M Chaussumier
email: michel.chaussumier@insa-toulose.fr.

M Shahzad

C Mabru

R Chieragatti

F Rezaï-Aria

Fatigue

Keywords: anodizing, pits, fatigue, aluminium alloy, short crack, long crack

It has been shown that decrease of the fatigue life of aluminium alloys treated with anodization can be explained by the degradation of surface condition due to pickling. In order to predict fatigue life of anodized aluminium alloys, a multi-site crack growth model is developed by considering the pickling pits sites as initial flaws from which fatigue cracks develop. A map of the pickled surface is built from topography measurement with a contact profilometer. Then the pits are detected and their sizes are defined (depth, length and width). At the beginning of the calculation, a short crack growth law is used for crack having depths less than grain size. Then Paris long crack growth law is used. The coalescence of cracks is considered when their lengths increased by its crack tip plastic zone are large enough to interact with other neighbouring short cracks. The fatigue life is calculated for K max achieving 70 % K IC .

Introduction

Aluminum alloys 2xxx and 7xxx are extensively used in aeronautical industry due to their superior strength to weight ratios. As the natural oxide layer thickness is not sufficient to protect structure when subjected to severe environmental conditions, aluminium alloys are often anodized by electrolytic process that produces controlled columnar growth of amorphous aluminium oxide on the surface [1]. Despite the benefits obtained for wear and corrosion resistance, it has been observed an important decrease of the fatigue resistance [START_REF] Cree | Effect of anodized coatings on fatigue cracks rates of aluminium alloy[END_REF][START_REF] Kermanidis | Fatigue damage tolerance behaviour of coroded 2024 T351 aircraft aluminium alloy[END_REF][START_REF] Lonyuk | The effect of oxide coatings on fatigue properties of 7475-T6 aluminium alloy[END_REF][START_REF] Sadeler | Effect of a commercial hard anodizing on the property of a 2024-T6 aluminium alloy[END_REF]. This decrease seems to depend on stress level. Generally this reduction in fatigue life is attributed to the brittle as well as porous nature of oxide layer and the tensile residual stress induced during anodizing process [START_REF] Camargo | Influence of anodization on the fatigue strength of 7050-T7451 aluminium alloy[END_REF][START_REF] Cyrik | Effect of anodic oxidation on the fatigue performance of 7075-T6 alloy[END_REF]. It can also be attributed to the surface degradation by pickling process before anodizing. The pickling pits have been identified to accelerate crack initiation [START_REF] Rokhlin | Effect of pitting on fatigue crack initiation and fatigue life[END_REF][START_REF] Pao | On fatigue crack initiation from corrosion pits in 7075-T7351 aluminium alloy[END_REF][START_REF] Dolley | The effect of pitting corrosion on fatigue life[END_REF]. In addition, it has been shown that this effect depends upon initial machining roughness: the more it is, the less the effect on fatigue life is [START_REF] Chaussumier | Influence du décapage sur la tenue en fatigue de l[END_REF].

The aim of this study is to compare several multi-crack propagation models for pickled specimens to experimental test results. These pits were initially identified on pickled plate specimens with topography measurement which allowed to define the number, the size (depth, length) of the pits and their respective position. Four approaches for predicting fatigue life have been used. In all approaches, the relative position of the pits has been considered. These approaches distinguish one from another by the initial flaws size, the applied stress level at the deepest point of the pit calculated with finite elements models built from topography measurement, and the propagation law. In the last approach, we used the local stress concentration at the deepest point of the pit. This kind of approach has been already proposed for a two-dimensional model [START_REF] As | Fatigue life prediction of machined components using finite element analysis of surface topography[END_REF][START_REF] Suraratchaï | Modelling the influence of machined surface roughness on the fatigue life of aluminium alloy[END_REF]. During their propagation, cracks emanating from pits can coalesce to form new crack front that then continue to propagate until cracks reach a critical size corresponding to 0.7 K IC .

