Javier Toral Vazquez

Bruno Castanié

Jean-Jacques Barrau

Nicolas Swiergiel

MULTI-LEVEL ANALYSIS OF LOW-COST Z-PINNED COMPOSITE JOINTS PART 1: SINGLE Z-PIN BEHAVIOR

Keywords:

In the framework of the EU program ALCAS (advanced low-cost aircraft structure), a new Z-Pinning technique was developed by EADS Innovation Works. It was used to manufacture low-cost Z-pinned junction demonstrators (L and T shaped specimens) typical of aeronautical structures. In order to understand load transfer mechanisms in this kind of assembly, a multi-level analysis was performed. Firstly, tension and shear pin behaviour was investigated as well as pin pull-out from net resin. It was demonstrated that the mechanical transfer is mainly through bonding, even when the pins are twisted. Secondly, an analytical model was proposed which predicts the maximum load capacity of a single pin. Finally this study provides the basis for a design methodology to compute ultimate loads of Z-pinned junctions under complex loading.

As reported by Mouritz [START_REF] Mouritz | Review of z-pinned composite laminates[END_REF], z-pinning is one of the most promising reinforcement techniques for classical laminates. Z-fiber®, developed since the eighties by Aztec Inc is the most widespread pinning technique. The technology uses an ultrasonic horn to push pins, held in low density foam, through the laminate before curing. The manufacturing method is detailed in [START_REF] Partridge | Manufacture and performance of z-pinned composites[END_REF]. Although titanium rods were used at the beginning of the technique's development, currently, carbon pins are the most common. For this technology, the pins are typically less than 0.5 mm in diameter and a low percentage value of z-pin density is required to obtain significant reinforcement in the out-of-plane direction. Logically, investigations carried out on z-pinned laminates have focused on the in-plane mechanical properties, their delamination behavior and the impact or post-impact response [START_REF] Zhang | Improvement of low velocity impact and compression-after-impact performance by z-fibre pinning[END_REF]. Z-pinning slightly decreases the in-plane mechanical properties of laminates due to the pin insertion process [START_REF] Chang | Properties and failure mechanisms of zpinned laminates in monotonic and cyclic tension[END_REF][START_REF] O'brien | Influence of compression and shear on the strength of composite laminates with z-pinned reinforcement[END_REF][START_REF] Mouritz | Compression properties of z-pinned composite laminates[END_REF][START_REF] Steeves | In-plane properties of composite laminates with through-thickness pin reinforcement[END_REF] which creates an undulation of the fibres (see Figure 1). The eye-shaped resin pockets produced can influence the response under thermal loading or fatigue [START_REF] Sweeting | The effect of thermal mismatch on Z-pinned laminated composite structures[END_REF][START_REF] Mouritz | Tension fatigue of fibre-dominated and matrixdominated laminates reinforced with z-pins[END_REF][START_REF] Zhang | Experimental Study and Computer Simulation of Z-Pin Reinforcement under Cyclic Fatigue[END_REF].

As expected, the delamination strength can increase significantly in mode I [START_REF] Partridge | Delamination resistant laminates by Z-Fiber pinning: Part I manufacture and fracture performance[END_REF][START_REF] Partridge | Delamination of Z-pinned carbon fibre reinforced laminates[END_REF][START_REF] Dai | Experimental study of z-pin bridging law by pullout test[END_REF].

The mechanisms that increase the delamination strength are the bridging forces that are created while the pins are pulled out from the laminate under opening loads. The strength of an interface can also be reinforced under shear loading by Z-pinning [START_REF] Partridge | Delamination of Z-pinned carbon fibre reinforced laminates[END_REF].

In this case the behaviour of a pin is complex. It is subjected to a combination of pullout, internal shear which can cause splitting of the pin and transverse shearing.

Moreover, the failure mode is also dependent on the stiffness of the resin area and the laminate itself [START_REF] Partridge | Delamination of Z-pinned carbon fibre reinforced laminates[END_REF]. Pin behaviour under pull-out loadings has been investigated experimentally [START_REF] Dai | Experimental study of z-pin bridging law by pullout test[END_REF], [START_REF] Cartié | Mechanisms of crack bridging by composite and metallic rods[END_REF] and numerically [START_REF] Gornet | Inelastic Interface Damage Modeling with Friction Effects: Application to Z-Pinning Reinforcement in Carbon Fibre Epoxy Matrix Laminates[END_REF] showing a two or three slope loaddisplacement response. Dai et al have also shown that small diameter pins are more efficient than larger ones. The research carried out on the Z-Pinned joint will be presented in part 2 of this publication. It will be demonstrated that the behavior of the junction under pull-out, shear and unfolding is globally homothetic to the mechanical behaviour of a pin. In spite of the works on the Z-pinned structures, the applicability of the results on EADS IW technology is still to be demonstrated. Moreover, no sizing methodology was undertaken to design joints. So, in the following section, the new pinning technique will be presented and the choice of a multi-level analysis justified.

2-EADS Innovation technology and methodology of analysis.

In the framework of the European ALCAS program (Advanced Low-Cost Aircraft Structure), investigations were conducted to develop low-cost composite junctions enabling automatic manufacturing and diminishing or avoiding manual layup. A new technology of automatic insertion of Z-Pins developed by EADS IW is analysed in this study [START_REF] Lefèbure | Assembly of dry Carbon Preforms by a Z-pinning technology[END_REF]. The technology uses twisted carbon fibre pins produced by Hexcel. Their diameter here was 0.7mm and their length 40 mm (Figure 2). The pin insertion system is held in a 7-axis machine so the pinning process can be automated (Figure 3). The pins are longer than Z-fiber® carbon pins [START_REF] Partridge | Manufacture and performance of z-pinned composites[END_REF] in order to be able to manufacture T-joints without a stiffener flange, thus creating a "low-cost" junction. A vibrating holed needle with the Z-Pin inside (see Figure 3) penetrates into a dry preform provided by Dassault Aviation. Once the desired depth is reached, the needle comes out leaving the Z-Pin inside the preform (Figure 4). Finally, the pins are cut before injection of the preform using an resin transfer molding (RTM) system. The objective of the present study was to provide users a pre-sizing method for this new technology. In aeronautics, such sizing methods are related to two concepts: the notion of limit and ultimate loads and the "no-crack growth" principle. The limit loads are the maximum loads that any aircraft can be subjected to during its life. Under the limit loads (LL), the structure must be calculated for use under fatigue loading and it is the reason why absence of crack growth is mandatory for composite parts. The ultimate loads (UL) are equal to 1.5 LL and the structure must sustain this load only statically. This ability is validated by static certification tests which are required to allow the aircraft to fly. Between the Limit and the Ultimate Loads cracks can propagate because a real aircraft will never see such loadings during its life.

According to the literature, delamination energy is absorbed by Z-pins via frictional pull-out so this technology does not provide real advantages until limit loads.

Moreover, we show in the companion paper that cracks initiate in joints at about the same loads with and without Z-pinning as already pointed out by Greenhalgh et al [START_REF] Greenhalgh | Evaluation of toughening concepts at structural features in CFRP -Part I: Stiffener pull-off[END_REF]. However, real structural gains are demonstrated for ultimate loads because, for a classical bonded joint, the propagation of a crack is faster. So the design methodology for such pinned joints can be separated into two: classical design methods for limit loads which are not the subject of the present research and a new pre-sizing method for ultimate loads. The assumptions made and the methodologies adopted in this paper are presented hereafter.

