Abstract
In the present paper, semi-analytical design sensitivity analysis of a morphing forward wing section modeled as a flexible multibody system is developed. The flexible multibody dynamics model consists of a flexible external wing skin and an actuation mechanism with rigid bodies. The floating frame of reference formulation is used for the formulation of the equations of motion, while time integration is carried out with the generalized-\(\alpha\) method and Newton–Raphson iterations for the nonlinear analysis. These steps are considered in both the primal analysis and the developed efficient design sensitivity analysis. The calculation of the design sensitivities is based on direct differentiation and carried out with a semi-analytical approach. The responses of interest include the deviations of the wing profile from the targeted morphed geometry and the resulting stresses in the wing skin. The design variables considered in this work are geometric parameters, material parameters and loading parameters. The calculations with the shown method provide reliable and accurate results combined with high computational efficiency. The visualization of the computed sensitivity values gives an easy interpretation of the sensitivities and facilitates the understanding of the design engineer on how to change the design variables to improve the design of the system. The introduced sensitivity analysis enables future investigations, including the application of the method to gradient-based design optimization and uncertainty analysis.
Similar content being viewed by others
Avoid common mistakes on your manuscript.
1 Introduction
Flexible multibody dynamics (FMBD) is a powerful tool for the analysis of mechanical systems that undergo dynamic loading and large displacements and rotations. In addition to the system responses, their design sensitivities are of great use in the design, design optimization and sensitivity analysis, amongst other uses. In this work, we develop the sensitivity analysis for FMBD and apply it to the design of a morphing wing of a high-performance sailplane.
High-performance sailplanes (also known as gliders) are unpowered aircraft and thus have a large envelope of operating speeds. A typical competition flight comprises of two typical flight phases: Low-speed circling phases in thermals, followed by mid- to high-speed interthermal flight. Fixed-geometry aircraft are designed to have a performance compromise in this wide speed range. Form-variable – or morphing – wings have shown to increase performance, particularly when the forward wing section is morphed in combination with a conventional trailing-edge flap. Morphing wings can achieve an increase in speed up to \(20~\frac{\text{km}}{\text{h}}\) with equal glide ratio compared to a conventional sailplane with a camber changing trailing-edge flap [1–3].
The morphing forward wing section is found in literature modeled both with hinged mechanisms [1, 4] and with specially designed hingeless compliant mechanisms [3, 5]. FMBD is applied here to a hinged mechanism that includes both rigid and flexible components. Via flexible multibody simulation, the mechanism is analysed to consider large rigid-body displacements and rotations as well as flexible deformations and stresses. Flexible multibody dynamics enables the dynamic analysis of the mechanism in contrast to coupled simulations based on rigid multibody dynamics and (quasi-) static finite-element analysis [1, 3–5]. These values can then be integrated into a design optimization routine. The methods developed here are implemented to extend an in-house flexible multibody simulation code SiMuLi, which carries out both the simulation and sensitivity analysis of a morphing forward wing section with the goal to be integrated in a future work in a design optimization framework.
The approach for FMBD in this work is a further development of that developed and introduced in the application to cleaning mechanisms of Tyrolean weirs, intake systems of small Alpine hydroelectric plants [6–8]. This work specifically builds upon [9] to include the semi-analytical sensitivities of the invariants to allow for a multibody simulation and its sensitivity analysis, decoupled from the finite-element solver after the initial modeling. This is enabled by the analytical derivatives w.r.t. the design variables of the system parameters that depend on the finite-element model, see [10] for the analytical expressions. This semi-analytical approach thus takes the analytical differentiation to a further layer of the partial derivatives leading to significant computational efficiency improvements to the method shown in [9].
The centerpiece of this methodology is the efficient analytical sensitivity analysis. This can then be used i.a. in gradient-based design optimization of a wide range of flexible mechanisms. The developed methodology is shown here for the simulation of a morphing forward wing section concept consisting of a flexible external wing skin and an actuation mechanism with rigid bodies. Actuating the mechanism leads to the deformation of the external wing skin from the low-lift (high-speed) profile configuration to the high-lift (low-speed) profile configuration. Particular attention is paid to the sensitivity analysis, including the computation, analysis and interpretation of the sensitivity values. The system responses considered here are the stress and the geometric deviations from the target geometry. These are differentiated with respect to the design variables, which include geometric parameters, material parameters and loading parameters. The method is general and not limited by those design variables explored here. Future application fields of the developed sensitivity method include gradient-based design optimization and uncertainty analysis.
2 Flexible multibody dynamics
Flexible multibody simulation is comprised of three components [8, 11]: governing equations, time integration and nonlinear solver. The main simulation to calculate the responses is referred as primal analysis to differentiate from the design sensitivity analysis, in which the sensitivity of the responses with respect to design variables are calculated. The simulation routine is shown as a flowchart in Fig. 1 and includes the components, which are introduced in § 2 for the primal analysis and in § 3 for the design sensitivity analysis.