Experimental Procedures

Materials

The material investigated in this work is 7010-T7451 aluminium alloy whose chemical composition is given in Tab. 1. It was provided in rolled plate form of 70 mm thickness. Microstructural investigations revealed the presence of unrecrystallized and recrystallized grains. The latter are highly elongated in the rolling direction as shown in Fig. 1. Grain size is about 350 μm in the rolling direction and about 150 μm and 60 μm in TC and TL directions respectively. Only two types of inclusions were mainly found in this material (Mg 2 Si, Al 7 Cu 2 Fe) that were generally located in recrystallized grains. The average size of these particles varied between 8 to 10 µm. Mechanical properties of this alloy in the rolling direction are; yield strength 464MPa, ultimate tensile strength 526 MPa, Young's modulus 70.7 GPa and elongation 9.8%.

Fatigue specimen preparation

Cylindrical test specimens were prepared in such a way that maximum load is applied perpendicular to the rolling direction, as shown in Fig. 2. They have been machined by turning without using lubricant. Turning was performed on 2 axes numerical lathe RAMO RTN20. Two types of machined surfaces (R a =0.6µm and R a =3.2µm) were produced. Machining conditions are shown in Tab.

Surface treatments

For each initial roughness, specimens were categorized in 4 groups; the first group, with no treatment was used to build reference fatigue curve; the second group had specimens which undergone degreasing, the third group specimens were pickled and finally fourth one were anodized.

Specimens degreasing was carried out in aqueous solution of sodium tripolyphosphate Na 2 P 3 O 4 and Borax Na 2 B 4 O 7 at 60°C for 10-30 minutes followed by demineralised water rinsing. Pickling was done in aqueous solution of H 2 SO 4 acid and anhydride chrome CrO 3 at 60°C between 1-10 minutes followed by rinsing. Chromic acid anodization was accomplished in anhydride chrome CrO 3 solution at 45°C under 50 volts for 55 minutes. The average thickness of oxide layer produced by the process was measured to be about 3µm

Surface measurement

Comparable surface roughness to that produced by turning of the cylindrical rotating bending specimens has been reproduced on plate specimens in order to better measure surface topography after pickling process. These surfaces have been obtained machined by a shaper with the same parameters as turning or for preparation of surface treated, especially in pickling process. Then surface topography measurements have been done using Mahr PKG120 profilometer. It is a conical diamond stylus instrument (90°-2 μm radius) that can give conventional roughness parameters with horizontal and vertical resolutions of 0.5 µm and 0.1µm respectively. Several 2×2 mm 2 areas have been measured in order to lead a statistical analysis of the pickling pits. Specimens were placed on profilometer table such as the measurement directions correspond with applied stress direction. The resolution in this direction and in the transverse one is chosen equal to 5 µm. Pits characterisation is realised with a specific program. From the surface topography data, a plane orientation correction is made using the least square method. In this step, pits are neglected in face of the selected area and normal distribution factor is calculated. Then, surface profile, calculated from machining conditions, is subtracted from the real surface. The depth of the pits is defined from this new surface under the following condition:

Ra z z ij ij < -ˆ(1)
where represents the measured depth, the theoretical depth considering machining conditions and Ra represents the arithmetic average roughness.

ij z ˆij z
All adjacent points matching the above condition are assumed to belong to the same pit. By this way, each pit can be characterised through its maximal depth, its length (along transverse direction) and width (along applied stress direction) but also its position on the surface. Figure 3 illustrates this characterization method; pits are colored in red. Statistical treatments of this characterization allow establishing depth and length distribution graphs of pits.

Modeling

Finite elements models

From corrected surface topography, 3D finite elements models using the FE code Abaqus have been developed for each measured area. Linear bricks elements with 8 nodes have been used. The element size corresponds to the resolution used for topography measurement in each direction of the surface. The local surface crater depths correspond exactly to the measured depths. The thickness of the meshed volume, the number of elements trough the thickness and the ratio between element thickness are then defined. Displacements in the direction of applied stress are applied on the extreme transverse section of the finite element model. These models allow calculating for each pit a local mean stress, denoted . Then, for each pit, a stress concentration factor

Crack propagation models

In order to predict the fatigue life, an analytic model has been developed based on linear fracture mechanics with local plasticity correction. In this model, each pickling pit is considered as a semi-elliptical micro crack with a shape factor c/a. Four approaches have been developed with considering the crack coalescence. The crack propagations on surface and inward the width of the specimen are calculated simultaneously. However the initial size of cracks and the fatigue crack growth law are different for each approach.