As for Z-fiber technology, the insertion process creates resin pockets around the pin as shown in Figure 1. In the present case, visual observations indicate that the pins are surrounded mostly by resin. Only few laminate fibres are in contact with the pin, probably due to the fact that the diameter of the needle is necessarily greater than the diameter of the pin. Thus, the main assumptions made during this investigation were:  Load transfer between pin and surrounding matrix is mainly due to bonding.

 The behaviour of a pinned joint is homothetic to the behaviour of a single Z-pin.

The observations made led us to present a multi-level analysis of pinned joints (Figure 5). The first level includes pin and resin elementary mechanical characterisation. Resin analysis consisted of a standard matrix characterisation showing slight non-linear behaviour [START_REF] Vasquez | Identification et modélisation du comportement des structures composites assemblées par cloutage[END_REF] and will not be presented here. Only single pin tests under traction and shear will be presented. The second level considered pin pull-out behaviour when bonded in resin with an analytical model to predict pull-out strength. At this level, pull-out behaviour of single pins inserted into laminates will show similarities with pull-out behaviour from resin thus validating the first assumption. A numerical model is developed able to calculate the debonding load.

The third level will be presented in the companion paper and will deal with L and T shaped pinned structural joints tested under tension, shear and unfolding loads. The test results demonstrated the utility of pinning structural joints from a static point of view. Ultimate loads were calculated assuming homothetic behaviour of the pins in the structure and no interaction between pins. The pins were modelled as non-linear springs. The non-linear laws under tension and shear of the springs were obtained from the combined experimental and numerical results of the first and second levels of the analysis thus demonstrating the suitability of the pre-sizing method presented.

3-Single pin testing

Tension

The carbon pins used in this study consisted of T700 carbon fibre tows impregnated with M18 epoxy matrix (Figure 2 (a)). The carbon fibre tows were slightly twisted to obtain an irregular outer surface (Figure 2 (b)). The pins were delivered with a matrix polymerisation of 80%, to be completed during joint manufacturing to allow chemical links with laminate resin to be created. Carbon pins were tested under tension to measure their strength and stiffness. Before testing, the pins were cured for 3 hours at 180ºC to complete resin polymerisation. The tests were performed on an Instron 10kN tension test machine. The tension load was applied on a single pin through mechanical grips. To protect the pin from damage by the grips, small carbonepoxy plates were bonded to both pin ends. The plates were 2 mm thick and had a hand-made V-shaped groove to fit the pin ends (see Figure 6). The two plates were not in contact to allow good clamping of the pin. Load was measured by the load cell while two inductive sensors gave the displacements [START_REF] Vasquez | Identification et modélisation du comportement des structures composites assemblées par cloutage[END_REF].

Three tests were performed. Load/displacement curves are shown in Figure 7.

The curves show a slight non-linear behaviour with a decrease of the tangent modulus for the highest load values. The same non-linear behaviour has been observed for steel cables also composed of twisted tows [START_REF] Jiang | A concise finite element model for simple straight wire rope strand[END_REF]. Thus, it seems very likely that the source of the pins' non-linearity is the twist in the carbon tows. From the test results, a stress/strain behaviour law was established using engineering stress and strains (based on initial circular section). The behaviour of the 0.7 mm diameter pins is shown in Figure 7. It can be fitted with a polynomial law of degree 3 using least squares [START_REF] Vasquez | Identification et modélisation du comportement des structures composites assemblées par cloutage[END_REF] so that a tangent Young's modulus at low strains of 92 092 MPa can be obtained as well as a final tangent Young's modulus of 32 255 MPa before failure. The pins failed most often near the end grips. The order of magnitude of the mean maximum stress was 2000 MPa, indicating that the test method used is globally correct as this value matches well classical UD carbon-epoxy tensile strengths.

Shear

Pins were tested under shear loadings using a specially designed test assembly. The test assembly consisted of two coaxial cylinders with a hole perpendicular to the cylinder axis as shown in Figure 8. The external cylinder was attached to a rigid rig fixed to the machine frame. The internal cylinder slipped inside the external one to apply a shear loading on the pin in the hole. Three pins were sheared with just one end in the device and three others with the pin crossing through the two holes. The cylinders were made of brass to minimise friction and for manufacturing reasons. The load-displacement curve of the pin shear test is presented Figure 9. The first stage, where load increases slightly corresponds to the displacement due to the clearance and to initial contact of the twisted pin in the test device. This initial contact also involves crushing the rough points on the pins. This is followed by a linear load increase until fragile pin failure. The strength of the pins and the shear stiffness are reported in Table 1. The stiffness value is probably influenced by the clearance between the test rig components (measured at 0.1 mm), a point which will be discussed in the second paper which focuses on pinned joint modelling and testing. From these values of strength and stiffness, pin shear behaviour can be modelled as a linear fragile behaviour law.

4-Pull-out testing

Pull-out tests were performed on pins bonded in epoxy resin. Specimens consisted of a carbon pin bonded in the centre of a resin block with ( Figure 10(a)).

The specimens were manufactured using a metal mould of internal dimensions 50 x 50 x 15 mm (see Figure 10(b)). The specimen preparation process in fully described in [START_REF] Vasquez | Identification et modélisation du comportement des structures composites assemblées par cloutage[END_REF]. The first step involved heating the resin to 120°C for 10 min in order to reduce its viscosity. Then, the still liquid resin was poured into the mould while pin was held vertically. When filling the mould, resin height was obtained with a precision of +/-1 mm. The specimen was then cured following the appropriate curing cycle. Different pin-resin bonded lengths were tested by modifying resin block height. Twenty-three specimens were manufactured. To carry out the pull-out tests, the vertical displacement of the specimen was blocked over the upper surface of the resin brick 5 mm from the pin centre as seen in Figure 11 (a). Pin pull-out tests were performed using a standard tensile machine and a special grip to hold the upper part of the pin to apply the load (see Figure 11 (b)). The applied load and the vertical displacement of the pin were recorded during the test by two Linear Variable Differential Transformer (LVDT) sensors. A mass was added to minimize dynamic effects during fragile failure. A zoom on the grids is shown in figure 11 (c). Initial tests were conducted with dial recorders to validate the measures given by the LVDT.

Figure 12 shows three typical load-displacement curves from pull-out tests.

Three stages can be distinguished: first, the pin's elastic deformation before reaching maximal load. This is followed by an applied-load drop due to resin-pin interface failure and, finally, the debonding stage where friction effort decreases while the pin slips inside the resin block. For all the tests, the initial debonding load spike was higher than the friction load during the third stage. These three stages have been described and observed during pin pull-out tests from laminates performed by Dai et al [START_REF] Dai | Experimental study of z-pin bridging law by pullout test[END_REF]. Finding the same global behaviour gives the first similarity indication between pins bonded in pure resin and pins inside laminate. We investigated the influence of bond length over maximum pull-out load. Twenty-three pull-out tests were performed with different bonded-lengths: from 2 to 12 mm. The results are presented in terms of maximum pull-out load versus bond length. Figure 13 shows the test points. The test results are quite scattered. Some post-mortem analysis using scanning electron microscopy (see Figure 14 and15) showed that the fracture surface matches up with the pin-resin interface but can also go though the pin itself. This observation indicates that the pin to resin joint is as resistant as pin matrix and fibre joint. However, the irregular pin surface (see figure 2) that produces different stress concentration points at the pin-resin interface for each specimen can produce scatter. The scatter could also be due to the influence of the meniscus on the upper face of the resin block. The radius of the meniscus was measured and scatter also appeared with radii of 1.5 mm to 4 mm for a block thickness of about 5 mm and a radius reaching 9 mm for block of 11.6 mm height. As this area is also an area of stress concentration, the differences can also lead to scatter. Finally, even if tests were made on some specimens showing no defects, it is still possible that voids appear in the resin block or at the interface. The points in Figure 13 can be extrapolated to the origin as, for a null bonded length, the resistance is zero. Thus, despite the scatter, a general trend can be observed: pin resistance increases for bond lengths between 0 and 4 mm. For bond lengths over 4 mm, the debonding load remains fairly constant. This behaviour corresponds to that of a typical bonded joint where, after a given bond length, strength saturates.

