Composites: Part A 40 (2009) 1534–1544

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Composites: Part A

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Assessment of cure-residual strains through the thickness of carbon–epoxylaminates using FBGs Part II: Technological specimen

M. Mulle a,*, F. Collombet a, P. Olivier a, R. Zitoune a, C. Huchette b, F. Laurin b, Y.-H. Grunevald c

a Université de Toulouse, INSA, UPS, ICA (Institut Clément Ader), 133, Avenue de Rangueil, F-31077 Toulouse, Franceb ONERA, BP 72 – 29, Avenue de la Division Leclerc, 92322 Châtillon Cedex, Francec Composites Expertise and Solutions, 31320 Castanet Tolosan, France

a r t i c l e i n f o a b s t r a c t

Article history:Received 7 November 2008Received in revised form 11 May 2009Accepted 19 June 2009

Keywords:A. Technological specimenB. Residual/internal stressC. FBG sensorsD. Process monitoring

1359-835X/$ - see front matter � 2009 Elsevier Ltd.doi:10.1016/j.compositesa.2009.06.013

* Corresponding author. Tel.: +33 05 62 25 88 72.E-mail address: matthieu.mulle@iut-tlse3.fr (M. M

This paper is a continuation of our previous study [Mulle M, Collombet F, Olivier P, Grunevald Y-H.Assessment of cure residual strains through the thickness of carbon–epoxy laminates using FBGs, partI: elementary specimen. Compos Part A 2008. doi:10.1016/j.compositesa. 2008.10.008] pertaining tothe assessment of autoclave cure-induced strains through the thickness of carbon–epoxy laminates. Inthis first part, postulates and measurement procedures were established for cure of elementary speci-mens. Based on these, this study undertakes investigation on what are called technological specimens.These specimens are of the beam type and contain geometrical specificities which represent typical struc-tural issues. In-plane process-induced strains were studied through the thickness of a thick reinforcedzone using several optical fibre Bragg gratings (FBGs) sensors embedded at different levels of the plystack. A non-uniform distribution of residual strains was detected. Once cured, the technological speci-men was subjected to a heating test whose cycle was comparable to the cure cycle. Thermally inducedstrains were measured with the embedded FBGs. The values recorded were compared with those ofcure-induced residual strains and FEM simulation. Discrepancies were observed that strongly suggestthe possible influence of environmental effects and the need for the calculation to take into accountthe through-the-thickness variability of thermal properties.

� 2009 Elsevier Ltd. All rights reserved.

1. Introduction

Many obstacles are encountered in designing composite struc-tures. These obstacles are related to the multiplicity of the scalesinvolved, the complexity of the mechanical response, the lack ofknowledge of the material state in the structure, and the unavoid-able variability induced in properties by the various manufacturingphases. Usually, the material properties are identified through testscarried out on elementary coupons. This step is the base of anexperimental building block procedure. However, identifying thematerial parameters is only meaningful if the response of the com-posite material is studied within the structure. To address this is-sue a technological specimen was designed [2–4]. This is anobject which relates the elementary coupon to the industrial struc-ture. It is of the beam type and contains design singularities such asply drops and reinforced zone (Figs. 1–3). The specimen enablesstudy of various thermo-mechanical aspects [5–7]. The first isone that determines the specimen’s initial state in terms of cure-induced residual stresses.

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ulle).

The preliminary phase of such an analysis was carried out inPart I of this study [1]. In-plane cure-induced residual strains wereassessed through the thickness of a carbon–epoxy elementaryspecimen using optical FBG sensors. Earlier, postulates and mea-surement procedures were established that were used to monitorthe curing process.

Based on these procedures, the present investigations were fo-cused on the central reinforced zone of the technological specimenwhere several embedded optical sensors were distributed allthrough the thickness of this zone.

This paper is divided into four main parts. The first describes thegeometry, the material, and instrumentation of the technologicalspecimen. A detailed account of the calibration procedure of theoptical sensors is also given. The second part relates to the exper-imental set up and measurement procedures adopted. As regardsthe monitoring of the autoclave curing of the specimen, the infor-mation delivered by the FBG sensors was gathered all through themanufacturing process. The distribution and amplitudes of theprocess-induced strains were measured. Simultaneously, anembedded highly birefringent FBG (HIBI), particularly sensitive totransverse strains, provided a means to quantify any of the result-ing plane effects [8–10]. After cure, the specimen was subjected toa series of heating tests in an oven in order to establish the thermal

50 mm

FBGs in plies 3, 10, 19 and 26

30 mm

20 mm

Scheme without scale

Ply drop off zone

Longitudinal axis

x z

x y

10°

300 mm

Thin zone

Optical fibres

Reinforced zone

5.5 mm 7.7 mm

Fig. 1. Plan of the technological specimen. Dimensions and FBG instrumentation.

Fig. 2. Photograph of two technological specimens.

Fig. 3. Photograph of the ply drop-off zone.