Modeling of crack coalescence

Crack propagation rate of a crack is influenced by the proximity of other cracks [START_REF] Debartolo | Effects of constituent particle clusters on fatigue behaviour of 2024-T3 aluminium alloy[END_REF]. During propagation, cracks interact with each other even if the crack tips do not touch each other. At this state of the development of the model, two cracks become a single crack without being necessary a semi-elliptical crack. During the coalescence step, the single cracks transform to semi-elliptical cracks. Then a single crack propagates until it interacts with another crack and so on. We consider the coalescence is occurring when the crack tip plastic zones of two cracks touching each other. This coalescence condition depends on the relative position of the two cracks. For example, if one of the cracks is entirely on one side of the other one, the condition of coalescence can be expressed as follow:

j i ij zp zp d + ≤ 2 max ) 1 ⋅ = Rp K zp π (3) ( ) ( 2 2 j j i i j i i ij c y c y x x d - - + + - =

Propagation models

The first approach consists in considering only the multicracks and coalescence of long fatigue cracks emanating from the pits. The size of initial flaw is calculated using Kitagawa relation between stress intensity factor threshold th K Δ and fatigue limit D σ ( )

2 0 0 1 1 ⋅ - Δ ⋅ = D th R K a σ π (4)
In this first approach, it is considered that the initial flaw depth is grater than the pit depths. In this condition the local stress at the deepest point of the crack front is considered to be equal to nominal stress. Therefore local stress raising is not taken into account and thus the local stress factor is 1. On the contrary, the stress concentration factor at the surface is considered by using Kt local method [START_REF] Chieragatti | Procédés de caractérisation de la tenue en fatigue d'une pièce à partir de son profil de surface[END_REF]. The propagation is calculated using a long crack growth rate law. In order to predict more realistic the life, the crack initiation life (N i ) and propagation life (N p ) are considered. The S-Ni curves machined specimens [START_REF] Santus | Physically short crack propagation in metals during high cycle fatigue[END_REF] are obtained by:

p R i N N N - = (5)
where N p is calculated using the long crack growth law.

x

y crack i crack j y i y j zp i x i x j c i c j d ij zp j
In the second approach, the size of the pits is considered but the stress concentration factor is still neglected. Alike the first approach, the stress concentration factor on surface with a particular roughness due to machining, is considered and again the propagation law is calculated as the approach 1.

In the third approach, the size of the pits is set as initial flaw size and the stress concentration factor at the deepest point of the pits is also considered. Two stages of propagation are considered. In the first stage, the stress concentration factor at the maximum depth of pit controls the crack growth rate. This stress concentration factor has been calculated using previously mentioned FEM model. In the second stage the cracks shape factor c/a, is taken into account for the crack growth. This consideration is intuitively in according with the experimental evidences of small cracks propagation. Several studies on aluminium alloys have shown that micro cracks could propagate even if the stress intensity factor range K Δ is less than the threshold stress intensity factor range th K Δ usefully defined for long cracks [START_REF] Pearson | Initiation of fatigue cracks in commercial aluminium alloys and the subsequent propagation of very short cracks[END_REF][START_REF] Lankford | The influence of microstructure on the growth of small fatigue cracks[END_REF]. It is often observed that short crack growth rate is greater than the long crack for given range of stress intensity factor less than th K Δ . Furthermore, for some small cracks the propagation can stop, for others, the crack growth rate first decrease and the increase to join long crack growth rate law. In this model, the transition between the two region of crack propagation is fixed to be th K Δ . For stress intensity range factor less than th K Δ , the stress concentration factor at the deepest point of the pit is consider while over th K Δ the stress is equal to nominal stress.

In the last approach, a specific small fatigue crack growth rate law is associated to the first stage of crack growth. The small crack growth rate is calculated using th K Δ corresponding to a pit depth a using El Hadad's model [START_REF] Haddad | Prediction of non propagating cracks[END_REF] based on Kitagawa's theory.

a a a K K th a th + ⋅ Δ = Δ 2 0 , α (6)
where a 0 is given by Eq. 4.

