geometrically nonlinear analysis of beam structures via...
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Research ArticleGeometrically Nonlinear Analysis of Beam Structures viaHierarchical One-Dimensional Finite Elements
Y. Hui,1,2,3 G. De Pietro,2,3 G. Giunta ,2 S. Belouettar,2 H. Hu,1
E. Carrera ,3 and A. Pagani 3
1School of Civil Engineering, Wuhan University, 8 South Road of East Lake, 430072Wuhan, China2Luxembourg Institute of Science and Technology, 5, avenue des Hauts-Fourneaux, 4362 Esch-sur-Alzette, Luxembourg3Politecnico di Torino, C.so Duca degli Abruzzi 24, 10129 Turin, Italy
Correspondence should be addressed to G. Giunta; [email protected]
Received 16 July 2018; Accepted 12 November 2018; Published 27 November 2018
Academic Editor: Xiao-Qiao He
Copyright Β© 2018 Y. Hui et al. This is an open access article distributed under the Creative Commons Attribution License, whichpermits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.
The formulation of a family of advanced one-dimensional finite elements for the geometrically nonlinear static analysis of beam-like structures is presented in this paper. The kinematic field is axiomatically assumed along the thickness direction via a UnifiedFormulation (UF). The approximation order of the displacement field along the thickness is a free parameter that leads to severalhigher-order beam elements accounting for shear deformation and local cross-sectional warping.Thenumber of nodes per elementis also a free parameter. The tangent stiffness matrix of the elements is obtained via the Principle of Virtual Displacements. Atotal Lagrangian approach is used and Newton-Raphson method is employed in order to solve the nonlinear governing equations.Locking phenomena are tackled by means of a Mixed Interpolation of Tensorial Components (MITC), which can also significantlyenhance the convergence performance of the proposed elements. Numerical investigations for large displacements, large rotations,and small strains analysis of beam-like structures for different boundary conditions and slenderness ratios are carried out, showingthat UF-based higher-order beam theories can lead to a more efficient prediction of the displacement and stress fields, whencompared to two-dimensional finite element solutions.
1. Introduction
Many structural elements, such as aircraft wings, rotorblades, robot arms, or structures in civil construction, canbe idealised as beams. Furthermore, the hypothesis that theunstrained and deformed configurations are coincident atequilibrium is often not true and geometrical nonlinearitiescannot be neglected. Engineering fields such as aeronautics,space, and automotive need more and more accurate modelssince an accurate prediction of the mechanics of beamsplays a paramount role in their optimal design. Therefore,geometrically nonlinear modelling of beams represents animportant and up-to-date research topic.
A general overview on linear and nonlinear structuralmechanics can be found in Nayfeh and Pai [2]. Nonlin-ear structural analysis via finite elements was thoroughlydiscussed in Crisfield [3] and Bathe [4]. Hodges et al. [5]provided a variational-asymptotical method that allowed
obtaining an asymptotically correct strain energy for theapproximation of stiffness coefficients for the predictionof geometrically nonlinear behaviour of composite beams.A paralinear isoparametric element for the geometricallynonlinear analysis of elastic two-dimensional bodies waspresented by Wood and Zienkiewicz [6]. Newton-Raphsonmethod was used in order to solve the nonlinear equilibriumequations. Surana [7] provided a geometrically nonlinearformulation for two-dimensional curved beams. A totalLagrangian approach was used and the beam element wasderived using linear, paralinear, and cubic-linear plane stresselements. Dufva et al. [8] presented a two-dimensional shear-deformable beam element for large deformation analyses.Cubic interpolation was used for the rotation angles causedby bending and linear interpolation polynomials were usedfor the shear deformations. The absolute nodal coordinateformulation was used for the finite element discretizationalong the beam axis. Further works on large rotations and
HindawiMathematical Problems in EngineeringVolume 2018, Article ID 4821385, 22 pageshttps://doi.org/10.1155/2018/4821385
2 Mathematical Problems in Engineering
large displacements analysis of shear-deformable beams byusing an absolute nodal coordinate formulation accountingfor a nonrigid cross-sectional kinematics were carried out byDufva et al. [9] and Omar and Sharana [10]. A geometric andmaterial nonlinear analysis was carried out by Chan [11] forbeam-columns and frames. An optimum nonlinear solutiontechnique within the Newton-Raphson scheme was obtainedby minimizing the residual displacements. The evaluation ofgeometrically exact beam theories andmodels based on a sec-ond order approximation of finite rotations for the bucklingand postbuckling analysis of beam structures was carried outby Ibrahimbegovic et al. [12]. Yu et al. [13] developed a gen-eralised Vlasov theory for composite beams by means of thevariational-asymptotic method.The geometrically nonlinear,three-dimensional elasticity problem was split into a linear,two-dimensional cross-sectional analysis and a nonlinearone-dimensional beam analysis. A novel two-dimensionalfinite element solution for the nonlinear buckling and wrin-kling of sandwich plates has been developed by Yu et al. [14].Kirchoff βs theory was adopted for the kinematics of the skins,whereas a higher-order displacement field was consideredfor the core mechanics. The nonlinear governing equationswere derived by the Principle of Virtual Displacements andsolved via the Asymptotic Numerical Method (ANM). Basedon this work, Huang et al. [15] proposed a more effective two-dimensional model by using the technique of slowly variableFourier coefficients. Garcia-Vallejo et al. [16] introduced anew absolute nodal coordinate finite element together witha reduced integration procedure in order to mitigate thelocking phenomena in dynamic structural problems.