M. Mulle et al. / Composites: Part A 40 (2009) 1534–1544 1535

expansion behaviour of the reinforced zone [11]. The heating pro-cedure corresponded to the cure cycle, subjecting the material toconditions close to those encountered during its manufacture. Thisimplies need for a discussion about the stress free temperature.The third part describes a numerical model (FEM) that was devel-oped to simulate the heating tests. Thermal strains were predicted.In the last part, experimental results are analyzed before beingcompared to theoretical results. Some discrepancies observedraised questions related to environmental and structuralinfluences.

2. Technological specimen

2.1. Geometry, material and stacking sequence

The technological specimen is of a beam type whose length/width ratio is 10. It contains a central reinforced zone and thus aply drop-off at each end of the reinforced zone. Dimensions ofthe specimen are presented with photographs in Figs. 1–3. Thespecimen is intended to represent the skin of a wing of an aero-plane with a local reinforcement likely to support a load with abolted assembly. The tapered angle of the ply drop-off is 10�. It lieswithin typical values found in the aeronautics field [12,13]. The

reinforced zone consists of plies added in the form of a patch andis sandwiched between two half thicknesses of the thin zone.These specimens were manufactured in a series of 16 pieces fromof the same mother plate. Pre-preg laying up of the mother plate(in the clean room) is shown in Fig. 8b. After the autoclave cure,specimens were cut out of the plate by means of abrasive waterjet machining.

Specimens are made of a carbon/polymeric Hexply M21 T700(reference: M21/35%/268/T700GC) pre-preg manufactured byHexcel Composites. They were autoclave cured based on the Man-ufacturer’s Recommended Cure Cycle (MRCC, cf. Section 3.1). TheM21 matrix combines a thermoset epoxy with thermoplasticphases (in both liquid- and solid phases. Fig. 4 shows the presenceof thermoplastic nodules and the manner of their random distri-bution through the laminate although a large portion of them liesbetween the plies and yarns. The carbon fibres are of TorayT700GC.

For the stacking sequence, a typical aeronautical sequence wasadopted [12,14] (Fig. 5). Specimens are made of 20 plies in the thinlateral zones (5.5 mm thick) and of 28 plies in the reinforced zone(7.7 mm thick). The thin zone has the following reinforcement dis-tribution: 50% at 0�, 40% at ±45� and 10% at 90�. The reinforced cen-tral zone has a complementary embedded patch with the followingreinforcement distribution: 25% at 0�, 50% at ± 45� and 25% at 90�.The 0� direction of the laminate is in the longitudinal direction ofthe specimen (x axis).

2.2. Instrumentation

As stated in part I [1], the working principle of FBGs has beendetailed in various papers in the last decade [15–17]. FBGs aremainly sensitive to temperature (T) and axial strains (exx). They de-liver information which is related to a variation of the Bragg wave-length (DkB) as shown in

DkB

kB¼ KTDT þ Keexx: ð1Þ

x 150

Fibres direction 0°

90°

45°

0.275mm

x 100 - 45°

90°

0°

Inter- yarns

Inter- plies

Fibres direction 45°

0.275 mm

Fig. 4. SEM observations of a sample of M21 T700 GC composite showing distribution of the thermoplastic nodules.

N° Orientation28 0°27 45°26 0°25 -45°24 90°23 0°22 -45°21 0°20 45°19 0°18 45°17 90°16 -45°15 0°14 0°13 -45°12 90°11 45°10 0°9 45°8 0°7 -45°6 0°5 90°4 -45°3 0°2 45°1 0°

Ply

stac

king

seq

uenc

e

10 %

50 %

40 %

10 %

50 %

40 %

25 %

25 %

50 %

FBG

FBG

FBG

FBG

0°

Fig. 5. Stacking sequence of the technological specimen and position of embeddedFBGs.

1536 M. Mulle et al. / Composites: Part A 40 (2009) 1534–1544

Coefficients KT and Ke depend on the nature of the optical fibreand the FBG inscription parameters and are related, respectively, tothe temperature and mechanical strain sensitivity. As temperatureand mechanical stress may act on the embedded sensor at thesame time during the manufacturing process, it is necessary to sep-arate their influences on the information from the FBG. Variousmethods of discrimination exist and are available in literature[5,15–17]. In part I [1] we used a combination of a bare FBG withan encapsulated FBG. This technique is useful because only onesensing technology is employed, which minimizes uncertaintiesand simplifies the examination of the results. However, the tubeprotecting the end tip of the fibre is too intrusive to enablemechanical testing of the instrumented specimen. Since it hadbeen decided to analyze the mechanical behaviour of the techno-logical specimen through a series of bending tests [7], such a meth-od for discrimination was not appropriate.

The technique employed here is a combination of an FBG and athermocouple. The latter was embedded in the mother plate at asufficient distance from the FBG so that it does not affect the spec-imen’s mechanical behaviour, but in similar conditions in order toenable sensing of identical heat changes during the curing process.