5-Pull-out modelling

The pin pull-out process presents some similarities with fibre pull-out from resin. Fibre pull-out tests were developed to characterise the fibre/resin interface by measuring the maximum shear stress at the interface. Several models are proposed in the literature to analyse fibre pull-out. Analytical models from Zhang et al [START_REF] Zhang | On steady-state fibre pull-out I The stress field[END_REF] and Fu et al [START_REF] Fu | Analyses of the micromechanics of stress transfer in single and multi-fiber pull-out test[END_REF] provide the stress distribution at the fiber/resin interface for a fibre bonded in net resin and a fibre bonded in net resin and surrounded by orthotropic material respectively. Based on these two models, a new model was developed to calculate the stress state at the pin/resin interface for the configuration of pin pull-out tests presented above. The model considers the geometry and configuration shown in Figure 16. Full details of the calculation can be found in reference [START_REF] Vasquez | Identification et modélisation du comportement des structures composites assemblées par cloutage[END_REF].

The pin is bonded in a resin cylinder of radius c. An equilibrium occurs between the pullout force F 0 and the contact pressure at the upper surface of the resin cylinder for radius greater than b. Pin and resin are assumed to be elasticisotropic materials. For a given horizontal section, the load balance over the X component can be written as follows:

      0 , 2 , 2 2 1 2          c r b r m x b r a r m x f x rdr x r rdr x r x a       (Eq 1)
Indices f, m1 and m2 indicate the pin, the resin for radius bellow b and the resin for radius bigger than b respectively. According to Fu et al [START_REF] Zhang | On steady-state fibre pull-out I The stress field[END_REF], shear stress in the resin can be written:

      2 , 1 ,    i for r x q r x p x r mi mi mi  (Eq 2)
Where p m and q m are the functions to be determined. The stress balance between shear stress at the pin/resin interface  c (x) and normal stress inside the pin led to:

    x a dx x d c f x   2 
(Eq 3)

In the same way, stress balance between shear stress inside the resin at r=b and the axial stress over the outer region of the resin can be written:

    0 , 2 , 2 2 2       c r b r m x m rdr dx x r d x b b     (Eq 4)
At r=a and r=b, shear stress over the different domains is related by the following expressions:

          x x a x b x b x b C m m M m           , , , , 1 2 1
(Eq 5)

Considering that normal radial and hoop stress are negligible compared to axial stress, the axial strains can be written as follows:

  2 , 1 , ) , ( ) , ( , m m f j for E x r x r dx x r du j j x j xx j x      (Eq 6)
The boundary conditions are:

      0 0 0 , 0 2      x L x x c f x f x m     (Eq 7)
Solving the above equations [START_REF] Vasquez | Identification et modélisation du comportement des structures composites assemblées par cloutage[END_REF], gives the axial stress distribution along the pin and also the shear stress distribution along the pin/resin interface:

            x L a x x L x c f x                 cosh sinh 2 sinh sinh 0 0 (Eq 8)
Where  is a function of radii a, b and c and of the mechanical properties of pin and resin. The very complex expression of  is given in Appendix 1. The shear stress distribution shows its maximum value at the highest point of the interface x=L. Thus, it can be assumed that pin disbonds from resin when the shear stress reaches a critical value at this point. This assumption involves:

  L a F MAX         tanh 2 max (Eq 9)
As the maximum shear stress is an interface pin/resin characteristic, the debonding load can be written as follows:

    L a L F         tanh 2 max max (Eq 10)
From this expression, the critical shear stress of the pull-out tests can be obtained by fitting the F max (L) analytical expression with the experimental data points.

In this manner, a critical value of 58 MPa was found corresponding to the curve in Figure 15. Thus, the analytical model gives a reference for the variation of the pull-out load versus bond length. The scatter of the experimental points makes it difficult to validate this curve but overall, it follows the main trend. However, some discrepancies can be found, in particular for bond lengths above 8 mm where the model seems to under-estimate the pull-out strength experimentally obtained. It is possible that for lengths of over 8 mm, cracks have to propagate in the pins thus increasing the failure energy required (see figure 14 and15).

The analytical model is based on strong assumptions, especially the isotropic linear behaviour of resin and pin and the fact that the menisci around the pin were not take into consideration. In order to estimate the validity of the analytical model, its results were compared to those provided by a FEM model. The FEM model was developed using SAMCEF software [START_REF]Samcef User's Guide[END_REF] and consists of an axisymetric model including the resin menisci around the pin and with boundary conditions in line with the tests. The geometry, the mesh and the boundaries of the model are reported in Figure 18. The typical size of the elements was 0.01 mm in the most refined area; using smaller elements did not change the stress values obtained.

The mechanical properties of pin and resin were those obtained from the characterisation tests presented in section 3. The pin material traction law in the axial direction corresponds to the curve shown in Figure 7. Pin transversal properties correspond to a typical carbon-epoxy UD ply (E t =9000 MPa,  lt = 0.25, G lt = 4000 MPa). Resin is assumed to present an isotropic plastic behaviour as determined through tension tests [START_REF] Vasquez | Identification et modélisation du comportement des structures composites assemblées par cloutage[END_REF]. The shear modulus of the resin used was 1443 MPa and  =0.3. The first study compared the shear stress at the pin/resin interface between the FE model and the analytical model. The case study corresponds to a resin block 5.5 mm high and an applied load of 300N. This load level was of the same order of magnitude as the debonding load measured during pull-out tests. The shear stress comparison in Figure 19 shows a good agreement between the two models. The agreement was worse in the upper part of the interface where the analytical model over-estimated shear stress. The mismatch could be explained by the fact that the analytical model did not consider the plastic behaviour of the resin. Thus, the real maximum value of the shear stress should be close to the FEM results and consequently lower than the critical shear stress used by the analytical model. After shear stress comparison, the FE model was employed to calculate the changes in debonding load with the bond length. Two criteria are proposed: the Tresca criterion and a maximum strain based criterion. To compare analytical and numerical models, the critical criteria value was set at a common point over the curve of strength vs. bond length for the two models. This point corresponds to L=5.5 mm and F max = 328 N given by the analytical model after matching the experimental results. At this point, the values are 48 MPa for the Tresca criterion and 0.0409 for the maximum strain criterion. Then, by modifying the bond length, the maximum applied load that fulfils failure criteria was calculated. Figure 17 shows debonding load vs. bonded length for both criteria applied to the FE model compared to analytical model results. The analytical model and the two criteria used with the FE model present globally the same variation, in particular with a maximum load plateau for bond lengths over 6 or 7 mm. The Tresca criterion gives strength values higher than strain-based criteria for bond lengths bellow 5 mm. This could be attributed to resin plastic behaviour. Indeed, as the stress vs strain relationship is no longer linear, close to the failure point strains increase faster than stress explaining the fact that the Tresca criterion is more conservative. Comparison between the analytical and numerical models shows that, despite the simplifying assumptions made, after a single experimental adjustment, the analytical model seems valid, especially for low bond lengths. This point is quite important in practice as in a typical aeronautical pinned T-joint (see Figure 20), pins always fail in pull-out by debonding from the skin, where the insertion length is shorter. Despite its complexity, the FE model also presents some limitations as it cannot explain the load scatter encountered in the test results.

