored by thermocouples. Figure 2 shows the results obtainedfor the transverse coefficient of expansion, at ; the output of the strain gaugebridge is plotted versus temperature after subtracting the expansion strain coeffi-cient of the stainless steel. The change of slope occurring between 120°C and160°C is associated with the glass transition of the PEEK matrix. From thesegraphs the best fit straight lines through the experimental points were used for thedetermination of the coefficients of thermal expansion.The measured values in both longitudinal and transverse direction are in good

agreement with experimental results obtained from ICI plc.The thermoelastic properties used in the analysis are summarised in Table 1.

Stress Free Temperature

It has been suggested by Hahn [6] that with thermosetting resins the tempera-ture at which stresses in a laminate begin to build up may be lower than the curetemperature. If this is not taken into account there will be an overestimate ofresidual stress. These considerations do not apply to thermoplastic matrixmaterials; however, stress build up will not necessarily start upon cool down butmay occur at the onset of crystallisation or at the glass transition. The precisevalue of the stress free temperature may vary depending on materials and oncooling rate.

In this paper the method used to measure the stress free temperature (Ts/) isthe logical extension of the method used to measure residual stresses inunbalanced laminates, where curvature is developed in proportion to residualstress. A number of laminates and strips cut from laminates, all of varying thick-ness, were heated in a small enclosed space until they became flat. At this tem-perature all residual stress was assumed to have been removed. From all the

samples tested the Tsf was found to be 310°C. This figure corresponds to peakcrystallisation temperature for APC-2 cooled at 3°C/min [9] i.e. the cooling ratemeasured with the one press system used in this work.Other authors [11] using similar experiments with APC-1 found higher values

for Tf , 328°C on heating and 319°C on cooling; details of cooling rate were notincluded. It should be noted that APC-1 and APC-2 are not identical and that theAPC-1 laminates were moulded at a lower pressure (0.5 MPa). The small dif-ferences between the two sets of measurements may be due to these factors.

THEORY AND PREDICTION OF RESIDUAL STRESSES

Classical laminate theory was used to calculate residual stresses and curvaturesin cross-ply laminates, symmetrical and non-symmetrical [12]. In the absence of



406

external loads the midplane strains (0°) and curvatures (k ,0) can be related tothermal forces and moments by:

(A], [B], [D] are the extensional, coupling and bending stiffness matrices ofthe laminate, ef, E y , E;;’ are the midplane strains and k °, ~~, k °, the midplane cur-vatures (referred to laminate axes x and y). NX , NJ, N’ , and Mx , MJ, M y, arethe thermal force and moment resultants respectively. For cross-ply laminates thethermal forces and moments are given by:

with

where

Q ; = stiffness constants of the kth layera x , a y = coefficient of expansion in x,y direction of lamina kzk , zk-1 = coordinate of upper and lower surfaces of lamina relative to the

midplane (z = 0)AT = temperature interval over which residual stress has built up

The residual stresses within each lamina are calculated from:

Two residual stress distributions, calculated using Equations (1-3) and the dataof Table 1, are shown in Figures 3 and 4 for [902/04/902] and [On /90n] ] laminates



407

Figure 3. Calculated residual stress distribution m a [90¡OJ90J balanced cross-ply lami-nate. The maximum tensile stress m the 90° pl1es and the maximum compressive stress inthe 0° pl1es have a value of 42 MPa.

respectively. These same configurations have been used for the measurement ofresidual stresses. The temperature dependence of the thermal and elastic proper-ties was taken into account by calculating the residual stresses incrementally: twosteps were used, the first from 310°C to 143°C and the second from 143°C to20°C, using the appropriate values given in Table 1. The choice of 1 GPa for E,in the first step was made for two reasons: (a) it is the value corresponding to themid point of the temperature interval 310 -143 °C on the graphs obtained from

Figure 4. Calculated residual stress distribution in a [OJ904J unbalanced cross-ply. The ten-sile stress in the 90° layer varies from 24 MPa on the outer surface to 34 MPa on the innerone. In the 0° layer the residual stress varies from -195.5 MPa (compressive) on the innersurface to + 138.3 MPa on the outer one.



408

the dynamic mechanical tests when the rapidly decreasing part around 300°C isextrapolated to 310°C, (b) it gave a good agreement with the initial set of mea-surements and thereafter was used consistently throughout this work.