The thermocouple thus delivered the thermal component of the to-tal FBG information (Eq. (1)). Such a technique calls for a very pre-cise calibration procedure of KT and Ke coefficients.

The calibration of KT and Ke was carried out using an ovenequipped with holes at the top and bottom through which the opti-cal fibre passes as shown in Fig. 6. One end of the optical fibre isconnected to the interrogation unit and the other is attached to asmall tray to receive variable loads. Two sets of experiments werecarried out in the loading and temperature range that encounteredthe sensor during the composite curing process. First, the FBG wastested at various isothermal dwells during which the load varied.This enabled the determination of the strain sensitivity (Ke kB) forvarious temperatures. For the second set of experiments, the FBGwas tested during heating ramps while the load remained con-stant. This enabled the determination of the temperature sensitiv-ity (KT kB) for various values of tensile stress.

For this study, the FBGs employed were inscribed in the stan-dard SMF-28e� single mode optical fibre by exposing a segmentof the fibre to an interference pattern (using a phase mask) of ul-tra-violet light. The Bragg gratings (10 mm long) were uniform,apodized and centred on a 1535 nm wavelength.

Results of the calibration procedure showed that the strain sen-sitivity coefficient remains constant in the [0�180 �C] temperaturerange, Ke = 0.78 � 10�6 le�1, whereas the temperature sensitivitycoefficient varies with temperature (Fig. 7): for FBG withkB = 1535 nm, KT = 5.6 � 10�9T + 6.2822 � 10�6 �C�1.

Four FBGs were embedded in the reinforced zone during thepre-preg lay up at different levels, in the longitudinal direction ofthe beam (Fig. 8). They are positioned 20 mm from the longitudinalcentre of the specimen, in plies 3, 10, 19, and 26, as shown in Figs.1, 5 and 8a.

3. Experimental set up and procedures

3.1. Autoclave cure procedure, optical measurement system, andmould

The cure cycle consisted of raising temperature at the rate of2 �C per minute up to 180 �C, an isothermal dwell at this tempera-ture for duration of 120 minutes and a cooling phase with the tem-perature falling at the rate of 2 �C per minute. The whole cycle wascarried out at a pressure of 7 bars and �0.7 bars of vacuum.

The experimental set up was almost identical to that used forthe curing procedure of elementary specimens described in theprevious paper [1]. The optical measurement loop essentially in-cluded a Bragg wavelength interrogator (MicronOptics si425) andan optical spectrum analyzer (MicronOptics si720). The only differ-ence laid in the material used for the mould. Here it was a 5 mmthick aluminium plate. As the thermal expansion of aluminium

(a) (b)

Variable heating procedures

Variable loading

OF connected to Interrogation units

FBG

TC

Fig. 6. Sketch (a) and photograph (b) of the experimental set up for the FBG calibration procedure.

Fig. 7. Variation of temperature sensitivity coefficients of two FBGs versus temperature. For cure monitoring FBGs centred on kB = 1535 nm were used.

M. Mulle et al. / Composites: Part A 40 (2009) 1534–1544 1537

(a = 23�6 �C�1) is greater than that of steel (a = 12�6 �C�1), it wasexpected that mould/part interactions would be accentuated andfacilitate the analysis.

3.2. Complementary optical line: FBG on HIBI fibre

In part I [1] we tried to detect ultimate out-of-plane strains onthe FBG that could be induced by the onset of pressure and vacuumor by thermal contraction of the matrix. None of these effects wereperceptible to the optical sensors. In fact, FBGs written on standardSMF28 fibres are not very sensitive to transverse effects. Anotherquality of optical fibre (OF) was therefore tested. The fibres em-ployed here were highly birefringent fibres (HiBi fibres). TheseOFs are manufactured so that the optical guide is characterizedby the two principal polarization axes. When written on such a fi-bre, an FBG is able to reflect two wavelength spectra because theindex profiles differ from one polarization axis to the other. Thanks

to these characteristics, these fibres are particularly sensitive totransverse effects. E. Udd exploited these characteristics to showthat an FBG written on a HiBi optical fibre becomes a very effectivemulti-axial strain sensor [8]. For advantageous use of this opticalfibre, specific tools are necessary to complement the other opticalmeasuring apparatus. A polarizer and a polarization controller areneeded. In the measurement loop, these components are placedbetween the source and the circulator [1]. Thus, in order to in-crease our chances of observing transverse effects at the time ofthe curing process, the mother plate containing the technologicalspecimen studied here was instrumented with such an OF. TheFBG written on the HiBi fibre was embedded in ply no. 3 of themother plate, in a position equivalent to that of the standard fibreFBG (Fig. 8a). More precisely, both FBGs were placed symmetricallydistant from the longitudinal axis of the mother plate such that,theoretically, they were subjected to the same curing stress andstrain conditions.

Fig. 8. (a) Schematic plan view of the instrumented mother plate and (b) photograph of the instrumentation phase during the pre-preg lay up of the mother plate.