6-Pull-out from laminates

The objective of this part of the study is to analyse the behaviour of a single Zpin in a laminate and to compare it with the behaviour in net resin. Specimens with a single pin which mimics the inclusion of pins in the web with different depths were manufactured (see Figure 21). Specimens were made with pins in the skin (replacing the resin block in figure 10 by a laminate) but it turned out to be impossible to manufacture in an autoclave. The specimens consist of two half carbon/epoxy plates jointed by a single pin and separated by foam with two layers of Teflon to avoid bonding between the edges of the plates. The plates were made of twelve 914/G803 pre-preg tissues for a total thickness of 3.4 mm. Pins were 0.7 mm in diameter and 40 mm long. During lay-up, the pin was placed manually in the middle of the layers of laminate (see Figure 22 (a)) and the insertion depth L in one of the two plates is checked (see Figure 22 (b)). Table 2 lists the 47 specimens tested, obtained from 8 plates and an insertion depth varying from 3 to 20 mm. Pull-out tests were performed by fixing both specimen ends inside the 10kN Instron tension machine grips. The applied load was measured by the machine load cell while displacements were obtained from LVDT sensors. In order to isolate and correct any parasitic rig displacement, some tests were instrumented with cameras to obtain specimen displacement by digital image correlation (see Figure 23).

Load-displacement curves are globally similar to those obtained from pin pullout from net resin as presented in Figure 12. However, looking at a smaller displacement scale, a load-displacement curve indicating a different behaviour of the type shown in Figure 24 for a 15 mm insertion depth, can be observed. Indeed, the decrease in tangent modulus is significant before reaching the maximum load level.

By comparing with the non-linear response of the free length of the Z-pin between the two plates (3mm), it can be shown that other mechanical properties are involved [START_REF] Vasquez | Identification et modélisation du comportement des structures composites assemblées par cloutage[END_REF].

For a better understanding of the mechanical process behind this phenomenon, tests were stopped at different stages of debonding. Micrographic analysis of specimen 5G, stopped before reaching the maximum load, shows some cracks at the pin/matrix interface (Figure 25 (a)). Crack length is around 7 mm beyond which the interface seems not to be damaged (Figure 25 (b)). In contrast, in specimen 6G, stopped after final load drop-off, shows the pin fully debonded from the resin (Figure 25 (c)). Here, cracks appear all over the interface as well as in areas inside the pin itself (figure 25 (d)). Crack propagation inside the pin may be due to the twisted tows that form the pin interfering with the crack path. Thus, the main difference between pin pull-out from resin and pull-out from laminate is the more progressive debonding of pin from the laminate. Micrographic analysis showed that the tangent modulus decrease was due to initiation of interface cracking.

The influence of insertion depth on pin pull-out behaviour can be analysed by comparing the plots in Figure 26. The curves appear to be globally superimposed over both the first quasi-linear load increase and the second slope corresponding to interface crack propagation. The maximum applied load increased as the insertion length increased. From these curves, it is possible to obtain the crack initiation load level and the maximum load level for each insertion depth (Figure 27). The next step consists of comparing these results with experimental and numerical values of pin pull-out from net resin. The debonding load of a pin bonded in resin is comparable to the load initiating pin interface cracks in laminates (Figure 28). The crack initiation load increase for insertion depths of over 6 mm is less steep here than for the maximum applied load presented above. Note that the pin requires a higher load to be pulled out from a laminate than from resin if the insertion depth is greater than 6 mm. These analyses are consistent with the failure scenarios identified through micrographic analysis and they indicate that crack initiation at the pin/resin interface or pin/laminate interface occurs for the same load level. If the pin is bonded in net resin, the crack spreads sharply over the whole interface and causes pin debonding while if the pin is inserted in a web, crack propagation at the interface shows a stable stage before complete debonding.

The initial assumption that pin behaviour in a laminate is similar to that in net resin is only partially true. For insertion depths of up to 6 mm, it seems to be a fairly accurate approximation and the analytical model can provide a maximum force. For deeper insertion, however, it is no longer the case but the ultimate load is underestimated which remains acceptable in pre-sizing. Finally, in our specimens we placed the pins manually between two plies without using the vibrating needle before curing under pressure. Thus, there is certainly less space between the plies and the pins, and more numerous contacts. The contacts may play the role of crack arrestor increasing the energy required for crack propagation and thus the maximum force.

They probably provide an upper boundary for this force by comparison to a pin in a skin (see picture [START_REF] Zhang | On steady-state fibre pull-out I The stress field[END_REF]. At this stage of the study, the assumption of identical behaviour in net resin and in laminate remains viable even if certain limitations and special behaviour in a web were highlighted. In fact, in the companion paper, we report that pins never debond from web but only from skin even with an equal insertion depth, thus the load calculated by the numerical model is able to predict the debonding load.

Nevertheless, to achieve an accurate representation of behaviour of a pin inserted into a web, a more complex model able to represent crack propagation between a pin and a laminate should be developed but this point was beyond the scope of a presizing method.

7-Conclusions.

In this first part of the study, the behaviour of single Z-pins was analysed.

Tensile and shear strength of the pins was measured and will allow the pinned joints to be modelled, as reported in the second part of this publication. Then, an experimental study of pin pull-out from net resin was carried out. Pin debonding seems to be controlled by shear stress at the pin/resin interface as for generic bonded joints. An analytical model based on the principles of continuum mechanics made it possible to find this behaviour with only one experimental adjustment. Also pull-out test of a pin inserted into laminate, which mimics the pin in a web was carried out.. The main difference with pin in neat resin is stable crack propagation at the pin/laminate interface. However, crack initiation load level is mostly the same for both pins in net resin and pins in laminate which is coherent with sudden interface crack propagation when the pins are bonded in net resin. However, the debonding load increases with pin insertion depth. This effect is probably due to the fact that pin is parallel to the ply in this experiment. Stable crack propagation requires more energy and can explain why, in joints, pins always debond from the skin even with equal insertion depth. To design joints for ultimate load, it is necessary to know the maximum force on the pin/laminate junction and the assumption that the debonding load of a Z-pin can be calculated with the numerical model remains valid and will be confirmed in the second part of this article.

8-Acknowledgment

The authors would like to thank Thierry Vilain (Dassault Aviation and ALCAS WP3 leader) and Patrice Lefebure (EADS IW) for their support.

9-Appendix

The expression of coefficient  is given below. The details of the calculation can be found in [START_REF] Greenhalgh | Evaluation of toughening concepts at structural features in CFRP -Part I: Stiffener pull-off[END_REF].