α is the geometric factor at the deepest pits depth that is also depending on shape factor c/a. Then short crack growth rate is calculated by reporting stress intensity factor range threshold

m a th a K C C , Δ ⋅ = (7)
where C et m are long crack Paris' law parameters. For the present alloy investigated, these parameters are equal to 3.17 10 -11 and 3.41, respectively with th K Δ equal to 3.5 MPa m for a 0 about 200 μm [START_REF] Suraratchaï | Modelling the influence of machined surface roughness on the fatigue life of aluminium alloy[END_REF].

Several models have been proposed for the calculation of the geometric factor α. In the present model, Murakami's geometrical correction factor is chosen due to its simplicity [START_REF] Murakami | Stress intensity factor for interacting two equal semi-elliptical cracks in tension[END_REF]. Figure 6 illustrate the evolution of this factor for the deepest depth of the crack and for the tip point at the surface. For the depth, the curve has been extrapolated for shape factor value over 2 and was limited to 0.2 as it can be seen (see white triangular points). For the surface, the correction factor has been limited to 1 for shape factor over 4.5.

Correction factor for semi-ellipical crack alpha_depth= 0.

Results and discussion

Figure 7 illustrate fatigue life as a function of alternating stress for each group of test specimens. It can be observed an important decrease in fatigue life scattering between machined specimens and anodized specimens. This decrease depends on stress level. When the applied stress decreases this diminution becomes worth in accordance with literature. It can be also observed that fatigue life curves for pickled specimens and anodized specimens are much closer. That indicates the primacy of pickling on the global decrease of fatigue resistance of anodized specimens. All these results have been treated using Kt local method. It has been shown that results for machined specimens, degreased ones and pickled ones are set on the same S-N curve, Fig. 8. The fracture surface of specimens with different treatments were examined by the SEM to identify the crack origin sites. The examination revealed multi-site crack initiation for the pickled and anodized specimens. The pickling pits initiating the fatigue cracks were about 8µm deep.

The statistical distribution of surface (2×2 mm 2 ) topography depth is shown in Fig. 9a. 117 pickling pits have been characterized at the surface of low initial roughness. Most of them are of small depth with an average being about 6 μm. Only 2% of these pits are more than 10 μm depths. In the same manner, the pit's mean length is about 10 μm (see Fig. 9b) and the longest ones are about 30 μm depth. Shape factor is between 0.58 and 3.28 with a mean value equal to 1.3. For 86% of the pits, the shape factor is less than 2. In the same way, it is possible to define the statistical distribution of stress concentration coefficient. Mean stress concentration coefficient is about 2.24. Figure 10 illustrates the results of one of these FEM models. Figure 11 presents stress concentration distribution for a given simulated area. The following figures illustrate the prediction of fatigue life given by the four approach presented in this paper. In the first approach, all the pits are considered and their respective position too. But their size is not considered. The size of the initial flaws is equal to the size of the crack corresponding to threshold stress intensity factor range. The effect of multicracks on propagation can be demonstrated through the decrease of Np of a single pit (Fig. 12). As it can be seen, the predicted fatigue life is lower than experimental life of the pickled specimens. However, initiation step is not considered in this approach. When correction factor is taken into account with the initiation curve, the predicted fatigue life is more close to the experimental data. The very good agreement obtained with this first model allows concluding on the importance of multicrack considerations associated with coalescence phenomena.

Figure 13 illustrates the results obtained with the second approach. As it can be seen, predicted life is very much optimist in comparison to the experimental results. This can be easily explained by the fact that the initiation step is not well represented because the actual stress at the pit's tip is not considered and also probably due to using an improper short crack growth rate law.

The predictions given with the third model are also too optimists comparing to experimental results but are more closed as it can be seen in Fig. 14. In this model, the influence of stress concentration at the pits until size reach a same critical size and the use of a single long propagation crack growth rate law reduce the difference between predictions and results comparing to the second models In the last model, small crack growth rate laws depending on initial size of pits are considered. The influence of stress concentration until the size of the crack rich a transition size depending on initial size of the pits are considered too. The predictions given by this last model are in very good agreement with experimental results as it can be seen in Fig. 15. Agreement is better when stress level is high where propagation period in total life is predominant. But life predictions become conservative for low level stress level where the crack initiation is predominant in total life. Such trend traduces the difficulty to characterize small crack growth rate. This is an important aspect to be investigated in the future.