1538 M. Mulle et al. / Composites: Part A 40 (2009) 1534–1544

3.3. Oven heating tests

After the autoclave cure analysis, the specimen was subjected toa thermo-mechanical analysis. The objective here was to study thestrain distribution pattern of thermal origin through-the-thicknessof the specimen during a thermal loading procedure carried out inan oven. The analysis conditions were less complex than those ofthe curing process since the specimen was subjected only to thetemperature parameter (without application of pressure or vac-uum or chemical transformation etc.). The temperature cycle cor-responded wholly to that of the curing cycle except for the180 �C dwell, which was limited to only 30 minutes instead of120 minutes. The most important was to replace the specimen un-der thermal conditions of its manufacture (i.e. 180 �C) and thatprolongation of these conditions would not yield additional infor-mation. The cooling phase differed from that of the autoclave pro-cess as no cooling control was available except by ventilation. Asthis was not sufficient, the oven was opened to accelerate coolingto the extent possible.

3.4. Stress-free temperature

Theoretically, the temperature at which the laminate is free ofstresses is usually considered for the determination of residualstrain and stresses. Thus, before commencing measurements andanalysis of theoretical results, it is necessary to determine thistemperature. An experimental approach consists in submittingunsymmetrical laminates from the ambient temperature (at whichtheir shape is curved as a consequence of the curing process) to thetemperature at which their shape becomes flat. This latter temper-ature is then regarded as the temperature at which the laminate isfree of stress. However, despite many studies and frequent discus-sions on the subject no agreement on this point has been reachedso far [18–23]. Certain authors found the stress-free temperatureto be lower than the curing dwell temperature, while others con-sider it to be higher. The former supports a cure-residual stress(CRS) relaxation either under viscoelastic effects related to coolingor by moisture absorption of the matrix, which also has a liberatingeffect on CRS. When this temperature is found to be higher thanthat of the curing dwell, the authors suggest the presence of chem-ical origin CRS which supplements the thermal origin CRS. Further,as explained in part I of this work [1], the tool/laminate interac-

tions are also considered to be at the origin of CRS. The effects ofthese interactions are believed to be trapped in the material andare hence likely to modify the total level of CRS [19].

Still, the most common approach on this subject stems from thehypothesis that, at the curing temperature, laminates are stress-free. In the context of the present study we decided to rely on thishypothesis. Thermal residual stress and strain will be treated for atemperature variation of �155 �C, corresponding to the differencebetween the curing dwell temperature (of 180 �C) and the roomtemperature (of 25 �C).

4. Numerical model

The technological specimen was modeled by the finite elementmethod (FEM). The meshing is carried out with 3-D quadratic re-duced elements. It must be pointed out that each composite plywas represented by one layer of elements. The mesh (especiallythe mesh of the ply drop-off zone) and the boundary conditionsare depicted in Fig. 8. It can be seen that all the ends of the pliesconstituting the resin rich pockets of the patch were modeled withpentaedric elements. The thermal test simulation was governed bya thermo-linear elastic behaviour law and the temperature duringthe cooling (DT = �155 �C) was assumed to change uniformlyacross the entire structure. Details are shown in Fig. 9. The associ-ated material properties were determined through elementarytests using specimens cured together with the technological spec-imen (Tables 1 and 2). In particular, thermal expansion coefficientswere obtained through a series of thermo-mechanical analyses(TMA) of a [08] UD sample. The mean ply thickness of the techno-logical specimen was measured by microscopy observations. As forthermoplastic nodules, their distribution (as revealed by the SEMobservations reproduced in Fig. 4) being highly random their ma-trix was considered homogeneous.

As stated above, CRS calculations were carried out treating thetemperature variation applied as DT = �155 �C.

5. Results and discussions

5.1. Spectrum analysis of highly birefringent (HiBi) FBG

Fig. 10 depicts the evolution of the wavelength spectra reflectedby an FBG written on HiBi fibre. Fig. 10a relates to the first half of

3D quadratic element 35584 elements with 483159 dof One element per ply

Mesh of the ply drop off zone

x

y z

300 mm

100 mm

88 mm

Ply thickness: 0.275 mm

30 mm

7.7

5.5

Uz=0

Uz=0

Ux=Uy=0

ΔT=-155°C

T700/M21: 50/40/10 T700/M21: 25/50/25 T700/M21: 50/40/10 Matrix M21

Fig. 9. FEM mesh characteristics of the technological specimen.

Table 1Mechanical and thermal properties of HexPly M21 T700GC.

Young’s modulus (GPa) Shear modulus (GPa) Poisson’s ratio CTE (10�6 �C�1) Ply thick (mm)

E11 E22 G12 m12 m21 m13 a1 a2 a3 e

118 8.4 3.7 0.32 0.023 0.51 �0.23 38.7 38.7 0.275

Table 2Mechanical and thermal properties of matrix M21.