A static analysis of beam-like structures via a hierarchicalone-dimensional approach is addressed in this paper. Thekinematics along the thickness is axiomatically assumed via aUnified Formulation (UF). UF had been previously proposedfor plate and shell structures; see Carrera [17]; and it has beenapplied to the study of beams byCarrera et al. [18] andCarreraand Giunta [19]. Thanks to this approach, the derivationof a family of higher-order beam theories is made formallygeneral regardless of the through-the-thickness approximat-ing functions and the approximation order. UF has beenextended to large deflections and postbuckling analysis ofbeam structures in recent works by Pagani and Carrera[20, 21], where the capability of such approach to investigateglobal and local deformations in solid and thin-walled beamstructures was demonstrated by using Lagrange polynomialsas approximating functions for the cross-section kinematicswithin a layer-wise approach. An elastoplastic analysis viaUF-based one-dimensional finite elements has been carriedout by Carrera et al. [22], showing that such formulation canlead to a 3D-like accuracy in terms of displacements andstresses for compact and thin-walled structures subjected tolocalized loadings. Applications of the UF approach for theinvestigation of multiphysics problems, vibration analyses,and static structural analyses of composite structures can befound in Carrera et al. [23], Biscani et al. [24], Koutsawa et al.[25], and Giunta et al. [26β29]. In this study, a large displace-ments, large rotations, and small strains analysis of beam-like structures are carried out using UF-based finite elementsand Taylor expansion of the displacement field along the
z, w
x, u
l
β
Figure 1: Beam geometry and reference system.
thickness using an equivalent single-layer formulation. Theapproximation order is a free parameter and, therefore,several kinematic models can be straightforwardly obtained,accounting for nonclassical effects, such as transverse shearand local cross-sectional warping. The number of nodes perelement is also not a priori fixed. Linear, quadratic, and cubicelements are formulated. The tangent stiffness matrix of theelement is derived from the weak form of the governingequations obtained via the Principle of Virtual Displacements(PVD). A total Lagrangian (TL) formulation is used andthe global problem is solved by classical Newton-Raphsonprediction/correction method. As far as the stress predictionis concerned, once the second Piola-Kirchoff stresses havebeen obtained by the TL formulation, they are transformedinto the trueCauchy stresses to have a direct comparisonwithresults coming fromupdated Lagrangian formulations imple-mented in the commercial software ANSYS. As opposed tothe Lagrange layer-wise approximation [20, 21], in both linearand nonlinear analyses, the use of Taylor polynomials allowsan enrichment of the cross-sectional kinematics by simplyincreasing the order N and with no need for additional cross-sectional nodes. This feature makes Taylor-based refinedmodels particularly suitable for the analysis of multilayeredstructures in the framework of an equivalent single-layerapproach that will be presented in a future work. As furthernovelties with respect to [20, 21], the correction of shearand membrane locking phenomena in the nonlinear regimevia the MITC method has been introduced, allowing animproved convergence performance of the proposed finiteelements in slender structures. Finally, a detailed descrip-tion of the stress analysis under large displacements, notoften encountered in the literature, has been also provided,together with the discussion of some limitations of theproposed formulation with respect to finite elements withlarge strains capabilities.