1 3 1 2 2 2 2 4           (Eq A-1)
With:

4 13 2 12 32 1 C A C A C    (Eq A-2) 4 13 1 11 32 12 2 C A C A C A     (Eq A-3) 11 3 A   (Eq A-4)
Where:

2 2 2 1 a b E E C f m   (Eq A-5) 1 2 2 2 2 mb A a b C   (Eq A-6)              1 2 2 2 2 2 3 1 2 2 I b c b A a b C m mb  (Eq A-7) 3 2 2 32 2 C b b c C    (Eq A-8) 2 2 2 4 b c C    (Eq A-9) With:                  a b b a b a b a A m mb ln 2 1 1 2 2 2 2 2 2 1  (Eq A-10)                  a b a a b a b b A m mb ln 2 1 1 2 2 2 2 2 2 2  (Eq A-11)   8 4 ln 2 2 2 2 2 2 2 2 1 b c b c b c c c I             
(Eq A-12)

And:

  1 1 2 11 2 C M a A    (Eq A-13)   2 2 12 2 C M A   (Eq A-14)   32 32 13 2 C M A   (Eq A-15) With:   2 2 2 1 a b E E M f m   (Eq A-16)                              a b b a b b a b a b a M m ln 2 4 8 1 2 2 2 2 2 2 2 2 2 2 2  (Eq A-17)                             

FIGURESFigure 1 :Figure 2 :Figure 3 :

 123 FIGURES

Figure 4 :

 4 Figure 4: Insertion process.

Figure 6 :

 6 Figure 6: Single pin specimen with loading plates.

Figure 7 :

 7 Figure 7: Typical load-displacement curves for Z-pins 0.7 mm diameter under tension.

Figure 8 :

 8 Figure 8: Shear device for single Z-pins.

Figure 9 :

 9 Figure 9: Shear tests force/displacement curves for 0.7 mm pins.

Figure 10 :Figure 11 :

 1011 Figure 10: Specimen for pull-out test a single pin in net resin (a) and its mould (b).

Figure 12 :

 12 Figure 12: Typical pull-out force/displacement curves. 5 mm

Figure 13 :

 13 Figure 13: Experimental pin pull-out strength values.

Figure 14 :

 14 Figure 14: Detail of a strand of fibres which remained bonded to the resin.

Figure 15 :

 15 Figure 15: Detail of isolated fibres which remained bonded to the resin.

Figure 16 :

 16 Figure 16: Modelling Z-pin pull-out.

Figure 17 :

 17 Figure 17: Experimental pin pull-out strength values compared to analytical model results.

Figure 20 :

 20 Figure 20: Typical T-Joint.

Figure 21 :Figure 22 :

 2122 Figure 21: Pin inside laminate: Pull-out specimen diagram principle.

Figure 23 :Figure 26 :Figure 28 :

 232628 Figure 23: Pin pull-out from laminate test assembly

Table 1 :

 1 Shear test results

			TABLES
		Maximum applied load (N)	Linear increase slope (N/mm)
	Diameter (mm)	Mean value	Standard deviation	Mean value	Standard deviation
	0.7	288	16.8	1396	228
	Insertion depth (mm) Number of Specimens Reference Displacement measurement
	3	7	E		LVDT
	5	4	D		LVDT
	6	4	H		LVDT
	7	3	H		LVDT
	8	7	F		LVDT
	10	4	D		LVDT
	15	3	G		LVDT
	15	3	I	LVDT + Cameras
	20	4	A		LVDT

Table 2 :

 2 Pin inserted in laminate specimens


Bruno Castanié
email: bruno.castanie@insa-toulouse.fr

Javier Toral Vasquez

Jean Jacques Barrau

Nicolas Swiergiel

MULTI-LEVEL ANALYSIS OF LOW-COST Z-PINNED COMPOSITE JUNCTIONS. PART 2 : JOINTS BEHAVIOR

Keywords: A. 3-Dimensional reinforcement, B. Stress Transfer, E. Joints/joining

In the framework of ALCAS (Advanced Low-Cost Aircraft Structure) E-U program, a new Z-Pinning technique was developed by EADS Innovation Works. This technology was used to manufacture low-cost Z-pinned junction demonstrators (L and T shaped specimens) typical of aeronautical structures. In order to understand load transfer mechanisms in this kind of assembly, a multi-level analysis was performed. Firstly, tension and shear pin behaviour was investigated as well as pin pull-out from neat resin. It has been demonstrated that the mechanical transfer mode is mainly by bonding even if the pins are twisted. Secondly, an analytical model was proposed which enables to predict the maximum load capability of a single pin. Then, in the second part of this publication, it will be demonstrated that the behaviour of the junction under pull-out, shear and unfolding is globally homothetic to the mechanical behaviour of a pin. Finally this study provides the basis for a design methodology for Z-pinned junctions under complex loading.

1-Introduction.

As reported by Mouritz [START_REF] Mouritz | Review of z-pinned composie laminates[END_REF], the z-pinning is one of the most promising reinforcement technique for classical laminates. Logically, most of the investigations carried out on zpinned laminates have focused their attention on the in plane mechanical properties, their delamination behaviour and the impact or post-impact response. However, a little number of authors has investigated the use of the Z-Fiber technology in joints. Lap-joint specimens reinforced by Z-pinning were tested [START_REF] Chang | Properties and failure mechanisms of pinned composites lap joints in monotonic and cyclic tension[END_REF] illustrating the behavior of the reinforced joint mainly under shear. It was shown that the static and the fatigue strength were both increased by 40 %. However, this result depends on the density of pinning, in this case 2%. Delamination appears for lower densities and higher densities where a static failure occurs at the border of the pinned zone. The same order of magnitude in the strength increase was obtained by Grassi et al [START_REF] Grassi | Simulation of pin-reinforced single-lap composite joints[END_REF] for joints in pipes. In order to increase the shear strength of such a joint, the idea of using inclined pins were analysed by Rugg et al [START_REF] Rugg | Damage mechanism for angled through thickness rod reinforcement in carbon-epoxy laminates[END_REF]. For +/-45° angle and the configuration studied, the static strength were increased by 100%. Nevertheless, the crack initiation at the interface is only slighty influenced by the pins. One interesting point is the fact that the mechanical behavior of the pins is very different where pushed or pulled depending on their relative orientation to the applied load: pins been pulled disbond easily while pins been pushed support high shear load level before failure. However, no influence of pins inclination have been observed by Chang et al [START_REF] Chang | Elevated Temperature Properties of Pinned Composite Lap Joints[END_REF] for lower inclination angles between 13º and 23º and for temperatures bellow 150ºC. Z-fiber technology has also been considered for increasing joint behaviour of aeronautic joints. Nevertheless, the only known application is the reinforcement of skin to stiffener joints of F/A-18 E/F "Super Hornet" [START_REF] Exline | Z-Pins strengthen Super Hornet structures, save weight and cost[END_REF]. This kind of reinforcement on cocured stiffener feet has been investigated by analogy through DCB tests by Byrd and Birman [START_REF] Byrd | Effect of temperature on stresses and delamination failure of zpinned joints[END_REF][START_REF] Byrd | Effectiveness of z-pins in preventing delamination of co-cured composite joints on the example of a double cantilever test[END_REF]. T-joint pull-out tests have also been performed by Rugg et al [START_REF] Rugg | Mixed mode delamination of polymer composite laminates reinforced through the thickness by z-fibers[END_REF].

Depending on the web to support distance, delamination starts at the web root or at the flange end what can be justified by energy balance analysis. Cartié et al [START_REF] Cartié | 3D reinforcement of stiffener-to-skin T-joints by Z-pinning and tufting[END_REF] have compared experimentally T-joints tufted and pinned and also proposed a model based in fracture mechanics. By knowing the cracks propagation paths and using a linear mixed mode criterion, the authors were able to correlate experimental load/displacement curves. Greenhald et al [START_REF] Greenhalgh | Evaluation of toughening concepts at structural features in CFRP -Part I: Stiffener pull-off[END_REF] have investigated pinned T-joints behaviour under pull-out showing a load carrying capacity 229% higher than the baseline. However, delamination initiation load level is only slightly affected. In summary, unlike stitched joints ([12]- [START_REF] Stickler | Parametric analysis of stitched composite T-joints by the finite element method[END_REF]), few studies have focussed their attention on composite structures joint by pinning.

It is true that Z-Fiber ® is constraint to be applied principally to stiffeners foot to skin joints, EADS IW technology [START_REF] Lefèbure | Assembly of dry Carbon Preforms by a Z-pinning technology[END_REF] allows deeper pin insertion and thus new junctions' designs can be investigated. In this part of multi-level analysis, representative specimens of typical aeronautics joints have been used to evaluate pin contribution to static behaviour and to propose a modelling methodology.