Conclusion

Rotating bending fatigue test have been realized on aluminum alloy 7010 after machining, degreasing, pickling and anodizing. It has been observed an important decrease of fatigue life between machined specimens and anodized ones. It was also observed an importance effect of pickling on this decrease. Cracks initiate on pickling pits, (assimilated tt small cracks or flaws) and multicracks was observed. In order to predict the fatigue life of pickled specimens, a multicrack propagation model considering the coalescence of cracks has been developed. The depth and length repartition of pits have been characterized via surface topography measurements and statistical treatments. Finite elements models have been developed based on these same surface topographies in order to calculate stress concentration coefficient at the pits. It is considered that the local stress controls small cracks growth rate. Several approaches have been developed in order to distinguish the respective influence of the multicracking, the local stress effects and the small crack growth rate laws. The importance of multicracking coupled with coalescence of cracks is emphasized. The stress concentration at the bottom of pits and the small fatigue crack growth rate law seems equally play an important role. One of the further investigations will concern the small fatigue crack growth rate law characterizations.

Fig. 2 .

 2 Fig. 2. Fatigue specimens position.

Fig. 3 .

 3 Fig. 3. Pickled surface topography for low roughness machining conditions.

Fig. 4 .

 4 Fig. 4. Finite element model constructed from surface topography.

Fig. 5 .

 5 Fig. 5. Coalescence criteria conventions.

Fig. 6 .

 6 Fig. 6. Murakami's geometrical correction factor as function of shape factor c/a.

Fig. 7 .Fig. 9 .

 79 Fig.7. Fatigue results for all the surface treatments and low roughness. Fig.8. Fatigue results using Ktlocal method for all surface treatments excepted anodizing and for the two initial surface roughness.

Fig. 10 .

 10 Fig. 10. Finite elements results. Fig. 11. Statistical distribution of stress concentration coefficient.

Fig. 12 .Fig. 14 .

 1214 Fig. 12. Fatigue life prediction given with model 1. Fig. 13. Fatigue life prediction given with model 2.

Table 1 .

 1 chemical composition of 7010 Aluminium alloy

	Element	Si	Fe	Cu	Mn	Mg	Cr	Zn	Zr	Ti
	Weight%	0.12	0.22	1.98	0.20	2.86	0.15	6.45	0.27	0.14

Fig. 1. Microstructure of 7010-T7451.

Table 2 .

 2 Specimens machining conditions

	Ra (μm)	Vc (m/min)	f (mm/tr)	p (mm)	ρ (mm)	lubrication
	0,6	180	0,1	0,5	0,8	none
	3,2	180	0,3	0,5	0,5	none


Anh Vu Duong
email: :anh-vu.duong@isae.fr

Samuel Rivallant

Jean-Jacques Barrau

Caroline Petiot

Benoît Malherbe

Influence of speed on the crushing behavior of composite plates

Keywords: Composite, crushing, energy absorption, speed, CFRP

This paper presents a study on crushing of CFRP composite laminated plates. The objective of this study is to analyze the influence of crushing speed on the rupture behavior and the energy absorption capability of composite materials.

Two materials are tested: a carbon-epoxy unidirectional prepreg, and a carbon-epoxy fabric. For both materials, three different plate configurations are tested (three different stacking sequences).

The crushing device is an improved plate crushing setup with stabilizing guides, which can be used both on a standard static traction/compression test machine, or on a drop tower. For static tests, a 20 mm/min velocity is applied to the specimen all along the 100mm crushing stroke. For dynamic tests, only the initial velocity can be imposed. Three initial speeds were chosen: 2, 5 and 9 m/s.

A high-speed camera is used during tests to visualize the global behavior of the plate, and the rupture modes in the crushing front. The load is also recorded, which enables to calculate the load/displacement curve of plates crushing and determinate the energy absorption capability.