2. Preliminaries
A beam is a structure whose axial extension (π) is predom-inant with respect to any other dimension orthogonal to it.The cross-section (Ξ© = β Γ π) is defined by intersectingthe beam with planes orthogonal to its axis. Figure 1 presentsthe beam geometry and the reference system, being β andπ, respectively, the beamβs thickness and width. A fixedCartesian reference system is adopted. The π₯ coordinate iscoincident with the axis of the beam and it is bounded suchthat 0 β€ π₯ β€ π, whereas the π¦- and π§-axis are two orthogonaldirections laying on Ξ©. The displacement field in a two-dimensional approach is
uπ (π₯, π§) = {π’ (π₯, π§) π€ (π₯, π§)} (1)
Mathematical Problems in Engineering 3
where π’ and π€ are the displacement components alongthe π₯- and π§-axis, respectively. Superscript βπβ representsthe transposition operator. For the sake of convenience, thedisplacements gradient vector π is introduced:
π = {π’,π₯ π’,π§ π€,π₯ π€,π§} (2)
Subscripts βπ₯β and βπ§β, when preceded by comma, representderivation versus the corresponding spatial coordinate.
Geometrical nonlinearity is accounted for in a Green-Lagrange sense. Large displacements and rotations are, there-fore, considered. The strain vector E is
πΈπ₯π₯ = π’,π₯ + 12 (π’2,π₯ + π€2,π₯)πΈπ§π§ = π€,π§ + 12 (π’2,π§ + π€2,π§)πΈπ₯π§ = π’,π§ + π€,π₯ + π’,π₯π’,π§ + π€,π₯π€,π§
(3)
Equation (3) can be written in the following matrix form:
E = [H + 12A (π)] π (4)
where
Eπ = {πΈπ₯π₯ πΈπ§π§ πΈπ₯π§} (5)
H = [[[
1 0 0 00 0 0 10 1 1 0
]]]
(6)
A (π) = [[[
π’,π₯ 0 π€,π₯ 00 π’,π§ 0 π€,π§π’,π§ π’,π₯ π€,π§ π€,π₯
]]]
(7)
A virtual variation of the strain vector cam be written as (seeCrisfield [3])
πΏE = πΏ {[H + 12A (π)] π} = [H + A (π)] πΏπ (8)
where πΏ stands for the virtual variation operator.The vectorial form of second Piola-Kirchhoff βs stress
tensor S is
Sπ = {ππ₯π₯ ππ§π§ ππ₯π§} (9)
The material is supposed to withstand small strains. Hookeβslaw is, therefore, considered:
S = QE (10)
In the case of an anisotropic material, the reduced materialstiffness matrix Q reads
Q = [[[
π11 π13 π15π13 π33 π35π15 π35 π55
]]]
(11)
Coefficients πππ are not reported here for the sake of brevity.They can be found in Reddy [30]. The Cauchy stress tensor πcan be derived from the deformation tensor F and the Piola-Kirchoff stress tensor through the following relation:
π = 1π½F[ππ₯π₯ ππ₯π§ππ₯π§ ππ§π§]F
π (12)
where
F = [1 + π’,π₯ π’,π§π€,π₯ 1 + π€,π§] (13)
and J = det(F) is the determinant of F. The weak form of thegoverning equations is obtained by means of the Principle ofVirtual Displacement:
πΏL = πΏLπππ‘ β πΏLππ₯π‘ = 0 (14)
whereL is the total work andLπππ‘ the internal one:
πΏLπππ‘ = β«π0
πΏEπS ππ (15)
π0 is the volume of the reference undeformed configuration.Lππ₯π‘ is the work done by the external forces. An infinitesimalvariation of the total work reads
π (πΏL) = β«π0
[πΏEππS + π (πΏEπ) S] ππ (16)
After few manipulations (see Crisfield [3]), (16) can berewritten in the following form:
π (πΏL) = β«π0
[πΏEπQπE + πΏππSππ] ππ (17)
where S β R4Γ4 is
S =[[[[[[
ππ₯π₯ ππ₯π§ 0 0ππ₯π§ ππ§π§ 0 00 0 ππ₯π₯ ππ₯π§0 0 ππ₯π§ ππ§π§
]]]]]]
(18)
The variation of the virtual work is finally written in terms ofthe actual and virtual variation of the gradient vector:
π (πΏL)= β«π0
πΏππ {[Hπ + Aπ (π)]Q [H + A (π)] + S} ππππ (19)