2.1 Governing equation
The governing equations of flexible multibody systems are given by the equations of motion and the kinematic constraint equations, which lead to a set of index-3 differential–algebraic equations given by
where \(\underline{q}\) is the vector of generalized positions, \(\underline{\lambda}\) is the vector of Lagrangian multipliers of the kinematic constraints, \(\underline{\underline{m}}\) is the mass matrix, \(\underline{\underline{d}}\) is the damping matrix, \(\underline{\underline{k}}\) is the stiffness matrix, \(\text{$\underline{\varPhi }$}\) is the vector of kinematic constraints, \(\underline{\underline{\mathrm{J}}}{}_{\varPhi }\) is the Jacobian matrix of the constraints (i.e. the partial derivative of the constraints with respect to position), \(\underline{F}_{\text{ext}}\) is the external force vector and \(\underline{F}_{v}\) is the quadratic velocity force vector. Overdots represent the first \(\dot{\left (\cdot \right )}\) and second \(\ddot{\left (\cdot \right )}\) temporal derivatives. Symbols with single underline \(\underline{\left (\cdot \right )}\) denote vectors, symbols with double underline \(\underline{\underline{\left (\cdot \right )}}\) denote two-dimensional matrices, symbols with triple underline denote three-dimensional terms, symbols with quadruple underline denote four-dimensional terms and those without underlines are scalars. Further, symbols with an overline \(\overline{\left (\cdot \right )}\) are expressed in floating coordinates of the body reference frame. To reduce numerical problems related to index-3 differential–algebraic equations [12, 13], the kinematic constraint equations are differentiated twice with respect to time to obtain a system of index-1 differential–algebraic equations expressed in matrix form by
where \(\underline{R}\) is the residual vector and \(\underline{F}_{c}\) is the right-hand side of the kinematic constraints in acceleration form. As index-1 differential–algebraic equations are affected by drift, Baumgarte stabilization [14] is used as described by [15].
FMBD is reviewed by [13, 16–18]. Available flexible multibody formulations include the absolute nodal coordinate formulation (ANCF) [19], the absolute coordinate formulation (ACF) [20], the equivalent rigid link system (ERLS) [21, 22], geometrically exact beam formulation [23, 24] and the floating frame of reference formulation (FFRF) [25, 26]. The lattermost is used here, which is suitable for beam, shell and solid FE elements and can be used with a linear-elastic material model as well as with geometric or material nonlinearities [13, 27].
In the following derivation of the equations of motion for FFRF, the index of each body is omitted for a simpler notation. The equations are shown for the general three-dimensional case. The generalized positions in terms of FFRF are paramterized by the position \(\underline{r}\) and orientation \(\underline{\beta }\) of the reference frame and flexible deformations in local coordinates \(\overline{\underline{q}}_{f}\), see Fig. 2,
The continuous position of a material point \(P\) on a flexible body can be seen in Fig. 2 and is described in global coordinates as
with the transformation matrix \(\underline{\underline{A}}\) and the local position vector \(\overline{\underline{u}}_{P}\) that is decomposed by the undeformed term \(\overline{\underline{u}}_{P,o}\) and the deformed term \(\overline{\underline{u}}_{P,f}\).
In this implementation, FFRF is used in combination with finite-element analysis. Each flexible body is discretized and the flexible deformations in local coordinates are given by
with the matrix of shape functions \(\overline{\underline{\underline{S}}}\) expressed in floating coordinates and the nodal deformations \(\overline{\underline{q}}_{f}\) of the finite-element mesh.
A linear-elastic material model is used, which leads to a FFRF stiffness matrix \(\underline{\underline{k}}\) that is linear,
with the finite-element stiffness matrix \(\underline{\underline{k}}{}_{ff}\). In contrast, the FFRF mass matrix \(\underline{\underline{m}}\) is highly nonlinear and defined by
with the volume \(V\), the density \(\rho \), the identity matrix \(\underline{\underline{e}}\) and the matrix \(\underline{\underline{B}}\). It should be noted that in the system matrices of FFRF, \(t\) denotes the translatory degrees of freedom of the body frame, \(r\) the rotational degrees of freedom of the body frame and \(f\) the flexible degrees of freedom. Accordingly, the index \(ff\) represents the system matrices from the finite-element model. The matrix \(\underline{\underline{B}}\) is defined by
where \(\tilde{\underline{\overline{\underline{u}}}}{}_{P}\) is the skew-symmetric matrix of positions in local coordinates and the matrix \(\overline{\underline{\underline{G}}}\) relates the angular velocity vector of the floating frame \(\overline{\underline{\omega}}\) and the time derivatives of the orientation parameters \(\dot{\underline{\beta }}\),
The quadratic velocity vector is highly nonlinear and defined by
In FFRF, the external forces are assembled in a vector as
The vectors and matrices of each body are assembled to obtain the system vectors and matrices. The governing equations are to be fulfilled and hence are solved at all time steps.