In both cases the contraction of the 90° ply against the restraint of the 0° plyputs the former in tension and the latter in compression. It should be noted thatit is the ratio of the 0° to 90° plies in the laminae which determines the residualstresses, rather than the total thickness of the laminate.The important result is the high transverse tension (42 MPa) experienced by the

90° plies of the [902/04/902] laminate. This stress corresponds to about 50% ofthe transverse tensile strength of the material (Figure 3). Although significant,this figure compares very favourably with the values measured by Flaggs andKural [13] on T300/934 carbon fibre/epoxy cross-ply laminates. The residualstresses they measured are slightly higher than in the corresponding APC-2laminate, about 45 MPa, but in the T300/934 system they are approximatelyequal to the transverse strength of the material and transverse ply failure is

prevented only by a crack suppression effect. Transverse cracking due to residualstresses alone has been observed in other systems by Nairn and Zoller [11] and itwas noted that the tendency to transverse cracking was much less in APC-2.

MEASUREMENTS OF RESIDUAL STRESSES

Balanced Laminates

FIRST PLY FAILURE METHODThis is a method similar to that used by Kim and Hahn [14] using graphite/

epoxy laminates. It has been predicted that in a [902/04/902] laminate the contrac-tion of the outer transverse plies against the restraint of the inner longitudinalplies would cause a tensile stress in the transverse plies of approximately 42 MPa.Therefore the strength of the transverse plies should be correspondingly reducedby 42 MPa.

Extensive tests on the specimens used previously gave the strength of APC-2in the transverse direction as 80 MPa (S.D. = 2 MPa) in tension (in agreementwith ICI published data) and 107 MPa in bending (S.D. = 10 MPa). The residualstrength of the transverse plies in a [902/04/902] laminate should therefore be 38or 65 MPa depending on test method.

Five specimens of dimensions 50 X 15 X 0.85 mm were tested in three pointbending in an Instron 1026 tensile test machine at a crosshead speed of 4

Table 2. Effect of temperature on first ply failure.



409

- ---

Figure 5. A transverse crack m the outer ply of a balanced laminate (outer ply thick-ness = 0 25 mm)

mm/min. A failure in the outer 90° ply on the tension side being shown by ajump and a change in slope in the load deflection graph. This occurred at anaverage load of 20N (range 18-22.5 N) which corresponds to a maximum stressin the 90° ply of 73 MPa.The tests were repeated at - 30 ° C and - 50 °C to increase the residual stress

which should in turn decrease the stress required to break the 90° ply. The resultsare shown in Table 2.The agreement between calculated and measured values is very good at room

temperature, less so at the lower temperatures and this may be due to increasedbrittleness of the matrix and increased sensitivity to defects.

In tension the detection of first ply failure through a change of slope in the load-extension curve does not work since the longitudinal plies carry most of the loadand the failure of the transverse ply has little effect upon the stiffness or load bear-ing capacity of the laminate. However, audible emissions occurring during thetest gave an indication of when a crack appeared in the transverse ply. This wasascertained by a series of tests on a number of specimens from various balancedlaminates the edges of which were polished for microscope observations and in-spected before testing to check that no cracks were present. In a few preliminaryexperiments the tests were stopped and the specimens unloaded before detectingany sound so as to ensure that no cracks formed without a corresponding noiseemission. The specimens were of dimensions 175 x 20 x 0.9 mm and weretested at a crosshead speed of 4 mm/min. As soon as a crack was heard duringa test the specimen was unloaded and examined under the microscope; in allcases there was an excellent one-to-one correlation between audible emission andcrack formation. Figure 5 shows a typical first ply failure on the outer 90° ply ofa [902/04/902] laminate. The average measured value for first ply failure was 39MPa (S.D. = 2.4 MPa) which agrees very well with the predicted value of 38



410

Table 3. Effect of laminate configuration on first ply failure(when crack suppression effect is reduced).

------

MPa, given by the difference between tensile strength (80 MPa) and residualstress (42 MPa).The laminate lay-up used in these experiments had the 90° plies on the outside

of the laminate whereas in similar work [13-16] these plies have been sandwichedbetween 0° plies. There are two reasons for this:

1. Any fibre misalignment due to flow in the mould will produce a slightlywarped laminate which causes difficulty in end tabbing. However with the 90°plies on the outside of the specimen the jaws of the testing machine can gripfirmly without end tabs and this simplifies procedures during tensile testing.

2. It has been noted by various workers [13-16] that in any balanced cross plylaminate the failure of the transverse plies can be suppressed as their thicknessis reduced. Parvizi, Garret and Bailey [15] and Garret and Bailey [16] do notmake any allowance for residual stresses while Flaggs and Kural [13] and Kimand Hahn [14] do but in all cases it was noted that the transverse strength of the90° plies increased with decreased thickness of the ply.In this work the need to have a constant transverse strength value as reference

for the 90° plies meant that these crack suppression effects had to be eliminatedor at least reduced. This was achieved by testing the balanced cross-ply laminatesin the 90/0/90 configuration with the fibres of the 0° ply parallel to the axial load.In this manner, the outer 90° layers are supported only on one side by the 0°layer, the other side being free of constraint. This is confirmed by tests on thelaminate configurations shown in Table 3 where first ply failure agrees well withpredictions and there is no evidence of crack suppression. It may be noted thatthe transverse residual stress level for each cross-ply laminate configuration isvery similar; this is not as suprising as it first appears, being a consequence ofthe large difference in moduli between the two principal directions. For differentthickness ratios of 90° to 0° plies the residual stresses in the 0° plies vary from-20 MPa for a ratio of 1:2, to -115 MPa for a ratio of 3:1.