3. Hierarchical Beam Elements
3.1. Kinematic and Finite Element Approximations. Withinthe assumed Unified Formulation and the finite elementframework, the displacement components are approximated
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along the beam thickness via the base functions πΉπ(π§) andalong the axis by the shape functions ππ(π₯):π’ (π₯, π§) = πΉπ (π§)ππ (π₯) ππ’πππ€ (π₯, π§) = πΉπ (π§)ππ (π₯) ππ€ππ
with π = 1, 2, . . . , ππ’, π = 1, 2, . . . ,πππ(20)
where ππππ : π = π’,π€ are the nodal unknowns. Einsteinβscompact notation is used in (20): a repeated index implicitlyimplies summation over its variation range.ππ’ is the numberof terms accounted in the through-the-thickness expansionand it is arbitrary. Index π varies over the element number ofnodes πππ and it is also a free parameter of the formulation.Linear, quadratic, and cubic elements along the beam axisare considered. These elements are addressed by βB2β, βB3β,and βB4β, respectively. The finite element shape functionsapproximate the displacements along the beam axis in aπΆ0 sense up to an order ππ β 1. For the sake of brevity,these functions are not reported here, but they can be foundin Bathe [4]. Taylor polynomials are used as the class ofexpansion functions πΉπ(π§). The generic explicit form of thedisplacement field expanded viaπ-order Taylor polynomialsis given by
π’π₯ = π’π₯1 + π’π₯2π§ + π’π₯3π§2 + β β β + π’π₯(π+1)π§ππ’π§ = π’π§1 + π’π§2π§ + π’π§3π§2 + β β β + π’π§(π+1)π§π
(21)
π is the order of the approximating polynomials along thethickness and it is a free parameter of the formulation. Bythis approach, several displacement-based theories and finiteelements accounting for nonclassical effects are straightfor-wardly derived. By replacing (20) within (2), the kinematicand finite element approximation of the displacements gra-dient vector readsπ = {πΉπππ,π₯ππ’ππ πΉπ,π§ππππ’ππ πΉπππ,π₯ππ€ππ πΉπ,π§ππππ€ππ}= Gππqππ
(22)
whereGππ β R4Γ2 and qππ β R2Γ1 are
Gππ =[[[[[[
πΉπππ,π₯ 0πΉπ,π§ππ 00 πΉπππ,π₯0 πΉπ,π§ππ
]]]]]]
(23)
andqπππ = {ππ’ππ ππ€ππ} (24)
3.2. Tangent Stiffness Matrix. Once (22) is replaced within(19), the variation of the total virtual work readsπ (πΏLπ)
= πΏqπππ β«ππ0
Gπππ {[Hπ + Aπ (π)]Q [H + A (π)] + S}β Gπππππqππ = πΏqπππ (Kππππππ + Kππ‘1ππππ + Kππ‘2ππππ) πqππ
(25)
where the superscript βeβ refers to the considered element,ππ0 = ππ β ππ β βπ is the element volume at the referenceunstrained configuration, and Kππππππ K
ππ‘1ππππ K
ππ‘2ππππ β R2Γ2 are
the βfundamental nucleiβ of the linear, initial-displacement,and geometric contributions to the tangent stiffness matrix:
Kππππππ = β«ππ0
GπππHπQHGππππ
Kππ‘1ππππ = β«ππ0
Gπππ [HπQA + AπQ (H + A)]GππππKππ‘2ππππ = β«
ππ0
GπππSGππππ(26)
The nuclei are very general regardless of the approximationorder π over the thickness, the class of approximatingfunctions πΉπ, and the number of nodes per elementπππ alongthe beam axis; see Carrera et al. [18]. Their explicit form canbe found in the Appendix. Once the approximation order andthe number of nodes per element are fixed, the element tan-gent stiffness matrix is obtained straightforwardly via sum-mation of the previous nucleus corresponding to each term ofthe expansion. Finally, the nonlinear system is solved via theclassical Newton-Raphson prediction/correction method.