The analysis of these results enables to discuss on the influence of crushing speed both on the rupture modes and, as a consequence, on the energy absorption capability.

INTRODUCTION

The use of laminated composite in vehicle structure requires a good understanding of their crash behaviour. But the range of crushing mode is very wide and strongly depends on numerous factors. Despite extensive experimental works done in the 90's [1:5], and more recently [START_REF] Ueda | Energy absorption of unidirectionam CFRP plate unde compression[END_REF], it is still very hard to predict the crash response of a complex structure.

A few works were done concerning dynamic crush tests, [START_REF] Lavoie | Dynamic crush tests of energy absorbing laminated composites plates[END_REF], but there are very few results concerning the influence of crushing speed on crushing behavior and energy absorption capability. Moreover, conclusions seem often contradictory (problems of dispersion, to few tests…)

The aim of this study is to obtain experimental results providing a better understanding of the speed influence in progressive crushing of composite laminated plates. The main experimental objective of this work is to perform crushing tests on two different materials, with three different stacking sequences for each material, at 4 different speeds (from quasi-static to 9 m/s). A second objective is also to compare energy absorption capability of laminates made of UD plies or fabrics, for the same equivalent modulus.

EXPERIMENTAL 2.1 Fixture design

The test fixture (figure 1) has been designed to improve plate crushing fixtures found in literature [START_REF] Guillon | Initiation and propagation mechanisms of progressive crushing in carbon-epoxy laminated plates[END_REF]. It is made of four vertical adjustable guides which avoid buckling of the specimen, and two horizontal adjustable guides. These ones are localised just beneath the vertical guides so that a gap exists between horizontal guides and the metallic base plate where the laminate crashes. Such a concept has been developed at the same time by Feraboli [9]. This avoid tearing of the plate observed in most of the plate crushing fixtures [START_REF] Daniel | The crush behaviour of carbon fibre angle-ply reinforcement and the effect of interlaminar shear strength on energy absorption capability[END_REF], and ensure that the boundary conditions are the same along the whole width of the plate, just above the crushing front. Two main uprights carry the guides without interference with the crush front. Visualisation holes have been made to let the crush front visible.

The thickness of the specimen can vary from 0 to 10mm. The gap (unsupported height between the base plate and the horizontal guides) can vary from 0 to 40 mm.

The laminate is fixed to a steel cylinder which is the interface with the static machine or the drop tower (dynamic tests). During crushing, plate and cylinder go down between the uprights and the vertical guides. At the bottom of the cylinder, a 120 kN Kistler piezoelectric sensor measures the crushing load. The small distance between the specimen and the sensor limits the mechanical filtering of the signal, allowing high precision in the dynamic crush load measurement. A high speed camera is used to have a real-time observation of the crushing front on the side of the plate. The filming speed is 20000 fps, with a 512x256 pixels resolution. The fixture is used both for quasi-static test on a screw-driven universal testing machine and for dynamic tests on a drop tower apparatus. The load is introduced through the top of the specimen with a steel cylinder. At the bottom of this cylinder, a piezoelectric sensor measures the crushing load. The small distance between the specimen and the sensor limits the mechanical filtering of the signal.

The use of a drop tower does not enable to have a constant speed during dynamic crushing : speed decreases as energy is absorbed. Depending on the test configurations (available weight on the drop tower, energy absorption capability of specimen, initial speed), speed can just slightly decrease, or reach 0 (in that case, the specimen is not totally crushed).

Design of experiment

Specimens are 160*60 mm flat plates. White graduations are drawn on the edge of the specimen each 5 mm. The trigger mechanism is a 45° chamfer machined at the bottom end of the plate. Each test is defined by two parameters:

-Crush speed: 20 mm/min for static tests, and three different impact speeds for drop tower tests: 2 m/s, 5 m/s, and 9 m/s.

-Laminate configuration: two materials were tested: 1). This enables to compare energy absorption efficiency for these two materials. The unsupported height is set to 20 mm for all tests. Each configuration and speed, two specimens are tested.