Young’s modulus (GPa) Poisson’s ratio CTE (10�6 �C�1)

E11 = 3.3 m = 0.42 a = �65.2

M. Mulle et al. / Composites: Part A 40 (2009) 1534–1544 1539

the autoclave process. One can observe a series of spectra that cor-responds to various stages of this phase. At the beginning of thecure the spectrum showed a peak which was about to separate.At this stage, no transverse pressure was exerted on the FBG exceptthat arising from the action of the vacuum. We treated this spec-trum as the referent. The spectra that followed were collected afterpressurization at the 7 bar. No change in shape occurred, evenwhen the temperature progressed towards the isothermal curingdwell.

Fig. 10b presents the series of the spectra recorded from thecommencement of the isothermal stage until the composite lami-nate cooled down completely. On tracing the evolution of the spec-tral shape, one observes the presence of side lobes and a wideningof the spectral base which indicate the evolution of the initialstress field. Its origin is not easily identified. In addition, the prin-cipal peaks did not separate enough to reveal the presence of sig-nificant transverse strains.

This experiment may appear futile, because we could not quan-tify these deformations. On the contrary, it endorses the credit as-signed to the wavelength shift measurements (obtained by thesi425 unit) to transduce axial strains. If the shape of a spectrumparticularly sensitive FBG to transverse effects remains unaffected,then it will be even less so for a standard FBG. Thus, the Bragg peakshift curve can be interpreted with confidence.

5.2. Autoclave cure-induced strain assessment

The changes occurring in the longitudinal strain during the curecycle are presented in Fig. 11 in various plies of the reinforcedzone. During the temperature rise phase, the curve takes relativelybroken forms. During this phase, the resin viscosity is low and thesensors are primarily subjected to the increase of pressure and vac-uum together with the matrix flow. As was shown in part I of thiswork [1], the onset of gelation for the HexPly M21 T700 occurredclose to the beginning of the curing dwell. From this point on-wards, we considered the load transfer from matrix to fibre to beeffective [1]. The OFs and their FBG sensors detect the axial strainsthe composite undergoes. All through the duration of the isother-mal dwell the outlines of the curves of the four FBGs show a gentledecrease. This indicates that the material undergoes a small con-traction in the course of the isothermal dwell.

At the beginning of the cooling phase, strains progress negativelyuntil around 270 min into the cure cycle. It may be noted thatchanges in the strain vary from one ply to another. For ply 3 it issteeper than for plies10 and 19 and ultimately, much steeper thanfor ply 26. Apparent coefficients of thermal expansion during coolinghave been calculated during this phase, between 180 �C and 85 �C.We obtained aply26 = 3.2 � 10�6 �C�1, aply10 = 2.2410�6 �C�1,aply19 =1.93 � 10�6 �C�1, and aply26 = 1.09 � 10�6 �C�1. Observations arereported in Fig. 12. We observed a non-linear variation that was as-sumed to reveal the influence of the embedded reinforcing patch. Asregards amplitudes, they display an increase in the proximity of thealuminium mould. They are three times higher for the lower pliesthan for the upper ones. Although the highest value is a long wayfrom that of the aluminium CTE (around 23 � 10�6 �C�1) it is clear

Fig. 11. Process-induced strain changes measured by FBGs through the thickness of the reinforced zone of the technological specimen during the autoclave cure cycle.

Fig. 10. Wavelength spectrum changes reflected by a HiBi optical FBG during the first half (a) and the second half (b) of the cure cycle.

1540 M. Mulle et al. / Composites: Part A 40 (2009) 1534–1544

that there is a strong mould–laminate interaction, i.e. the contrac-tion of the laminate is amplified in the proximity of the metal mould[1].

Fig. 11 shows a lack of acquisition before the end of the cure cy-cle. This was assumed to be due to an appreciable increase in con-traction not only of the laminate but also of the vacuum baggingproducts that generates a multitude of micro-bends in the OFs.This results in excessive attenuation of the optical signal and even-tually in a complete loss. These optical signals were recovered oncethe laminate was released from the mould and the vacuum bag-ging products. Residual cure strains were estimated after the spec-imen was cut from the mother plate. The results obtained arereported in Table 3. A non-linear gradient through the thicknesswas noted. It is believed that this gradient is not linear either be-

cause of the shape of the specimen or the patch incorporated inthe ply stack. The tool/part interactions which took place at thetime of the cooling phase must also be considered as a factorresponsible for these deformations [24–27]. This information isimportant for acquiring a better understanding of the materialcharacteristics because it reveals a tendency for variability of ther-mal properties through the thickness of the technological speci-men. It highlights the importance of characterizing the materialwithin the structure.

5.3. Thermal strain assessment during heating tests

The variations of strain in plies 3, 10, 19, and 26 occurring overthe temperature cycle are presented in Fig. 13. They appear very

Fig. 12. Apparent longitudinal coefficients of thermal expansion during the cooling phase at various depths in the reinforced zone of the technological specimen.