2.2 Time integration
The system of differential–algebraic equations shown in § 2 is solved in time with the generalized-\(\alpha\) time integration method originally introduced by [28] and implemented as a predictor–corrector scheme, as described in [8, 11]. Generalized-\(\alpha\) time integration is based on Newmark’s equations [29],
with intermediate approximations of the generalized-\(\alpha\) method,
The resulting effective system of equations for FMBD is
where the components are defined as
The position and velocity values are updated with the solved accelerations and continue to the next time step until the termination time is reached.
2.3 Nonlinear solver
The residual equations with generalised-\(\alpha\) time integration are
Newton–Raphson iterations (i.e. exact Newton method) are used to consider nonlinearity of the system by solving
with the Jacobians w.r.t. acceleration \(\ddot{\mathrm{J}}\) and Lagrangian multipliers \(\overset{\lambda}{\mathrm{J}}\) giving the expressions for the Jacobians of the residual by
The components of the Jacobian of the residual are
For a complete derivation of the terms above using generalized-\(\alpha\) time integration including time integration, the readers are referred to [11].
3 Design sensitivity analysis
Although referred to as design sensitivities, the sensitivities of the system responses with respect to parameters are useful in a wide range of applications. Such applications include design optimization (including sizing, shape and topology optimization), uncertainty analysis as well as the direct application of the sensitivities to assess robustness and parameter influence.
This paper investigates discrete sensitivity analysis, also referred as discretize-then-differentiate, where the differentiation w.r.t. the design variables is carried out on discretized governing equations [30]. These include temporal discretization into time steps of numerical time integration and spatial discretization into finite elements of the flexible bodies. This is in contrast to continuous sensitivity analysis, also known as the variational approach or differentiate-then-discretize, where continuum governing equations are differentiated before discretization.
To avoid the high computational effort and lack of precision of numerical sensitivity analysis, analytical methods including the direct differentiation method [13, 31] and the adjoint variable method [32–34] are preferred for the sensitivity analysis of flexible multibody systems. When the number of design variables is higher than the number of optimization responses including objective and constraint functions, the adjoint variable method is generally more efficient. In the reverse case, where the number of design variables is less than the number of optimization responses, the direct differentiation method is more suitable. The direct differential method is chosen here and integrated into the primal solve routine to avoid the solving backwards in time as needed with an adjoint variable methodology (therefore often referred to the backward method). Related challenges to the adjoint variable method include the handling of time-dependent system matrices (save and reload or recalculate), especially when using variable time steps and a required special solver for the adjoint system.
The design sensitivity analysis is performed here with direct differentiation using a semi-analytical approach. The differentiation is carried out through all three steps of the calculation routine: the governing equation, the sensitivity analysis and the nonlinear solver (cf. Fig. 1). The introduced method is similar to the staggered corrector method introduced by [35] with the backward difference formula. This method originates from the staggered direct scheme, where an iteration loop of the primal analysis is followed by the iteration loop of the sensitivities for each time step [36]. Efficiency improvements compared to the staggered direct scheme have been obtained by the simultaneous corrector scheme, where primal and sensitivity analysis are solved simultaneously with one iteration loop [37]. With the staggered corrector method as used here, further efficiency improvements are obtained using two sequential iteration loops for primal and sensitivity analysis and by reducing the updates of the residual Jacobians [35].
3.1 Governing equation
The direct differentiation of the governing equations of the primal analysis (3) results in turn with the governing equations for the sensitivity analysis,
where \(\nabla \) is the partial derivative with respect to the design variables,
It should be noted that (38) is of the same form as (3) and accordingly the same solving routine can be used. The pseudo load \(\underline{\underline{F}}{}_{\text{pseudo}}\) is comprised of the partial derivatives of the system parameters with respect to the design variables,
This method is shown in [9], where the partial derivatives of (40) are approximated numerically with forward differencing,
This thus results in a semi-analytical approach, where the governing equations are derived analytically, but the partial derivatives of these equations are computed numerically. In contrast, the partial derivatives of those parameters dependent on the finite-element model \(\underline{\underline{m}}\), \(\underline{\underline{k}}\) and \(\underline{F}_{v}\) are differentiated analytically by a further layer based on an invariant approach, see [10]. This is done to avoid loading the finite-element model at each time step and leads to a further decrease in computation effort compared to the method shown in [9].
3.2 Time integration
The sensitivity analysis with time integration is analogous to the time integration of the primal analysis. For the generalised-\(\alpha\) method, we utilize a predictor–corrector scheme with Newmark’s equations and intermediate approximations [8, 11]. The resulting effective sensitivity system is
where the pseudo load case is
The sensitivity values of position and velocity are carried out from the sensitivity values of acceleration analogously to the primal analysis.
