# A HIGHER-ORDER BEAM THEORY FOR COMPOSITE BOX BEAMS

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1 A HIGHER-ORDER BEAM THEORY FOR COMPOSITE BOX BEAMS A. Kroker, W. Becker TU Darmstadt, Department of Mechanical Engineering, Chair of Structural Mechanics Hochschulstr. 1, D Darmstadt, Germany SUMMARY A higher-order theory for composite box beams with a rectangular, closed cross-section is presented. The flanges and webs may have different layups. Starting from appropriate assumptions for the displacements, a set of differential equations is derived. The resulting coupled differential equation system can be solved in a closed-form analytical manner. Keywords: higher-order theory, laminated composite box beam, closed-form analysis, shear deformation, failure prediction INTRODUCTION For many years beams are very commonly used construction elements in civil and mechanical engineering. In the past they were primarily made of wood or steel. Over the last few decades composite materials are used in more and more technical applications in all fields of engineering and substantial progress has been made in manufacturing techniques. By using these new materials engineers especially have to consider reliability and durability problems. Therefore analytical and computational beam models are needed to predict the system behaviour. Today many theories for laminated plates exist in parallel, most of them for special cases only. Three main theories are very well established: The classical laminate plate theory (Kirchhoff kinematics), the first-order shear deformation theory (Reissner- Mindlin kinematics), and the third-order shear deformation laminate theory of Reddy [1]. In contrast to the first two mentioned theories, the theory of Reddy is also suitable for thicker plates because it takes shear deformation into account by using a cubic functional dependence on the thickness coordinate. In the present work the main ideas of Reddy s theory are transferred to laminated composite box beams with a single cell cross-section. However, instead of using a cubic formulation in thickness direction, higher-order functions in the in-plane direction are implemented. PROBLEM FORMULATION The beam theory is derived for rectangular, closed cross-sections. For that purpose the cross-section of the box beam is separated into four parts: two webs and two flanges.

2 All four parts consist of laminates, which may be different for the webs and the flanges. The laminates themselves are normally made of different fibre reinforced single layers with varying fibre angles from layer to layer. The aim is now to find an analytical model for describing the behaviour of the whole beam in detail. THEORY FORMULATION For each wall section, different representations for the displacements are chosen to describe the behaviour independently from each other. As it is known for thick laminates, warping effects can not be neglected for beams with high aspect ratios of height to width. Therefore higher-order theories are necessary. Basic Assumptions For the formulation of the beam theory three main assumptions are made for simplifying the equation system of the model, which are normally fulfilled in technical applications: The theory is developed only for symmetrical layups with balanced fibre angles. In this way there occurs no bending extension and no shear coupling in the laminates. The third assumption is that there is the same layup in both webs and both flanges. The flange layup may be different to the one in the webs. Several coordinate systems are used for the beam description. They are defined as follows (figure 1a). a) b) Figure 1: a) Global and local coordinate systems of the beam model, b) Shear flow distribution for a transverse load case In the beam centroid the global coordinate system is defined and labelled as x, y and z. There are also locally defined coordinate systems for each section wall separately. They have an additional index of the associated wall (i.e. 1, 2, 3 and 4). All x-coordinates

3 show always in the axial direction of the beam. The coordinate s follows the beam cross-section in the direction of the local y-axes. Kinematics The displacement field for each wall section is defined in terms of its local coordinate system. For a transverse loading case integration of the local Bernoulli beam results and FEM-results have shown, that there is predominantly linear shear flow distribution in the flanges and a quadratic one in the webs. To get this result from a first-order beam model, an integration of the derivative of the axial stress with respect to x across the cutting area of the cross-section has to be carried out. Starting from this shear flow distribution (figure1b) the shear strain field can be obtained by using Hooke s law in conjunction with the laminate stiffness matrix. The integration of the strain field yields the underlying displacement functions (eqs. 1-6). As a result the in-plane kinematics for the flanges are quadratic and for the webs of a cubic type - similar to the kinematics Reddy used in thickness direction for his third-order plate theory - as shown below. Because of symmetry reasons the equations are shown only for the top flange 1 and the left web 2. Flanges: Webs: Continuity (e.g. left top edge): The linear part of u f which means the function Ω x and the quadratic part (θ x ) of u w can be neglected because of symmetry reasons. An axial deflection of the complete beam cross-section will not be taken into account in this work. The unknown functions can be reduced to the following four functions of x by ensuring continuous displacements at the box edges (eq. 8): w 0 (x): vertical beam deflection; Φ x (x): inclination and λ x (x) warping mode of the webs; ψ x (x): warping mode of the flanges.