2-Experimental analysis.

Specimen and test description.

L and T shaped specimens, most typical for aeronautical joint were chosen for this study. The specimens were manufactured following the description made in the first part of this article and in [START_REF] Lefèbure | Assembly of dry Carbon Preforms by a Z-pinning technology[END_REF]. L-joints geometry is presented in Figure 2. The skin is manufactured from UD Uniwave PRIFORM IMS 200 plies with the following stacking sequence: [90/45/0 2 /45/90/-45/0 2 /45/0/-45/0/45] s for a total thickness of 7 mm and where 0º plies are parallel to specimen length (see figure 2). The vertical part attached to the skin corresponds to a rib foot. It is manufactured from NCF PRIFORM HTS 450 with a theoretical thickness of 3.8 mm. The skin to rib-foot interfaces is 0º/0º. Two pin lengths are investigated: 15 and 40 mm. Pinning pattern is presented in Figure 3 Tension and bending tests are instrumented with cameras over both sides of the specimen so crack initiation can be detected using plane image correlation technique through Correli® software [START_REF] Hild | Mesure de champs de déplacements 2d par corrélation d'images numériques : correli 2D , rapport interne LMT Cachan[END_REF]. Load application point displacements are measured by two LVDT sensors. Applied load have been measured through 100 kN load cell and displacement ratio is fixed to 0.2 mm/min. Within this paper, the authors have focused on maximum applied load results and behaviour as they are more impacted by pins presence. However, extra analyses are per-formed to investigate cracks initiation [START_REF] Vasquez | Identification et modelisation du comportement des structures composites assemblees par cloutage[END_REF]. The results from this analysis are in line with previous studies where crack initiation is found to be only slightly influenced by pinning [START_REF] Greenhalgh | Evaluation of toughening concepts at structural features in CFRP -Part I: Stiffener pull-off[END_REF]. This is also the conclusion of stitched T-joints studies [START_REF] Stickler | Experimental and numerical analysis of transverse stitched T-joints in bending[END_REF]. In addition, for the presented technology, some specimens present cracks at the interface before loading so no conclusion can be extracted. It has been demonstrated by using thermal FEM analysis that cracks appear due to thermal loading over a resin rich area [START_REF] Vasquez | Identification et modelisation du comportement des structures composites assemblees par cloutage[END_REF].

Pull-out tests results and analysis.

Load/displacement curves for non pinned L-joints are shown Figure Nº 8. A first irregular response appears up to 500N corresponding to the specimen positioning stage followed by a regular applied load increase until the final sudden failure. Failure load level is highly dependent on the presence of cracks in the specimen before testing. Indeed, correlation image analysis has shown that specimen 1 presented a crack along the whole width of the skin/rib-foot interface at 200 N [START_REF] Vasquez | Identification et modelisation du comportement des structures composites assemblees par cloutage[END_REF]. This crack grows and reaches 7 mm long when applied load reaches its maximum level. It is highly likely that crack appeared during manufacturing process. On the other hand, for specimen 3, no crack has been detected before failure and thus, load increase appears to be very regular and failure occurs without previous signs. Specimen 2 behaviour is between specimen 1 and 3 as no crack has been detected before failure but final delamination occurs through different plies interfaces what may indicate that crack was non visible. To sum up, tension failure load is 2806 N +/-789.

Figure 9 shows typical load/displacement response of configuration C (pinned Ljoint, pins 40 mm long). In this case, scatter is lower than non-pinned configuration with a maximum applied load of 5598 +/-463 N, which represents a 100% increase. Pinned L-joints show cracks appearing on the skin/rib-foot interface at the same load order of magnitude than non-pinned joints (point I). Then, crack propagates along the interface while load goes on increasing (point II). It can be concluded that, during this stage, tension load is mainly carried by pins. Finally, after a first load drop, load decrease slowly while pins slip from skin until total pin pull-out (III). Same behaviour has been observed for configuration B which maximum applied load is 4955 +/-355 N.

Pinned T-joint behaviour is similar to pinned L-joints. Due to low interface surface between skin and web, premature cracks are present so only pins carry the tension load.

The results of pull-out test are summarized in Figure nº 10. Globally, it can be observed that for both pinned L and T-joints and for the two insertion depths investigated, the results are equivalent within a +/-10% range. The differences are due to different boundary conditions for configuration C [START_REF] Vasquez | Identification et modelisation du comportement des structures composites assemblees par cloutage[END_REF], and to the fact that, for configuration B, 15 mm pins are inserted in a resin-rich region of the interface leading to a less effective load transfer capability. For all configurations, it has been observed that pins are not exactly perpendicular to the skin what could generate extra scatter. From all test performed, the mean maximum carried load by one pin is equal to 300N. This value is perfectly consistent with results from analytical model presented in the first part of this study that gives 336 N for a 7 mm pin insertion depth. What's more, pin mean load scatter is lower than scatter from single pin tests probably due to the RTM process producing more homogenous joints. This last remark validates the initial assumption of the homothetical response of pinned joints compare to a single pin.

Bending tests results and analysis.

Load/displacement curves for non pinned L-joints are given in Figure Nº 11. Failure appears suddenly without any appreciable previous crack propagation. Specimen post-mortem appearance is given also in Figure 11. Failure occurs when a crack appears at the outer radius over the UD fillet and propagates through the foot to skin interface.

Maximum applied load reaches 728+/-48 N.

Load/displacement curves for pinned L-joints with 40 mm pins are shown Figure 12 and the related behaviour is described hereafter:

1.Non linear web deflexion until reaching 1848 N of mean applied load (before first load drop in I). A crack appears and propagates over 8mm through the plait and the foot to skin interface (see Figure 12). During this stage, pins debond slightly from the skin. The combination of these two phenomena explains the non linear response 2.Load drop at point I (Figure 12) correspond to the delamination in the radius of the rib-foot. In this area, peeling and interlaminar shear stresses are high and, in addition to this, plies geometry is disturbed by pins insertion.

3.Second load drop is due to the failure of plies in the inner area of the radius under compression loads. Compression failure is a consequence of delamination described in 2 as it leads to a local bending stiffness decrease and so, a redistribution of tension and compression stress. Post-mortem section shows this failure (Figure 13) and also pins debonded from skin under bending.

Bending strength is 1842+/-58N which represents an increase of +153% compared to no-pinned L-joints. L-joints reinforced with 15mm long pins show a very different behaviour as given in Figure 14. Crack appears far over the fillet and propagates from there through the web/fillet interface until reaching the skin/foot interface. In this case, pins are not long enough to be correctly inserted in the web and carry bending loads. pins are slightly more resistant than the ones reinforced with 40 mm pins. However, the difference is thought to be due to the fact that T-joints with 15 mm pins have their three rows more widely spaced that 40mm pins T-joints, which increase bending stiffness and strength.

Shear tests results and analysis.

Figure 17 shows load/displacement curves from L-joints shear tests. Globally, pinned and no-pinned L-joints present the same failure scenario described hereafter: firstly a delamination (point A) appears over the opposite side of the applied load. Indeed, a parasite bending moment is produced as the application load direction may not be perfectly aligned with the specimen interface (figure Nº 18). However, delamination starting point could be different depending on configuration and initial specimen defects.