3.3 Nonlinear solver
The Jacobian matrix of the sensitivity analysis simplifies to the Jacobian matrix of the primal analysis [8, 11], allowing to reuse the already computed Jacobian of the primal analysis for an efficient computation method of the sensitivity analysis,
with the Jacobian with respect to acceleration sensitivity \(\nabla \ddot{\mathrm{J}}\) and with respect to the Lagrangian multipliers sensitivity \(\nabla \overset{\lambda}{\mathrm{J}}\). With this simplification, the Jacobian of the sensitivity analysis is reduced from four dimensions to two. This saves excessive memory usage and thus represents the key to efficient sensitivity analysis of multibody systems with direct differentiation.
4 Morphing wing model
In this work, the developed method is applied to the design of a morphing wing concept developed in [2, 3, 38, 39]. Specifically, the geometry of the morphing wing sailplane for both undeformed (high-speed) and morphed (low-speed) configurations is introduced in [2, 40]. The airfoil shapes are designed using a numerical optimization approach where aerodynamic lift and drag are considered [2, 40]. The geometries of the airfoils along with the aircraft for which they were designed are shown in Fig. 3. A morphing concept shows a significant performance advantage over conventional sailplanes with a camber changing flap: 20% higher lift coefficients are achieved at equal drag coefficients compared to conventional camber changing trailing edge flap laminar airfoils, thus allowing to reduce the wing area and decrease profile drag by the same ratio. This results in \(20~\frac{\text{km}}{\text{h}}\) higher interthermal dash speed.
In [5], the morphing actuation of this concept is achieved using compliant mechanisms designed with topology optimization. This concept uses a stack of six individual compliant mechanisms and is shown in Fig. 4. In contrast to a compliant concept, we will address a traditional hinged mechanism to drive the deformation of the wing in this work.
The forward wing section is modeled as flexible multibody system using FFRF. Figure 5a shows the planar system actuated with a conventional hinged mechanism and the two-dimensional case here is the simplification of the three-dimensional case shown in § 2. The flexible outer shell (body 1), i.e. wing skin, is modeled as flexible body using planar Euler–Bernoulli beams with a FE mesh consisting of 31 nodes and 30 elements. For this system, a wing section with the length of 1000 mm and a wall thickness of 1.5 mm is considered. On the upper end of the forward wing section (point \(\mathrm{F}\)), the wing skin is fully constrained in all three degrees of freedom, i.e. \(x\), \(y\) and \(\theta \). On the lower end of the forward wing section (point \(\mathrm{G}\)), the wing skin is constrained in two degrees of freedom \(y\) and \(\theta \), allowing for free motion in \(x\).
The morphing forward wing section is actuated by a hinged mechanism consisting of four rigid bodies 2, 3, 4 and 5. The motion of these bodies is constrained by six joints including five revolute joints in the points \(\mathrm{A}\), \(\mathrm{B}\), \(\mathrm{C}\), \(\mathrm{D}\) and \(\mathrm{E}\) and one rigid joint between the bodies 2 and 4 in the point \(\mathrm{B}\). For actuation, a torque is applied to body 2 on point \(\mathrm{A}\). The torque value is zero at the initial time and increases with the curve in Fig. 6 to a maximum value of \(4.2~\text{N}\cdot \text{m}\).
A generic aluminum alloy is used with a density \(\rho \) of \(2.7\times 10^{-9}~\text{t}/{\text{mm}^{3}}\), a Young modulus \(E\) of 70 000 MPa and a Poisson ratio \(\nu \) of \(0.35~{\left [-\right ]}\). The material is assumed to be isotropic and to remain in the linear elastic regime.
The transient solution of the flexible multibody system is obtained with generalised-\(\alpha\) time integration and Newton–Raphson iterations as nonlinear solver. The system is solved at each time step \(\varDelta t=0.05~\text{s}\). The deformation of the external wing skin and the motion of the rigid bodies of the multibody system from the low-lift (high-speed, undeformed) configuration to the high-lift (low-speed, morphed) configuration is shown in Fig. 5b.
The key for a proper design of the morphing forward wing section is the interaction between the actuation mechanism and the flexible external wing skin. Firstly, it is important that the deformation of the morphed forward wing section approximates the target wing profile as closely as possible. Secondly, the material limits of the wing skin must not be exceeded. Flexible multibody dynamics is applied to the morphing forward wing section to assess both morphed shape as well as the stresses. Figure 7 shows the stress distribution caused by tension or compression and bending in the wing skin during the motion of the mechanism from the low-lift (high-speed, undeformed) configuration to the high-lift (low-speed, morphed) configuration. The stress values increase during the motion and the highest values are found in morphed configuration close to the constraints of the wing skin (points \(\mathrm{F}\) and \(\mathrm{G}\)).