4 Strain Field Only moderate beam deflections will be investigated, as a result the strains will be small. The three in-plane strain components of the linearised Green-Lagrange s strain tensor are given by the following equations (9-14). Flanges: Webs: Stress Field The assumptions for the plane stress state are fulfilled in essence for sufficiently thin plates such as the beam walls. Therefore the stress field can be expressed for each single layer as a product of the common plane stress reduced stiffness coefficients ij and the strain field. For an orthotropic single layer with an arbitrary fibre angle the elasticity law in local wall coordinates can be written as follows (eq. 15): The behaviour of the laminate can be described in terms of Kirchhoff s laminate stiffness matrix by integration over all single layers in thickness direction.

5 Herein the strains are split up into terms which are constant or have a linear dependency on the thickness coordinate z (eq. 17). For laminates which have symmetrical layups with balanced fibre angles, as it was requested for this beam theory, the bending-extension coupling matrix B and the shear coupling coefficients A 16, A 26 are zero. With these assumptions the curvature of the wall laminates has no influence on the cross-sectional forces and the influence of the local z coordinate can be neglected in the displacement field. The out-of-plane bending moment of the flanges, which again is not taken into account, is of lower importance in comparison to the in-plane bending moment of the webs. In contrast to lower-order displacement approaches this higher-order theory makes it possible to obtain the shear force distribution directly without integration along the cross-section contour. Eventually it is to be stated that in the current setting there is no shear flow continuity condition at the edges implemented although if this would reduce the unknown functions by one. However this would complicate the definition of the boundary conditions. Equations of Motion With the presented displacement representation it is not possible to fulfil the equilibrium conditions in their strong form. Instead they can be met in a weak form by means of the principle of minimum total potential. The governing ODE-system i.e. the Euler- Lagrange equations and the boundary conditions for the static case can be obtained using the principle of virtual displacements (eq. 18). This principle states that the virtual strain energy equals the virtual work done by the applied forces. f stands for the volume forces of the body and t for the forces acting on the body surface. The equations for both webs and for both flanges can be summarised. The four unknown functions lead to the following linear differential system of four coupled equations (19-22) in terms of the unknowns of second order.

6 Boundary Conditions In the derivation of the equations of motion the terms of both webs and both flanges are summarised as stated before. The boundary conditions are expressed in terms of the top flanges and the left web. Four boundary conditions (eqs ) are required at each end of the beam in form of values for the introduced kinematical functions (i.e. the unknown functions) or by means of the corresponding forces and moments. The transverse force F is acting in global z-direction. M b is a bending moment about the global y-axis. The last two moments are the warping moments of the flanges and the webs. SOLUTION PROCEDURE By using an -ansatz (eq. 27) for all four unknown functions of x a solution can be found for the ODE-system in a completely closed-form analytical manner.

7 A substitution of equation (27) into the ODE-system (eq ) leads to a homogeneous equations system and the characteristic equation (eq. 28) is obtained. Its solution yields 8 eigenvalues K 1-8 (eq. 29) and eigenvectors E 1-8. The lengths of the eigenvectors still are open, which means that the constants D 1...D 8 are not yet determined. For the deformation functions w 0 (x), Φ x (x), λ x (x), ψ x (x) thus only the 8 independent constants D 1...D 8 remain to be identified from the boundary conditions. Herein the quantities c i are already known constants consisting of the beam geometry and material data. The extended terms are not shown here. RESULTS AND DISCUSSION Various layups and boundary conditions have been investigated in different examples for cantilever beams. For all following cases the beam is clamped at the left end for x=0 in all degrees of freedom. The right end (x=l) is welded to a rigid plate which is loaded by a bending moment and / or a transverse force. At this end of the beam there is no warping of the cross-section possible.