Table 3Summary of thermally induced strain values of four thermal tests conducted in anoven; respective comparisons with the autoclave cure-residual strains are shown.

No. Thermal cycle in oven; DT = �155 �C Strain (lm/m)

Ply 03 Ply 10 Ply 19 Ply 26

1 14/03/06 �200 �185 �282 �2722 03/05/06 �210 �195 �275 �2453 08/06/06 �217 �186 �290 �2634 20/06/06 �220 �190 �274 �274

Average �212 �189 �280 �264Standard deviation 8 4 6 11

Autoclave cure-residualstrain 6/01/06; DT = �155 �C

�190 �145 �255 �255

M. Mulle et al. / Composites: Part A 40 (2009) 1534–1544 1541

similar to the temperature change. It is noted however that straindoes not vary proportionally depending upon the position inthrough-the-thickness of the reinforced zone. Once the isothermaldwell is reached, one observes various strain levels which resemblethose observed during the curing process. The cooling phase corre-sponds to the cooling capacity of the oven as explained in Section3.2. However, the difference in strain levels between those at thecuring dwell and at room temperature (DT = �155 �C) is ofsignificance.

Four identical thermal tests were carried out in order to collectobservations and thus verify the consistency of measurements.These tests were separated by several weeks, even several months.The thermally induced residual strain in plies 3, 10, 19, and 26were estimated and are reproduced in Table 3. For each test, a gra-dient of in-plane strain was noted through the thickness of thespecimen. All the thermal tests exhibited very similar exx strain dis-tribution patterns. They are also similar to those obtained duringthe autoclave cure process (Table 3) showing a non-linear distribu-tion. Thus, since this procedure is reproducible, we may say thatthere is a typical- and permanent distribution. This means thermalelastic properties within the structure exhibit variability. Theseproperties are accentuated during the curing cycle and are fixedin the material.

However, the strain amplitudes of the autoclave cure and heat-ing tests differ by approximately 10%. Environmental effects suchas moisture absorption or other ageing parameters are certainlyresponsible for these discrepancies. It is obvious from Fig. 13 thatat the end of the cooling phase, the strain curves reached lower lev-els than those at the initial state. Although the differences are not

very significant (approximately�10 lm/m) they may be attributedto relaxation of the residual stress. Relaxation is also observed overtime. The evolution of Bragg wavelength for each FBG was trackedover a 6-month period commencing from the state of the specimenimmediately after the autoclave cure. These observations are pre-sented in Fig. 14. Globally, a slight reduction in wavelength isvisible.

A second assumption concerning the origin of this effect is pro-posed. It is based on the propensity of this composite to absorbmoisture, a characteristic that was highlighted by Lévèque et al.[28]. It is supposed that at the end of the thermal test, the speci-men loses sufficient moisture so that its state of stress differs fromthat at the commencement of the test.

A complementary test was undertaken to evaluate the watercontent in the material of the specimens and to measure the influ-ence of temperature on desorption. The test consisted of thermo-gravimetric analysis (TGA) of a sample of approximate dimensions4 � 4 � 5 mm in a freshly cut technological specimen (the age ofthe specimen itself was about 6 months). Fig. 15 depicts the lossof mass of the sample through a temperature cycle. This corre-sponds to the thermal tests except that the cycle terminated atthe end of the 180 �C isothermal dwell, and natural cooling wasemployed. The operation was repeated three times consecutively.

The resulting loss-of-mass curves are shown in Fig. 15. Oneclearly distinguishes the significant losses during the first test,whereas they are rather insignificant during the following tests.The first test revealed the non-negligible presence of moisture inthe sample that usually evaporates. In the Hexply M21 matrix, thisbehaviour manifests perhaps due to the presence of 30% of thermo-plastic, which is known to be hydrophilic.

As for the technological specimen, it is undeniable that it under-goes moisture desorption during thermal tests, hence its thermalelastic behaviour is accompanied by relaxation effects.

5.4. Comparison with FEM simulation

Because of the presence of ply drops the specimen has an asym-metric geometry. Cure-induced stresses are expected to induce acertain amount of bending. Indeed, the numerical results showeda deflection of 15 lm. To compare this result to that obtainedexperimentally the specimen was examined on a surface table.When it was laid on the side that was in contact with the mouldplate during the cure a slight curvature was detected. Using shims

Fig. 14. Bragg wavelengths at T = 25 �C of FBGs situated in the reinforced zone of the technological specimen.

Fig. 13. Thermally induced strain changes measured by FBGs through the thickness of the reinforced zone of the technological specimen during the oven reheating cycle.