3.3. Shear and Membrane Locking: MITC Beam Elements.In the geometrically nonlinear analysis of straight beams,the displacement components are coupled by the quadraticterms in the geometric relations; see (3). Therefore, mem-brane as well as shear locking phenomena will degrade theelement performance and need to be mitigated, especiallywhen slender structures and low-order shape functions areconsidered (see Reddy [31] and Malkus and Hughes [32] formore details). In this study, locking phenomena are overcomevia the MITC method (see Bathe et al. [33β35]), consisting inthe following interpolation of all the strain components alongthe beam element axis:
πΈπ₯π₯ = πππΈππ₯π₯πΈπ§π§ = πππΈπzπ§πΈπ₯π§ = πππΈππ₯π§
(27)
where π denotes an implicit summation and varies from1 to πππ β 1. πΈππ₯π₯, πΈππ§π§, and πΈππ₯π§ are the strain componentscoming from the geometrical relations in (3) evaluated at theπ-th tying point πππ and ππ are the assumed interpolatingfunctions. Their expressions as functions of the natural beamelement coordinate π β [β1, 1] can be found in Carrera etal. [36] and they are reported below. For linear elements, theinterpolation is reduced to a point evaluation, since
π1 = 1ππ1 = 0 (28)
Mathematical Problems in Engineering 5
For quadratic elements, the assumed interpolating functionsand tying points are
π1 = β12β3(π β 1β3)π2 = 12β3(π + 1β3)ππ1 = β 1β3ππ2 = 1β3
(29)
And for cubic elements
π1 = 56π(π β β35)
π2 = β53 (π β β35)(π + β35)
π3 = 56π(π + β35)
ππ1 = ββ35ππ2 = 0ππ3 = β35
(30)
4. Numerical Results
The beam support is [0, π] Γ [ββ/2, β/2] Γ [βπ/2, π/2]. Thecross-section is square with β = π = 1 m. Slender(π/β = 100) and short beams (π/β = 10) are investigated.Cantilever, doubly clamped, and simply supported (hinged-hinged) beams made of aluminium (πΈ = 75 GPa and ] =0.33) are considered. A concentrated load ππ§ is applied at(π₯/π = 1, π§/β = 0) for the cantilever case and at (π₯/π =1/2, π§/β = 0) for doubly clamped and simply supportedboundary conditions. A dimensionless load factor π =ππ§π2/πΈπΌ is used, πΌ being the moment of inertia of the beamcross-section. Both displacement and stress values are givenwith respect to the initial fixed coordinate system.
Results for a plane stress, large displacements, largerotations, and small strains analysis provided by the proposedfamily of one-dimensional finite elements are compared withtwo-dimensional finite elements based on a total Lagrangianformulation and small strains hypothesis, referred to asβFEM 2D TLβ (see Hu et al. [37]). Reference solutions fromthe available literature as well as classical one-dimensionalcorotational ANSYS finite elements βBeam3β, with bothEuler-Bernoulli (EBT) and Timoshenko (TBT) kinematics,are considered for comparison and validation purposes.Results given by two-dimensional large strains ANSYS finiteelements βPlane183β based on an updated Lagrangian formu-lation are also provided as a further assessment. About the
computational costs, in order to be able to predict an accuratestress field in both short and slender beams, the most refinedmodel used in the following numerical investigations for theproposed one-dimensional formulation is given by a mesh of121 nodes and beam theoryπ = 5, corresponding to 1.5 β 103degrees of freedom (DOFs), whereas a mesh of 240 Γ 24elements was used for 2D FEM solutions, corresponding to3.6 β 104 DOFs. It should be noticed that the computationaladvantage coming from theUF approach is evenmore signifi-cant in nonlinear analyses when compared to linear analyses,since an iterative solution procedure is required and, there-fore, a computational gain is obtained at every solution step.
4.1. Locking Assessment. In order to correct the shear andmembrane locking phenomena affecting nonlinear one-dimensional elements, MITC method was adopted. Figure 2shows the effectiveness of the MITC B2 elements in predict-ing the normalised displacement οΏ½οΏ½π§ = π’π§/π’Cubicπ§ for increas-ing slenderness ratios π/β, where π’Cubicπ§ is the convergedsolution obtained with 40 B4 elements. It can be noticed thatthe locking correction strategy is effective regardless the beamtheory order π and the considered boundary conditions.Due to the presence of membrane locking, simply supportedbeams are the most critical case among those investigated,as far as element performance is concerned. Figure 3 showsthat MITC correction in slender beams can significantlyreduce the number of nodes needed for convergence. Theconverged reference solution π’Cubicπ§ is here obtained with140 B4 elements. The improvement is even more significantwhen lower-order shape functions are used, such as in linearand quadratic beam elements. It should be also noticedthat, unlike a classical displacement-based finite element, anelement adoptingMITC correction strategy no longer assuresa monotonic convergence βfrom belowβ, as shown in Laulusaet al. [38]. Following the previous convergence analyses, 121nodes and MITC B4 elements are used for the plot results,whereas 121 nodes and MITC B2, B3, and B4 elements arecompared in the table results at a fixed load parameter.