The design is limited by the maximum stress of all elements and all time steps. To achieve this with a continuous, differentiable function, we approximate the maximum value via constraint aggregation (also called constraint lumping) with the Kreisselmeier–Steinhauser function \(\mathcal{F}_{\mathrm{KS}}\) [41],
The modified Kreisselmeier–Steinhauser function is used here to avoid numerical overflow and is defined by [42],
where \(n_{f}\) is the number of function values and \(\rho \) is the aggregation parameter (also referred to as penalty), for which a value of 200 is used here.
Specifically in this case, we approximate the maximum stress by
with the stress \(\sigma \), the number of time steps \(n_{t}\) and the number of elements \(n_{e}\). Table 1 shows a comparison of the primal analysis of the design described above, where the maximum stress is given by 111.586975 MPa. When using a small value for the aggregation parameter, there is a small error between the true maximum stress and the approximated value. Increasing the value of the aggregation parameter, the error reduces and with large aggregation parameters with \(\rho \geq 50\), the approximated value is exactly the true maximum stress and no relative error can be detected.
The undeformed high-speed configuration, the target morphed low-speed configuration and the achieved deformation is seen in Fig. 8a. The error between the target morphed geometry and the deformation field achieved is calculated for each node. For the \(k\)th node, the distance between the simulated node position and the target profile in morphed configuration is
where \(x_{k,\mathrm{sim}}\) and \(y_{k,\mathrm{sim}}\) are the simulated node coordinates, \(x_{k,\mathrm{tar}}\) and \(y_{k,\mathrm{tar}}\) are the target node coordinates and \(n_{N}\) is the number of nodes. Figure 8b shows the error between the simulated profile and the target profile in morphed configuration. The largest error is 1.185 mm and is located at approximated the front quarter point on the lower side of the wing profile.
We quantify the deviation of the achieved deformation from the target shape via root-mean-square error,
which is given by 0.6348 mm for the given design. The deviations from the target wing profile with a maximum value of 1.124 mm and a root-mean-square error of 0.6225 mm are deemed to be acceptable accuracy, considering the chord length of 550 mm, the wing height of 67.5 mm and the chord length of the forward wing section of 137.5 mm. As the shape of the wing profile is crucial for the aerodynamics and therefore the performance of sailplanes, it is desired to minimise the deviations.
The design variables for this investigation include the thickness of the external wing skin \(t_{w}\), the torque value applied to actuate the mechanism \(M_{t}\), the coordinates \(x_{A}\) and \(y_{A}\) of the torque application point \(\mathrm{A}\) and the Young’s modulus \(E\). The method shown in § 3 is applied to compute the sensitivity values and the chain rule is applied to all operations.
Figure 9 shows the design sensitivities of the stress during the motion of the flexible multibody system.Footnote 1 During the motion, the highest design sensitivities for all design variables are found at the end of the simulation in morphed configuration. The location of the highest values on the wing profile for the sensitivities with respect to the \(x\)-coordinate \(x_{A}\) of the torque application point are found on the tip of the wing profile and at the revolute joint in point \(\mathrm{C,}\) where the actuation mechanism is coupled to the external wing skin, while for all other design variables the location of the highest sensitivity values is close to the constraints of the external wing skin (points \(\mathrm{F}\) and \(\mathrm{G}\)). While the maximum stress values in the primal analysis are found on the lower constraint (point \(\mathrm{G}\)), the highest stress sensitivity values in the sensitivity analysis are found on the upper constraint (point \(\mathrm{F}\)). Comparing the Figs. 7 and 9 shows that a reduction of the stress \(\sigma \) can be obtained by increasing the thickness of the external wing skin \(t_{w}\), the \(x\)-coordinate \(x_{A}\) of the torque application point and the Young’s modulus \(E\) or by reducing the actuation torque \(M_{t}\) and the \(y\)-coordinate \(y_{A}\) of the torque application point.
Figure 9 shows the design sensitivities of the error between the simulated profile and the target profile in morphed configuration. Thereby stands out that especially in the nose of the forward wing section, the sign of the design sensitivities on the lower side is opposite to the sign of the design sensitivities on the upper side. The comparison of the error values on Fig. 8b with the error sensitivities on Fig. 10 shows how the design variables should be modified for reducing the error. The maximum values of the geometric deviation between the simulated profile and the target profile are found on the lower side of the forward wing section and a reduction of the error is therefore obtained by increasing the thickness of the external wing skin \(t_{w}\), the \(x\)-coordinate \(x_{A}\) of the torque application point and the Young’s modulus \(E\) or by reducing the actuation torque \(M_{t}\) and the \(y\)-coordinate \(y_{A}\) of the torque application point.
Table 2 gives the results of the sensitivity analysis of the maximum stress and the root-mean-square error. The differentiation of the Kreisselmeier–Steinhauser function is defined by
where the function of interest \(f\) is the stress \(\sigma \). The direct differentiation of the error is
and the direct differentiation of the root-mean-square error is
It can be seen that Eq. (53) is infinity when the error at one point is exactly zero. Where this can have adverse effects, e.g. gradient-based optimization, mean square error, which does not have this issue, should be used.