1542 M. Mulle et al. / Composites: Part A 40 (2009) 1534–1544

at the longitudinal centre of the beam, a maximum gap of 20 lmbetween the surface table and the technological specimen wasmeasured. Results obtained by numerical simulation and experi-ment results were very similar. The influence of this curvatureon cure-induced strains was estimated experimentally. With thisin view, the specimen was pressed down on the surface table untilthe gap between the two completely disappeared while the wave-length variations in all the four FBGs were measured. In terms ofstrain, two positive shifts of 52 and 31 lm/m were obtained,respectively, for plies 3 and 10 and two negative shifts of �25and �60 lm/m were obtained, respectively, for plies 19 and 26(cf. Table 4). As the influence of this curvature is not negligible,for comparison of values of strain obtained by experiment and bynumerical simulation it is important to take into account a similardeflection state for the beam.

These findings are presented in Fig. 16 and Table 4. The valuesobtained by numerical simulation were compared to those ob-tained for autoclave cure-residual strains and thermally inducedstrains. It was first noted that the distributions resulting from theexperimental tests differed from those obtained by numerical sim-ulation. Values obtained by numerical simulation exhibit a quasi-linear strain gradient. It is evident that such calculation does not

take into account the variation of the thermal elastic propertiessince only one longitudinal expansion coefficient was incorporatedinto the numerical model. Consequently, its results revealed only avery small strain gradient which may be attributed to a structuraleffect (non-linear geometry of the specimen).

As for the residual strain amplitudes, those relating to exteriorplies (plies 3 and 26) presented a rather satisfactory match (dis-crepancy lying between 5 and 21%). For ply 10, a non-negligiblediscrepancy was noted (40%) but still, results are of the same orderof magnitude.

6. Conclusions

This paper aimed to the assess through-the-thickness process-induced strains in a carbon–epoxy technological specimen. Forthis purpose, several optical FBG sensors were embedded in anddistributed all through-the-thickness of the reinforced zone ofthe specimen. Induced strains were measured during autoclavecure and reheating tests of similar temperature cycles. Both seriesof values showed identical distribution of residual strain. A non-linear gradient was observed which revealed variability in ther-

Fig. 15. Thermogravimetric analysis (TGA) for three consecutive thermal tests on a technological specimen sample; loss-of-mass measurements.

Table 4Through the thickness exx strain values obtained by numerical simulation and measurements recorded during autoclave cure and reheating procedure for the technologicaspecimen.

Autoclave cure-induced strains Oven heating-induced strains Numerical simulation

With curvatureexx (10�6)

Variation/num.sim. (%)

Without curvatureexx (10�6)

With curvatureexx (10�6)

Variation/num.sim. (%)

Without curvatureexx (10�6)

With curvatureexx (10�6)

Pli no. 26 �255 5 �315 �264 9 �324 �243Pli no. 19 �255 5 �280 �280 16 �305 �242Pli no. 10 �145 �40 �114 �189 �22 �158 �241Pli no. 3 �190 �21 �138 �212 �12 �160 �241

Fig. 16. Through the thickness exx in-plane strain distribution patterns obtained by numerical simulation, and measurements recorded during autoclave cure and reheatingprocedure for the technological specimen.

M. Mulle et al. / Composites: Part A 40 (2009) 1534–1544 1543

mo-mechanical properties through the thickness of the structure.Both mould/laminate interactions and structural effects (presenceof a reinforcement patch) are considered influent. The amplitudesof autoclave cure- and heating-induced strain values showed a

l

10% discrepancy. This may be due to humidity and/or relaxationeffects.

A numerical model (FEM) was developed to simulate thereheating tests. Due to the asymmetric geometry of the specimen

1544 M. Mulle et al. / Composites: Part A 40 (2009) 1534–1544

and cure-induced stresses a certain amount of bending was ob-served experimentally as well as numerically. The in-plane straindistribution is linear and shows almost no gradient. It is evidentthat incorporating only one thermal coefficient in the model is thusnot sufficient to reproduce behaviour of the material.

The analysis of a technological specimen instrumented withFBG allowed estimation of the initial strain state of the material,revealing variability in its properties. It stresses the importanceof characterizing the material with respect to the structure within.

Acknowledgments

The authors wish to thank the DGA (Direction Générale del’Armement) for its financial support through the upstream pro-gram AMERICO and the Hexcel Composites Company for providingus with HexPly� pre-preg material.

References

[1] Mulle M, Collombet F, Olivier P, Grunevald Y-H. Assessment of cure residualstrains through the thickness of carbon–epoxy laminates using FBGs, part I:elementary specimen. Compos Part A 2009;40:94–104.

[2] Grunevald Y-H, Collombet F, Mulle M. Integrated approach for designingcomposite structures: a new industrial challenge. In: Proceedings of 34th solidmechanics conference, Zakopane (Poland); 2002.

[3] Grunevald Y-H, Collombet F, Mulle M, Ferdinand P. Global approach andmultiscale instrumentation of composite structures in the field oftransportation. In: Proceedings of 8th Japan international SAMPEsymposium, Tokyo, vol. 2; 2003. p. 665–75.