RESULTS and DISCUSSION

The numerous experiments made enabled to observe the different crush modes of composite laminated plates with different configurations and different speeds. Crush mechanisms are homogeneous through the width of the plate, thus the real-time pictures of the side of the crushed laminate (figure 4 and5) allow observation and understanding of crush modes. Following is shortly described some of the main modes observed. With all specimens tested, there isn't undamaged splaying mode. The crushing mode observed is mixed mode and fragmentation mode. In mixed crush mode, exterior plies on both sides bend and evacuate in splaying crush mode. The proportion of plies that bend on each side is variable. An obstacle, like a debris wedge, in the front separating two splaying arms is the reason why this mode appears. Fragmentation occurs when plies reach the metallic base at right angle and can not slip on the base. Failure appears continuously at macroscopic scale. Energy absorbed can be very high: up to 50 kJ/kg, depending on the number of fragmented plies. But fragmentation is not a stable crush mode, plies crush evolving from fragmentation to bending, leading to undamaged or fractured splaying crush mode.

Influence of speed

The results of these tests for different stacking sequences of composites plates Cytec and T700/M21 with different initial speeds are shown in figure 6 and figure 7. The dataset of the maximum loads Fmax during the tests for each sample is created. Fmax is the initial peak load (figure 3). Globally, this peak load increases with the speed of the initial velocity before impact. number of plies in fragmentation is not stable during the test (and not the same from one test to the other). In figure 7, the dataset of each stacking sequence of both materials with different initial velocity of the test is shown. It can be concluded that there is no visible influence of speed on the energy absorption on the two materials tested with the three stacking sequences.

Figure 7: SEA with different initial speeds

Comparison unidirectional (UD)/fabrics

The Cytec samples and T700/M21 samples have the same equivalent Young modulus for each stacking configuration (iso-rigidity). To compare UD to fabrics, as no visible influence of speed on the results of SEA was found, an average value of SEA was calculated for each material and stacking configuration taking into account the SEA of tests at all speed for a given stacking configuration. Comparison is shown on Figure 8.

The results show that the energy absorption capability of Cytec fabric is much higher (approximately twice) than T700/M21, for the three tested configuration. This can be explained by the fact that in Cytec samples, plies damage happens earlier than in T700/M21, leading to fragmentation whereas in T700/M21, delamination appears between plies, before fiber rupture, developing splaying mode in the external plies, and reducing the number of fragmented plies.

CONCLUSIONS

The influence of initial velocity on the crushing behavior of composite plates is studied. With two materials tested, the influence of speed to the maximum load is easily observed. However, its influence on the energy absorption is not visible.

The results show clearly that Cytec specimen can absorbed more energy than T700/M21 specimen in crush test.

Figure 1 :

 1 Figure 1: Design of the fixture, with (on right) and without (on left) upright

1 )

 1 Material 1: Cytec fabric prepreg 5H (6KHTA) 977-2 Ply thickness: 0.35 mm Three stacking sequence configurations: P1-[(0/45)3]sym P2-[(0)6]sym P3-[(0/45)3,0]sym. 2) Material 2: Hexply unidirectional prepreg T700/M21 (M21/35%/268/T700GC) Ply thickness: 0.26 mm Three stacking sequence configurations: P4-[(0/45/90/-45)*2]sym, P5-[(0/90)*4]sym, P6-[(0/45/90/-45)*2,90,1/2 0°]sym Stacking sequences were chosen so that for each kind of configuration (1 to 3), weight, global thickness and equivalent modulus of the plates are the same for both materials (Table

Figure 3 .Figure 3 :

 33 Figure 3: some examples of load-stroke behaviour

Figure 4 :

 4 Figure 4: Examples of crush modes observed

Figure 5 :

 5 Figure 5: Images from high-speed camera P4

Figure 6 :Figure 8 :

 68 Figure 6: Fmax with different initial speeds However, the results of energy absorption with different speed are difficult to analyze: no clear tendency is observed. The dispersion of the results can be explained by the different observed crush modes which depends on the debris wedge created in the test. A debris wedge made of debris issued from the initiation process or from fracture of central plies can appear at different time and create different crushing modes in which the

Table 1 :

 1 Specimens characteristicsConfig. 1

		Config. 3
	0/90	0° oriented

iso Config. 2