The results in Table 2 show consistency with the results in Fig. 9 and Fig. 10. For the maximum stress, the highest value is found on the constraint at the lower end of the forward wing section (point \(\mathrm{G}\)). The design sensitivities of Fig. 9 in this location correspond exactly to the approximated Kreisselmeier–Steinhauser sensitivities reported in Table 2. In the computation of the root-mean-square error, all nodes are considered by the function and therefore it is not possible to perform a similar comparison. However, the order of magnitude of the values from Fig. 10 and Table 2 is the same.
The introduced semi-analytical method for the sensitivity analysis of flexible multibody dynamics is validated with a numerical sensitivity analysis with forward differences using a relative perturbation equal to \(10^{-6}\times \) the value of the design variable itself. The values coincide with the relative difference shown in Table 2.
Table 3 shows a comparison of the computational effort of both methods.Footnote 2 This comparison highlights the high computational efficiency of the introduced semi-analytical sensitivity method w.r.t. finite differencing. With numerical sensitivity analysis, the computational effort of the numerical sensitivity analysis is given by the sum of \(n_{\mathsf{x}}+1\) primal evaluations with a mean computation time of 62 s. With the semi-analytical sensitivity analysis, the primal analysis and the sensitivity analysis result in computational effort that is less than twice (\(1.403\times \)) the computational effort of one primal evaluation. The high efficiency of the semi-analytical method compared to the numerical method is already seen with the five design variables used here. This difference becomes even more evident for increasing numbers of design variables. The main reasons for the speed-up are the decoupling of the sensitivity analysis from the FE model explained in § 3.1 and the simplification shown in § 3.3 that allows to reuse the two-dimensional Jacobian of the primal analysis in the sensitivity analysis.
5 Conclusion
The present work introduces an efficient method for the computation of design sensitivities of flexible multibody dynamics using floating frame of reference formulation with a semi-analytical direct differentiation approach with generalized-\(\alpha\) time integration and Newton–Raphson iterations for a morphing wing concept. This concept for a morphing forward wing section in combination with a trailing flap for low-lift (high-speed) and high-lift (low-speed) configurations increases the performance in the wide speed range of high-performance sailplanes.
The morphing forward wing section is modeled with flexible elements for the external wing skin and rigid bodies for the actuation mechanism. As system responses, the error between the simulated forward wing profile and the target profile in morphed configuration and the stress values in the wing skin are investigated. The maximum stress is considered as material limit and knockdown factor, though the consideration of the fatigue (including sensitivities) would be of great practical use in future work. Special interest is devoted to the design sensitivities of the system responses.
A semi-analytical sensitivity approach is implemented for its computational efficiency and accuracy. A direct differentiation is chosen to avoid the reverse time integration needed for an adjoint variable methodology. That said, the adjoint variable method exhibits great computational effort advantages with high numbers of design variables, e.g. in topology optimization, and further investigation and comparison in the regard is warranted.
The high efficiency of the method is guaranteed by the reuse of the primal Jacobian for the sensitivity analysis and the decoupling of the sensitivity analysis from the FE model by the analytical differentiation of the system parameters w.r.t. the design variables to a further layer with an invariant-based approach. The design sensitivities calculated include those with respect to geometric properties, material properties and the position and value of the actuator load.
The present study is an integral part of a design optimization framework to optimally design a morphing wing and its actuation mechanism, which is being developed. In this context, the shown sensitivity analysis is the basis for efficient design optimization and uncertainty analysis specifically of the morphing wing and its actuation mechanism and flexible multibody dynamics in general.
Notes
Units of sensitivity results are not reduced to demonstrate the physical and engineering meaning of these values.
Computations are performed on a PC with Intel Core i7-8700 CPU @ 3.20 GHz × 12 and 32 GB RAM.