[4] Laurin F, Carrère N, Maire J-F. A multiscale progressive failure approach forcomposite laminates based on thermodynamical viscoelastic and damagemodels. Compos Part A 2007;28:198–209.

[5] Mulle M, Collombet F, Périé JN, et al. Instrumented technological specimen(ITS): curing process and mechanical characterisation using embedded fibergratings. In: Proceedings of ECCM11 (11th European conference oncomposite materials), 31 May–3 June 2004, Rhodos (Greece), CDROMpaper#C137; 2004.

[6] Collombet F, Mulle M, Grunevald Y-H, et Zitoune R. Contribution of embeddedoptical fiber with Bragg gratings in composite structures for test-simulationdialogue. Mech Adv Mater Struct 2006;13(5):429–39.

[7] Mulle M, Zitoune R, Collombet F, Robert L, Grunevald Y-H. Measuring strainsthrough the thickness of a composite structural specimen subjected tobending. Exp Mech 2008. doi:10.1007/s11340-008-9209-.

[8] Udd E, Schulz W, Seim J, Haugse E, Trego A, Jonhson P, et al. Multidimensionalstrain fields measurements using fiber optic grating sensor. In: Blue RoadResearch – 2000, SPIE proceedings on smart structures and materials, vol.3986; 2000. p. 254–72.

[9] Güemes JA, Menéndez JM. Response of Bragg grating fiber-optic sensors whenembedded in composite laminates. Compos Sci Technol 2002;62(7–8):959–66.

[10] Okabe Y, Yashiro S, Tsuji R, Mizutani T, Takeda N. Effect of thermal residualstress on the reflection spectrum from fiber Bragg grating sensor embedded inCFRP laminates. Compos Part A 2002;33:991–9.

[11] Mulle M, Zitoune R, Collombet F, Olivier P, Grunevald Y-H. Thermal expansionof carbon–epoxy laminates measured with embedded FBGS – comparison withother experimental techniques and numerical simulation. Compos Part A2007;38:1414–24.

[12] Meirinhos G, Rocker J, Cabanac J-P, Barrau J-J. Tapered laminates under staticand fatigue tension loading. Compos Sci Technol 2002;62:597–603.

[13] Cui W, Wisnom MR, Jones M. New model to predict static strength of taperedlaminates. Composites 1995;2.6:141–6.

[14] Gay D, Hoa SV, Tsai SW. Composite materials: design and applications. CRCPress; 2003. ISBN: 1.58716.084.

[15] Kuang KSC, Kenny R, Whelan MP, Cantwell WJ, Chalker PR. Embedded fibreBragg grating sensors in advanced composite materials. Compos Sci Technol2001;61:1379–87.

[16] Ferdinand P. Capteurs à fibres optiques à réseaux de Bragg. Techniques del’Ingénieur, traité Mesures et Contrôle, R 6 735-1; 1999 [in French].

[17] Rao Y-J. In-fibre Bragg grating sensors. Measure Sci Technol1997;8:355–75.

[18] Schapery RA. Inelastic behaviour of composite materials. In: Herakovicheditor. Proceedings of the 1st conference of the American society formechanical engineers on composite materials, New York, C.T. 127;

1975.[19] Hahn HT. Residual stresses in polymer matrix composite laminates. J Compos

Mater 1976;10:266–78.[20] Jones FR, Mulheron M, Bailey JE. Generation of thermal strains in GRP. Part 2:

The origin of thermal strains in polyester cross-ply laminates. J Mater Sci1983;18:1533–9.

[21] Tsai SW. Theory of composites design. Think composites; 1992.[22] Kim KS, Hahn HT. Residual stresses development during processing of

graphite/epoxy composites. Compos Sci Technol 1989;36:121–31.[23] Crasto AS, Kim RY. On the determination of residual stresses in

fiber-reinforced thermoset composites. J Reinf Plast Compos 1993;12:545–58.

[24] Tarsha-Kurdi KE, Olivier P. Thermoviscoelastic analysis of residual curingstresses and the influence of autoclave pressure on these stresses in carbon/epoxy laminate. Compos Sci Technol 2002;62:559–65.

[25] Potter KD, Campbell M, Wisnom MR. The generation of geometricaldeformations due to tool/part interaction in the manufacture of compositecomponents. Compos Part A 2005;36:301–8.

[26] De Oliveira R, Lavanchy S, Chatton R, Costantini D, Michaud V, Salathé J-R, et al.Experimental investigation of the effect of the mould thermal expansion onthe development of internal stresses during carbon fibre compositeprocessing. Compos Part A 2008;39:1083–90.

[27] Garstka T, Ersoy N, Potter KD, Wisnom MR. In situ measurements of through-the-thickness strains during processing of AS4/8552 composite. Compos Part A2007;38(12):2517–26.

[28] Lévèque D, Mavel A, Nunez P. Investigation of hygrothermal residual stressesin composite laminates and their consequences on the damage onset. In:Proceedings of JNC 15, Marseille, France; 2007. p. 125–32.