8 The boundary conditions for all following examples are set as stated in figure 2 and the equations (35) and (36). Q and M are the transverse force and the bending moment at the position x=0.5l. Figure 2: Boundary conditions for a cantilever beam For an assessment of the derived approach and for comparison reasons also detailed finite element analyses of the beam have been performed using ABAQUS S4 shell elements representing the beam walls. Different load cases and varying layups have been investigated. In figure 3 a few results of this new theory for the beam deflection are plotted over its axial length coordinate and a comparison is made with two finite element computations. The nodes - except the ones at the beam ends - of the first FE-model can move freely in all three directions whereas in the second model all nodes are fixed in the transverse y-direction. Figure3: Beam deflections for different layups and load cases

9 It can be seen that for all these cases there is very little difference between the FEresults and those of the analytical model. Only with the second layup, where webs and flanges are made of ±45 degree layers, a significant discrepancy of all three models is observed. The predicted deflection is too small. The analytical model can represent rather the results of the second FE-model well but not those of the first one. By blocking the transverse displacement, as this is not taken into account by the displacement field of the theory, the flanges cannot elongate or shrink in their normal way. This observation is reflected by the disagreement of the two FE-models. Figure 4 shows the cross-sectional shear force plotted along the coordinate s at different axial positions x of the beam. The distribution is presented for the beam with ±45 degree layers from above, where a big discrepancy between the analytical and the two FEmodels can be found. Figure 4: Cross-sectional shear force developing along axial beam axis Even if there is a larger variation in the beam deflection as shown in figure 3, it can be seen, that the influence on the distribution of the cross-sectional shear force is limited. A very good agreement is met at each axial position of the beam in comparison especially to the second FE-model where all nodes are fixed in y-direction. A Q-M-interaction plot for one cross-section of the beam can be created as follows. By variing the force F and the bending moment M b at the right beam end an arbitrary load vector can be set up for this cross-section. For the investigated cut an appropriate failure criterion e.g. Tsai-Wu, Puck etc. is calculated for each layer. The highest loaded layer restricts the norm of the load vector to a value where the beam is at its strength limit. This load combination is then plotted in form of a single point into a graph. Afterwards the procedure is repeated for several different load cases. The results are shown in figure 5. The previously described procedure is repeated for different laminates with varying fibre angels as stated in the figure legend whereas the thickness of each layer is held constant.

10 Figure 5: Q-M-interaction plot of one beam cross-section Six different configurations of fibre angles between 0 and 45 degree have been investigated, where the webs and the flanges have the same layup. The larger the angle the higher is the force resistance and the lower is the bending resistance of the beam. Up to a fibre angle of around 15 degree obviously there is only little change in bending resistance but the force resistance is nearly doubled. It is an advantage of the derived approach that it enables such an investigation without essential computing costs. This makes it suitable for efficient optimisation processes. Furthermore this load interaction plot can also be produced for a complete beam. CONCLUDING REMARKS This new beam theory has been developed for rectangular composite box beams to give engineers an analytical model for behaviour predictions and to speed up layer optimisation processes. Different examples for cantilever beams with varying layups were presented. The considered load cases were a combination of a bending moment and a transverse force. The obtained results showed a good agreement with comparative finite element computations and the warping deformation of the box beam cross-section is clearly revealed. Especially for those cases with a high shear force, the higher-order box beam theory is better than a lower-order approach. In contrast to the finite element analyses the derived beam solutions however are of a much higher computational efficiency and allow parameter studies without essential computational costs. There is also little effort needed to compute Q-M-interaction plots for any employed failure criterion and arbitrary layups. At last there is no meshing needed and the results have to be calculated only for points of interests and not at each mesh node, as it is in finite-element methods. ACKNOWLEDGEMENTS This work has been financially supported by Deutsche Forschungsgemeinschaft under grant no. BE 1090/22-1, which is gratefully acknowledged. References 1. Reddy, J.N. (2004): Mechanics of laminated composite plates and shells: theory and analysis. 2 nd ed. CRC Press, Boca Raton

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