References
Sinapius, M., Monner, H.P., Kintscher, M., Riemenschneider, J.: DLR’s morphing wing activities within the European network. Proc. IUTAM 10, 416–426 (2014)
Achleitner, J., Rohde-Brandenburger, K., Rogalla von Bieberstein, P., Sturm, F., Hornung, M.: Aerodynamic design of a morphing wing sailplane. In: AIAA Aviation 2019 Forum. American Institute of Aeronautics and Astronautics, Reston, Virginia (2019)
Sturm, F., Achleitner, J., Jocham, K., Hornung, M.: Studies of anisotropic wing shell concepts for a sailplane with a morphing forward wing section. In: AIAA Aviation 2019 Forum. American Institute of Aeronautics and Astronautics, Reston, Virginia (2019)
Rudenko, A., Hannig, A., Monner, H.P., Horst, P.: Extremely deformable morphing leading edge: optimization, design and structural testing. J. Intell. Mater. Syst. Struct. 29(5), 764–773 (2017)
Reinisch, J., Wehrle, E., Achleitner, J.: Multiresolution topology optimization of large-deformation path-generation compliant mechanisms with stress constraints. Appl. Sci. 11(6), 2479 (2021)
Gufler, V.: Multibody dynamics and optimal design of a Tyrolean weir cleaning mechanism. Master thesis, Free University of Bozen-Bolzano (2019). Advisors: E. J. Wehrle, R. Vidoni
Gufler, V., Wehrle, E., Vidoni, R.: Multiphysical design optimization of multibody systems: application to a Tyrolean weir cleaning mechanism. In: Advances in Italian Mechanism Science, pp. 459–467. Springer, Berlin (2021)
Wehrle, E., Gufler, V.: Lightweight engineering design of nonlinear dynamic systems with gradient-based structural design optimization. In: Proceedings of the Munich Symposium on Lightweight Design 2020, pp. 44–57. Springer, Berlin (2021)
Gufler, V., Wehrle, E., Achleitner, J., Vidoni, R.: Flexible multibody dynamics and sensitivity analysis in the design of a morphing leading edge for high-performance sailplanes. In: Proceedings of the 10th ECCOMAS Thematic Conference on MULTIBODY DYNAMICS. Budapest University of Technology and Economics, Budapest, Hungary (2021), online
Gufler, V., Zwölfer, A., Wehrle, E.: Analytical derivatives of the floating frame of reference formulation. Multibody Syst. Dyn. (2022). https://doi.org/10.1007/s11044-022-09858-5
Wehrle, E., Gufler, V.: Analytical sensitivity analysis of dynamic problems with direct differentiation of generalized-\(\alpha\)time integration. Submitted, 10.31224/osf.io/2mb6y (preprint)
Zhu, Y.: Sensitivity analysis and optimization of multibody systems. PhD thesis, Virginia Polytechnic Institute and State University (2014)
Gufler, V., Wehrle, E., Zwölfer, A.: A review of flexible multibody dynamics for gradient-based design optimization. Multibody Syst. Dyn. 53, 79–409 (2021).
Baumgarte, J.: Stabilization of constraints and integrals of motion in dynamical systems. Comput. Methods Appl. Mech. Eng. 1(1), 1–16 (1972)
Gufler, V., Wehrle, E., Vidoni, R.: Sensitivity analysis of flexible multibody dynamics with generalized-\(\upalpha \) time integration and Baumgarte stabilization: a study on numerical stability. In: Mechanisms and Machine Science, pp. 147–155. Springer, Berlin (2022)
Shabana, A.A.: Flexible multibody dynamics: review of past and recent developments. Multibody Syst. Dyn. 1(2), 189–222 (1997)
Wasfy, T.M., Noor, A.K.: Computational strategies for flexible multibody systems. Appl. Mech. Rev. 56(6), 553–613 (2003)
Dwivedy, S.K., Eberhard, P.: Dynamic analysis of flexible manipulators, a literature review. Mech. Mach. Theory 41(7), 749–777 (2006)
Gerstmayr, J., Sugiyama, H., Mikkola, A.: Review on the absolute nodal coordinate formulation for large deformation analysis of multibody systems. J. Comput. Nonlinear Dyn. 8(3), 031016 (2013).
Gerstmayr, J.: The absolute coordinate formulation with elasto-plastic deformations. Multibody Syst. Dyn. 12(4), 363–383 (2004)
Vidoni, R., Gasparetto, A., Giovagnoni, M.: A method for modeling three-dimensional flexible mechanisms based on an equivalent rigid-link system. J. Vib. Control 20(4), 483–500 (2014)
Vidoni, R., Gallina, P., Boscariol, P., Gasparetto, A., Giovagnoni, M.: Modeling the vibration of spatial flexible mechanisms through an equivalent rigid-link system/component mode synthesis approach. J. Vib. Control 23(12), 1890–1907 (2017)
Simo, J.: A finite strain beam formulation. The three-dimensional dynamic problem. Part I. Comput. Methods Appl. Mech. Eng. 49(1), 55–70 (1985)
Bauchau, O.A.: Flexible Multibody Dynamics. Springer, Berlin (2011)
Shabana, A.A.: Dynamics of Multibody Systems, 5th edn. Cambridge University Press, Cambridge (2020)
Zwölfer, A., Gerstmayr, J.: The nodal-based floating frame of reference formulation with modal reduction. Acta Mech. 232, 835–851 (2021)
Dibold, M., Gerstmayr, J., Irschik, H.: A detailed comparison of the absolute nodal coordinate and the floating frame of reference formulation in deformable multibody systems. J. Comput. Nonlinear Dyn. 4(2), 021006 (2009)
Chung, J., Hulbert, G.: A time integration algorithm for structural dynamics with improved numerical dissipation: the generalized-\(\alpha \) method. J. Appl. Mech. 60, 371–375 (1993)