MILLING METHODThis is an indirect method of measuring the stresses present in balanced lami-

nates ; it involves milling away part of the surface of the laminate to produce anunbalanced laminate, the curvature of which can be predicted as shown pre-



411

viously. The top 90° layer of a [902/04/902] laminate was removed by mill-ing-leaving a [02/904] laminate. The radius of curvature measured after millingwas 375 mm to be compared with a predicted radius of 311 mm.

This is a reasonable agreement especially in view of the difficulties of millingaccurately through one layer; any irregularity in thickness of the laminate meansthat part of the lower layers in one region may have to be removed in order toremove all of the upper layer. This will, of course, affect the curvature.

Unbalanced Laminates

CURVATURE MEASUREMENTIn unbalanced laminates the residual stresses cause out of plane deformation.

This feature has been used by Nairn and Zoller [11] to follow the build up ofresidual stresses as a small unbalanced composite strip was cooled. These cur-vatures have then been compared with predictions from a modified bi-metallicstrip model [17]. The measurements shown here were taken from the actual lami-nates and compared to predictions from classical lamination theory. Laminatecurvature is related to residual stresses, therefore, if the theory can predict thecurvatures accurately, it can be assumed that the residual stress levels are also ac-curate and that appropriate thermoelastic properties have been used.A number of [90,,/0,,] cross-ply laminates were made and their curvatures

measured and compared with the predicted values in Table 4.Each of these laminates and, indeed, all unbalanced laminates used in this

study are bi-stable, i.e., although classical lamination theory predicts a simul-taneous double curvature or saddle shape, laminates which are thin by compari-son with their in plane dimensions take a cylindrical shape which can be &dquo;flipped-over&dquo; to form a second cylinder. In the case of the two and four ply doublecurvature laminates in particular, one curvature tends to restrain the other. Inorder to measure the residual curvatures as accurately as possible, strips about 15mm wide were cut from each laminate in two orthogonal directions. No less thanthree laminates were used for each of the results in Table 4. The restraining effectwas most marked in the two ply laminates, rapidly becoming less with eightplies.

It has been noted by Hyer [18,19] that for laminates with a sufficiently large

Table 4. Radius of curvature of various crossply laminate configurationsas a function of thickness.



412

-

(D)

Figure 6. Effect of mould shape on residual curvatures, 0/90 crossply made on: (a) flatmould showing bistable behaviour on cooling with two curvatures of equal magnitude andopposite signs in the x and y directions, and (b) curved (cylindncal) mould showing, on cool-ing, flattening in the x direction and curvature in the y direction.

ratio of thickness to in plane dimensions the curvature of the cylinder is approx-imately the same as that predicted for the saddle, i.e., the same as that predictedby classical lamination theory. The ratio of thickness to in plane dimensions inthe laminates investigated in our work is of the order of 1:1000, 1:500 and 1:250for two, four and eight ply laminates respectively and comparing cylinder cur-vatures with strip curvatures it was found that the approximation held within thelimits of experimental error.

In the course of this work it has also been noted that the room temperature cur-vature of unbalanced laminates decreased with time. This effect is more pro-nounced than with thermosetting resins where moisture uptake can produce areduction in curvature. In APC-2 moisture uptake is negligible [20] and this ef-fect therefore seems to be related to stress relaxation characteristics of thematerials used and is being investigated.

CURVED MOULDINGFrom Table 4, the curvature acquired by a [04/904] laminate of thickness 1 mm

is 201 mm.In order to put our predictions to the test, a [04/904] cross-ply laminate was

made between a pair of matching curved metal moulds with one principal radiusof curvature equal to the predicted value of 201 mm, and the other, orthogonal tothe first, equal to infinity; during heating and consolidation the laminate is heldin the shape of a cylindrical shell.A 0/90 cross-ply laminate made on a flat mould acquires two principal cur-

vatures of opposite signs on cooling and the characteristic saddle shape (which at



413

this thickness becomes two cylindrical states). The [04/904] cross-ply made onthe cylindrical mould should flatten in the direction corresponding to the cir-cumference of the cylinder and acquire a radius of 201 mm in the direction of theaxis.The stacking sequence of the laminate is given in Figure 6 showing that the

transB °rse contraction of the upper ply would cancel out the moulding curvature.After cooling the laminate was removed from the mould and measured. As pre-dicted, the radius of curvature in the circumferential direction of the mould wasfound to be infinite, and the other 201 mm. Although the laminates made in theflat mould and in the curved mould look identical because of the small thickness,the former can flip between two states corresponding to +R and -R [Figure6(a)]; the latter can not [Figure 6(b)].