It can appear inside the fillet or at the end of the stiffener foot [START_REF] Vasquez | Identification et modelisation du comportement des structures composites assemblees par cloutage[END_REF]. After first delamination, a stiffness decrease can be observed as the cracks grow at the interface (point A to B, Figure N°17). Crack path is complex and it can go through a pair of plies interface to the higher or lower one [START_REF] Vasquez | Identification et modelisation du comportement des structures composites assemblees par cloutage[END_REF]. Finally, specimen fails suddenly because of a fast delamina-tion propagation in case of unpinned L-joints, or because of pin shear failure in case of pinned joints. Post-mortem analysis of pinned joints over a section of pin rows shows that most of pins have debonded before shear failure except for the ones under compression load generated by the parasite bending moment (Figure 18). Besides, no influence of pin length has been observed over failure mode or load/displacement curves [START_REF] Vasquez | Identification et modelisation du comportement des structures composites assemblees par cloutage[END_REF]. This observation has been confirmed in terms of maximum applied load as they are exactly the same for both L-joints with 15 and 40 mm pins (Figure Nº 20). Figure Nº20 also shows that pin reinforcement increase only 20% mean shear strength.

T-joints failure mode seems to be different: although delamination appears on the web to skin interface, it does not lead to a stiffness decrease. Some stiffness drops-off are observed over the load/displacement curve (Figure 19) but they are linked to the same pin debonding scenario as for L-joints as it has been deduced from specimen post-mortem analyses. T-joints pinned with 15 mm pins show a maximum carried load 15% higher than T-joints reinforced with 40 mm pins. It has also been observed that pins misalignment angle impacts strongly on shear strength. Curves on Figure nº 19 show a higher strength of specimens whose pins are oriented with the direction of the applied load. In that case, pins are mostly loaded on tension (pull-out) what increase joint strength. On the other side, when pins are oriented against the applied load, the shear component is higher and it makes the joint become weaker. On both cases, a higher strength has been observed for T-joints reinforced with 15 mm than 40 mm. The higher web stiffness due to pinning of 40 mm deep could produce additional loads at the interface that penalizes carrying load capability.

When comparing L-joint solution with T-joint configuration, L-joint strength remains far over T-joint's as the interface surface is much higher.

Conclusions.

Tests performed over T and L-joints with and without pinning reinforcement have made possible to analyze their behaviour under three static loadings: tension (pull-out), shear and bending. The results of this study will allow a future designer of this kind of joints to orientate his choices in terms of configuration. The parameters investigated (T or L shape and pin length) have an impact on joint behaviour but not necessarily under all loadings. For example, pin length has an impact on bending behaviour but modifies only slightly tension and shear response. Generally speaking, for test performed, pinned L-joint configuration shows a better behaviour that T-joint. However, this assertion could not be any more valid for complete stiffened structure where T-joints could be reinforced with inclined pins [START_REF] Petiot | New challenges in structural design and analysis of composite structures[END_REF]. Regarding the type of effort on the pins, they are always loaded under tension and they always fail by debonding from the skin: the thinner part where they are inserted. This means that pin debonding load is a good indicator of joint strength and allows comparison. Moreover, experimental mean pin pull-out load values match well with the prediction from the analytical model presented on the first part of this publication. It validates multi-level approach and allows to feed the model proposed in the next paragraph.

3-Computation of the strength of Z-pinned Junction.

Modelling principles.

The purpose of this section is to present a modelling methodology able to predict pinned joint strength. Experimental analyses have shown, with only one exception, that rib-foot to skin interface is broken on the pinned area before final joint failure. This observation is not verified for pinned L-joints under shear. Indeed, joint failure appears suddenly and it has not been possible to identify latest failure scenario: pins, interface or a mix. Moreover, in this case, interface surface is large enough to carry a significant part of the load so the real load distribution between pins and interface can not be obtained with a presizing model as presented here. In fact, this problem is similar to the bonded/ bolted joints behaviour for which a significant and specific research work needs still to be done ( [START_REF] Fu | Fatigue of hybrid (adhesive/bolted) joints in SRIM composites[END_REF]; [START_REF] Paroissien | Analytical two-dimensional model of a hybrid (bolted/bonded) single-lap joint[END_REF]). For this reason, only T-joints will be modelled under shear as the interface area remains small. General approach is described on Figure 21: interface is supposed to be broken and thus, web and skin are only joined by pins. Pins are represented by non-linear springs which tension and shear behaviour laws are obtained from elementary test presented in the first part of this publication. Within the next section, this general approach will be applied to the variety of specimens and loadings investigated.

Tension and bending.

Tension and bending tests over L-joints have been simulated with a 2D model presented in Figure 22. Skin and rib-foot are meshed with quadrangular elements taking into account plies orientation and considering plain strain. Pins have been represented by 1D elements (rods) with an equivalent section corresponding to the pin section per specimen width unity. Pin nodes are coincident with skin and web ones. At the skin to rib-foot interface, pins are modelled according to the principle described in section 3.1. Material properties applied to the different parts of the model are given in table 1. In terms of boundary conditions, the skin is considered clamped 10 mm away from the web or the end of the rib-foot. Loading under tension has been simulated by applying a vertical displacement over the web. For bending, a horizontal displacement has been applied to the web. Figure 23 shows load/displacement curve obtained by simulation and the one obtained from pinned L-joint tension test. Despite of a higher stiffness for the simulation results, a good agreement has been found between both curves, especially in terms of maximum applied load. The stiffness mismatch could be due to an insufficient clamping action at the skin during the tests. Indeed, while tightening clamping plates over the skin ends, some parasite stress concentration appear on the rib-foot to skin interface due to specimen geometrical defaults. In order to avoid this parasite loading, it has been decided to be used a light clamping load.

Load/displacement curve obtained from FEM shows an irregularity for 4760N. It corresponds to the debonding of the most outer pins row. Thus, from that moment, pins of the first row carry only the load by friction. Thereby, the model allows estimating the load level that produce first debonding of the pins that should correspond to the limit load. In terms of maximum applied load, the model predicts 5343 N, only 5% lower than mean experimental value of 5598 N. The difference is lower than experimental scatter.

Regarding pinned T-joints under tension, the model strength estimation is 5560N, 10% over mean experimental measurement. The difference remains acceptable taking into account the scatter observed in elementary pin pull-out tests. The model predicts a simultaneous disbonding of the three pins rows at the maximum applied load.

Pinned L-joint model under bending (folding) is quasi-identical to the one used for tension (see Figure Nº 22). The only difference comes from the fact that, for bending, the interface is considered to be delaminated only at the pinned area from the web to the beginning of the rib-foot. This delamination length can be justified by analysing bending test results where it has been observed that joint failure occurs when rib-foot brakes before the total delamination of the interface [START_REF] Vasquez | Identification et modelisation du comportement des structures composites assemblees par cloutage[END_REF].

Figure 24 shows load carried by the three pin rows as a function of the displace-ment of the load application point. On the same graphic, total joint carried load from tests and from the simulation are also plotted against load application point displacement. A good agreement between test and calculation can be observed even if the model doesn't represent delamination growth. Load carried by the three pins "row 1", "row 2" and "row 3" can be read over the right vertical axis whereas total load values correspond to the left vertical axis. According to the simulation, pins from row 1 (on the left on Figure 22) start disbonding when the applied load reaches 420N. Thus, this load level will correspond to the joint limit load as, for higher loads, pins on row 1 will carry load just by friction. Once the third row debonds at an applied load around 1100N, joint bending stiffness change and its integrity is only assured by the skin to the end of the rib-foot interface. Nevertheless, pins still carry some load by friction. From a sizing point of view, the failure of the joint can be consider occuring at 1100N even if total failure happens for a higher load level.

Regarding T-joint under bending, Figure 25 shows simulation results versus three test measures. Simulation curve matches globally well with experimental results. However, a difference on the slope change can be observed for an applied load of 150N. Indeed, the model shows two slope changes at 115N and 150N corresponding to the first and second pin row debonding. Tests have only shown one slope change for a mean load value slightly higher: 165N. The difference can be explained as the model considers a perfect geometry and identical pin pull-out laws. In reality, pins are not exactly over their theoretical position and small deviations can impact bending strength.