Newmark, N.M.: A method of computation for structural dynamics. J. Eng. Mech. Div. 85(3) (1959). https://doi.org/10.1061/JMCEA3.0000098
Haftka, R.T., Gürdal, Z.: Elements of Structural Optimization, 3rd edn. Kluwer, Dordrecht, Netherlands (1992)
Zhang, M., Peng, H., Song, N.: Semi-analytical sensitivity analysis approach for fully coupled optimization of flexible multibody systems. Mech. Mach. Theory 159, 104256 (2021)
Boopathy, K., Kennedy, G.: Adjoint-based derivative evaluation methods for flexible multibody systems with rotorcraft applications. In: 55th AIAA Aerospace Sciences Meeting. American Institute of Aeronautics and Astronautics, Grapevine, Texas (2017)
Held, A., Knüfer, S., Seifried, R.: Structural sensitivity analysis of flexible multibody systems modeled with the floating frame of reference approach using the adjoint variable method. Multibody Syst. Dyn. 40, 287–302 (2017)
Nejat, A.A., Moghadasi, A., Held, A.: Adjoint sensitivity analysis of flexible multibody systems in differential-algebraic form. Comput. Struct. 228, 106–148 (2020)
Feehery, W.F., Tolsma, J.E., Barton, P.I.: Efficient sensitivity analysis of large-scale differential-algebraic systems. Appl. Numer. Math. 25(1), 41–54 (1997)
Caracotsios, M., Stewart, W.E.: Sensitivity analysis of initial value problems with mixed ODEs and algebraic equations. Comput. Chem. Eng. 9(4), 359–365 (1985)
Maly, T., Petzold, L.R.: Numerical methods and software for sensitivity analysis of differential-algebraic systems. Appl. Numer. Math. 20(1–2), 57–79 (1996)
Sturm, F., Illenberger, G., Techmer, D., Hornung, M.: Static aeroelastic tailoring of a high-aspect-ratio-wing for a sailplane with a forward morphing wing section. In: 32nd Congress of the International Council of the Aeroanautical Sciences. Deutsche Gesellschaft für Luft- und Raumfahrt, Shanghai, China (2021)
Sturm, F., Hornung, M.: Morphing shell design of a sailplane with a morphing forward wing section. In: XXXV OSTIV Congress – Congress Program and Proceedings. TU Braunschweig – Niedersächsisches Forschungszentrum für Luftfahrt, online (2021)
Achleitner, J., Rohde-Brandenburger, K., Hornung, M.: Airfoil optimization with CST parameterization for (un-)conventional demands. In: XXXIV OSTIV Congress, pp. 117–120 (2018)
Kreisselmeier, G., Steinhauser, R.: Systematic control design by optimizing a vector performance index. In: International Federation of Active Controls Symposium on Computer-Aided Design of Control Systems, vol. 12, pp. 113–117. Elsevier, Amsterdam (1979)
Martins, J.R.R.A., Poon, N.M.K.: On structural optimization using constraint aggregation. In: 6th World Congress on Structural and Multidisciplinary Optimization (2005)
Acknowledgements
This work is supported by the project CRC 2017 TN2091 doloMULTI Design of Lightweight Optimized structures and systems under multidisciplinary considerations through integration of multibody dynamics in a multiphysics framework funded by the Free University of Bozen-Bolzano. Further support was provided in the framework MILAN – Morphing wings for sailplanes, funded by the German Federal Ministry for Economic Affairs and Energy under the grant of the German Federal Aviation Research Program (Luftfahrtforschungsprogramm, LuFo) V-3.
Funding
Open access funding provided by the Free University of Bozen-Bolzano within the CRUI-CARE Agreement.
Author information
Authors and Affiliations
Corresponding author
Ethics declarations
Competing Interests
The authors declare that they have no conflict of interest.
Additional information
Publisher’s Note
Springer Nature remains neutral with regard to jurisdictional claims in published maps and institutional affiliations.
Rights and permissions
Open Access This article is licensed under a Creative Commons Attribution 4.0 International License, which permits use, sharing, adaptation, distribution and reproduction in any medium or format, as long as you give appropriate credit to the original author(s) and the source, provide a link to the Creative Commons licence, and indicate if changes were made. The images or other third party material in this article are included in the article’s Creative Commons licence, unless indicated otherwise in a credit line to the material. If material is not included in the article’s Creative Commons licence and your intended use is not permitted by statutory regulation or exceeds the permitted use, you will need to obtain permission directly from the copyright holder. To view a copy of this licence, visit http://creativecommons.org/licenses/by/4.0/.
About this article
Cite this article
Gufler, V., Wehrle, E., Achleitner, J. et al. A semi-analytical approach to sensitivity analysis with flexible multibody dynamics of a morphing forward wing section. Multibody Syst Dyn 58, 1–20 (2023). https://doi.org/10.1007/s11044-023-09886-9
Received:
Accepted:
Published:
Issue Date:
DOI: https://doi.org/10.1007/s11044-023-09886-9