DISCUSSIONS AND CONCLUSIONS

The effects of residual stresses in thermoplastic matrix composite laminatescan be predicted using classical lamination theory provided that the changes ofthermal expansion coefficients and moduli of the matrix with temperature aretaken into account.There are however a number of problems. With APC-2 accurate measurements

well above Tg are difficult. The decrease in transverse modulus in particular,could not be monitored accurately above 280°C; there is some evidence (Figure1) to suggest that a significant drop occurs near this temperature. In our calcula-tions we have taken a &dquo;mean value&dquo; over the temperature interval TSf- Tg. This isjustified by the fact that the inaccuracies in E, between 280-310°C affect only 5 %of the total residual stress at 20°C.A second difficulty in calculating residual stresses is introduced by stress relax-

ation effects during cooling, especially during the early stages. These may be im-portant given the slow cooling rates used in this work (although tests carried outon standard, 40°C/min and fast cooled, 1000°C/min, unbalanced laminates gavesimilar residual curvature measurements to the slow cooled ones). These stressrelaxation effects in APC-2 have not yet been fully quantified but are now beinginvestigated. Neglecting stress relaxation in the calculations will overestimate re-sidual stresses; this is equivalent to overestimating the transverse modulus E,which decreases with both temperature and duration of application of stress. Thegood agreement obtained between predicted and measured values of residualstress and residual curvatures may mean either that the value chosen for E, com-

pensates for stress relaxation effects to some extent or that stress relaxation is sig-nificant only in the high temperature portion of cool down (400-310°C) where E,is very small anyway (0.1-0.2 GPa) and has little effect.The results obtained show that in APC-2 cross-ply laminates, the residual

transverse tensile stress at room temperature is about 50% of the transverse

strength of the material and, as with all composite materials, this needs to betaken into account in the design of components. Although first ply failures due totransverse cracking may not reduce significantly the strength and stiffness ofcross-ply laminates, they may lead to delamination of the interface between the



414

0° and the 90° layers with significant loss of the integrity of the laminate andreduction in compressive strength.The work reported in this paper does not take time effects into account but

there are indications that their importance in thermoplastic matrix composites isgreater than in thermosetting materials. Stress relaxation effects on residualstresses and residual strength need to be investigated further for a more completedescription pf the behaviour of these materials.

ACKNOWLEDGEMENTS

The authors wish to thank the Science and Engineering Research Council forthe CASE award supporting this work and ICI plc. for supplying materials andfor many helpful discussions.

REFERENCES

1. Daniel, I. M. and T Liber "Lamination Residual Stresses in Fiber Composites," IITRI ReportD6073-I, NASA CR-134826 (1975).

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3. Daniel, I. M. and T Liber. "Effect of Laminate Construction on Residual Stresses in

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(1984).9. Blundell, D. J. and B. N. Osborn. "Crystalline Morphology of the Matrix of PEEK-Carbon Fibre

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Composites," Proc. 5th Int. Conference of Composite Materials, San Diego (1985)12. Jones, R. M. Mechanics of Composite Materials. Tokyo:McGraw-Hill Kogakushu Ltd. (1975).13. Flaggs, D. L. and M. H. Kural. "Experimental Determination of the In Situ Transverse Lamina

Strength in Graphite/Epoxy Laminates," J. Composite Materials, 16:103 (1982).14. Kim, R. Y. and H. T. Hahn. "Effect of Curing Stresses on the First Ply-Failure in Composite

Laminates," J. Composite Materials, 13:2 (1979).15. Parvizi, A., K. W. Garret and J. E. Bailey. "Constrained Cracking in Glass Fibre Reinforced

Epoxy Crossply Laminates," J. Material Science, 13:195 (1978).16. Garret, K. W. and J. E. Bailey. "Multiple Transverse Fracture in 90° Crossply Laminates of

Glass Fibre Reinforced Polyester," J. Material Science, 12:157 (1977).17. Nairn, J. A. and P. Zoller. "Residual Stresses in Amorphous and Semicrystalline Thermoplastic



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Matrix Composites," ASTM Special Tech Publ 937 on Toughened Composites. Philadelphia.ASTM (1986)

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residual stresses in carbon fibre-thermoplastic matrix laminates

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