In this way, it has been shown that FE model proposed allows calculating pin joint strength under tension and bending loadings.

Shear.

T-joint pinned with 40 mm pins over three rows has been simulated with a 3Dmodel shown in Figure 26. Like for pinned L-joint under tension, only pins have been represented by non-linear springs at the web to skin interface. Material properties are the same as in previous models (Table 1). Boundary conditions are also clamping conditions over the skin 10 mm away from the web. Load application test rig has been modelled by rigid elements. Contact at the web to skin interface has been considered in order to avoid web elements to penetrate into the skin on the compression area.

Simulation results are compared to test measures on Figure 27. It must be reminded that pinned T-joints shear behaviour depends on relative pins angle against applied load. Thus, on the same graphic, the shear responses of specimens with 40 mm pins oriented "against" and "towards" de load are plotted. The model response shows stiffness slightly lower than the test measures. The difference could be due to the fact that the model does not represent the resin layer between web and skin. This layer could stiffen the specimen even if it delaminates during the test. The simulation stops suddenly when shear load on the most loaded pin reaches the pin shear strength. Model maximum applied load is coherent with test results.

Figure 28 shows axial and shear load carried by three pins as a function of the load application point displacement. "Pin 1" corresponds to a pin on the lowest side of the interface (lowest Z coordinate value) where tension load over the pins are higher due to the bending parasite moment. On the other side, "pin 6" is placed on the highest side of the interface, closer to the ball-joint used to introduce de load. On the graphic, the total applied load is also plotted in order to better understand the response of each pin compared to the global joint response. Load on pins can be read on the left vertical axis and joint global applied load over the right one. According to the simulation, the first pin to debond is "pin 1" over the lower part of the specimen for an applied load level of 2000N, a low value compared with joint shear strength. "Pin 6" is firstly loaded under compression but axial load changes it sign to tension load when the joint applied load reaches 2500N. Pin 6 debonds when applied load is 4000N. From that moment, pins carry tension load only by friction. At the same time, pins are also loaded under shear by forces increasing gradually with shear total applied load. Joint failure occurs when shear load over pins 6 and 3 reaches the shear strength (298 N). It can be observed that maximum shear pin load is reached almost at the same time over all the pins what explains the sudden specimen failure.

The simulation shows pins to be loaded mostly under tension. It could explain why the maximum applied load obtained from the model is in line with shear strength of specimens with pins oriented "towards" the load. Indeed, experimentally, as pins are inclined towards the applied load, they are more loaded in tension and it will produce a higher joint strength.

The proposed model allows representing only one of the configurations tested due to the particular behaviour of pinned joints under shear when pins are not perpendicular.

Numerical studies have allowed to represent T-joints pinned with pins oriented "against the load" by leading to a lower shear strength. However, these results have been obtained by modifying original single pin behaviour laws obtained in the first part of this paper. In particular, single pin response under combined loading (shear/tension, shear/compression) should be investigated in detail in order to feed a more capable finite elements model.

4-Conclusions.

For this study, pinned joints behaviour manufactured with EADS Innovation Works has been investigated. This technology allows manufacturing composite subassembly including skin/stiffener joints with a high integration level.

In this paper, experimental behaviour of pinned and non-pinned joints under static tension, shear and bending loading has been analysed. Two main joint families has been tested: L-joint (classical design) and T-joints (lower manufacturing cost consuming)

Globally, for all the configurations, pins increase joint strength strongly (more than 200% for L-joints under tension and bending). The benefit is lower under shear. The comparison between classical non pinned L-joint and pinned T-joint have shown that pinned pull-out and shear T-joint capabilities can be as good as non pinned L-joints but with a lower associated manufacturing cost. This is not the case for bending where Tjoint strength is low. Joint behaviour under shear is highly dependent on pins misalignment produced during the manufacturing process. This phenomenon could be avoided by using a pin pattern with alternative inclinations as proposed in [START_REF] Petiot | New challenges in structural design and analysis of composite structures[END_REF].

Pinned joint mechanical behaviour simulation has shown the validity of multi-level approach. A good agreement between test results and model prediction has been found in terms of tension and bending strength. Joints behaviour under shear is globally well captured by the model in case of pins oriented "towards the load". In order to obtain a more universal model, investigations on pins behaviour characterisation under combined loading should be done. From a numerical point of view, it could be interesting to switch from a model based on spring elements to a model using specific cohesive elements still to be developed. Indeed, cohesive elements allow an easier application on the simulation of big structures where spring-base models can be very hard to implement.

From a general point of view, this study represents the first stage on the under-standing of the mechanical behaviour of this new technology. Although first characterisations have been done here and even a model has been proposed, a significant work must be done still before the technology is mature enough to be used in an industrial project.

5-Acknowledgement.

Authors would like to thanks Thierry Vilain (Dassault Aviation and ALCAS WP3 leader) and Patrice Lefebure (EADS IW) for their support.

with 3 staggered

 3 rows clearly appreciated over the lower skin surface. It must be pointed-out that pin insertion produces a wavy effect in the highest plies of the skin over the joining line (figure nº 4). For L-joints, outer radius of the rib-foot is filled with a UD fillet along the width of the specimen (see Figure nº 4). Five joint configurations have been investigated: •Non pinned L-joints: type A •15 mm pinned L-joint: type B. •40 mm pinned L-joint: type C. •15 mm pinned T-joint: type D. •40 mm pinned T-joint: type E. Specimens are tested under pull-out, shear and bending. Three specimens are tested for each loading so 45 tests are performed in total. Boundary conditions for all the loadings and all configurations (except type B, D and E under tension) consist in clamping the skin 5 mm away from the web and the foot end. The load is applied over the web: tension and bending (Figure nº 6) and shear (figure nº 7). Shear test has been conceived to apply the load at the skin-to-web interface.

  Thus, failure load decrease to 801+/-76N, representing 10% increase from no-pinned Ljoints. Bending Load/displacement curves for T-joints pinned with 40 mm long pins are shown in Figure 15. A first linear load increase is observed until 165 N where a plateau appears. Visual analyses show that this happens when outer pins star disbonding from the skin and continue slipping from it. The load plateau extends until the displacement reaches 17 mm. At that moment, rotation angle at the web base is too high and pins fail by bending making load to drop-off. Bending failure load for each configuration is shown Figure Nº 16. Under bending load, pins length is an important parameter as only joints reinforced with 40 mm pins have shown a significant improvement. In that case, due to their insertion depth into the web, pins increase stiffness and also the strength. Pinned T-joints are not effective under bending due to their natural low mechanical strength. T-joints reinforced with 15 mm

FIGURESFigure N° 1 :

 1 FIGURES

Figure N° 2 :

 2 Figure N° 2: L-joints geometry.

Figure N° 9 :Figure N° 10 :Figure N° 11 :Figure N° 12 :Figure N° 13 :

 910111213 Figure N° 9: Pinned L-joints (40 mm pins) under tens ion): load/displacement curves and failure scenario.

Figure N° 14 :

 14 Figure N° 14: Failure initiation and final failure p attern for 15 mm pinned L-joint under bending.

Figure N° 15 :Figure N° 17 :

 1517 Figure N° 15: Load Displacement curves and failure sc enario of 40 mm pinned T-joints under bending

FinalFigure N° 20 :

 20 Figure N° 20: Shear maximum applied load for L and T -joints.

Figure N° 21 :Figure

 21 Figure N° 21: Modelling principle.