4.2. Cantilever Beam. A slender cantilever beam (π/β = 100)is first considered in order to validate the model towards clas-sical reference beam solutions. Dimensionless displacementsπ’π = π’π/π, Cauchy stresses πππ = πππ(2πΌ/ππ§πβ), and thicknessοΏ½οΏ½ = π§/β are considered. Figures 4 and 5 show the evolutionof the displacement components with the load parameter.The displacement οΏ½οΏ½π₯ is evaluated at (π, β/2), whereas οΏ½οΏ½π§at (π, ββ/2). For the case of slender beam, the referencesolution based on EBT kinematics can accurately predict thenonlinear deformation and it matches the higher-order beamtheories as well as the two-dimensional FEM results. On theother hand, if the slenderness ratio is reduced, the sheardeformation effects as well as local cross-sectional warpingbecome relevant and at least a beam theory with order πequal to 2 should be used for an accurate displacementprediction, as shown in Figures 6 and 7. A more detailednumerical comparison is given in Table 1, showing that beamtheories with π β₯ 2 can reduce the error given by TBTfrom 5.5% to 0.7%, when compared to 2D FEM solutions. As
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MITC B2B2
N = 2N = 5
100 500 10l/h
0.3
1
οΌ―οΌ΄
Figure 2: Locking correction via MITC method for linear elements and different beam theories, doubly clamped beam, π = 2, οΏ½οΏ½π§ evaluatedat (π/2, ββ/2).
B4MITC B4
0.6
0.7
0.8
0.9
1
1.1
1.2
1.3
1.4
οΌ―οΌ΄
61 109 301 205 157 13 253οΌοΌ¨
Figure 3: Convergence analysis of the normalised displacement οΏ½οΏ½π§ versus the total number of nodes ππ via cubic finite elements B4 andMITC B4, simply supported beam (π/β = 100 and π = 3), οΏ½οΏ½π§ evaluated at (π/2, β/2).
far as stress prediction is concerned, Figures 8β10 show thethrough-the-thickness profile of the stress components forthe short beam case. Stress components are given in the initialfixed reference system. A small difference can be noticedbetween the large strains 2D FEM solution βPlane183β andthe small strains βFEM 2D TLβ. Furthermore, the higher-order beam theories match the 2D small strains solution, withrelative differences being lower than 0.6% for π β₯ 3 and B4
elements, as shown in Table 2. Unlike classical theories, theproposed higher-order models can predict global as well aslocalized displacements and stresses, even in the proximity ofboundary conditions and throughout the nonlinear regime,as shown in Figures 26β30.
4.3. Doubly Clamped Beam. Clamped-clamped boundaryconditions are considered. The evaluation point for οΏ½οΏ½π₯ is
Mathematical Problems in Engineering 7
EBT [38]PLANE183FEM 2D TL
N=2N=3N=5
0
2
4
6
8
10
β4 β2β3β5 β1 0β610 Γ οΌ―οΌ²
Figure 4: Load factor π versus dimensionless displacement οΏ½οΏ½π₯ for a slender cantilever beam.
0 1 2 3 4 5 6 7 8 9
EBT [38]PLANE183FEM 2D TL
N=2N=3N=5
10 Γ οΌ―οΌ΄
0
2
4
6
8
10
Figure 5: Load factor π versus dimensionless displacement οΏ½οΏ½π§ for a slender cantilever beam.
(π/4, β/2), whereas οΏ½οΏ½π§ is evaluated at (π/2, ββ/2). Figures 11and 12 show the load-displacement curves for a slender beam,whereas the short beam case is shown in Figures 13 and14. A significant difference between large and small strainshypothesis can be noticed in this latter case. Similarly tothe cantilever case, Table 3 shows that higher-order beamtheories can improve the accuracy with respect to the 2D
FEM small strains solution from 4.6% (TBT) to 0.3% (π β₯ 3),at worse. Figures 15β17 as well as Table 4 show that higher-order beam theories are required in order to accuratelypredict the stress profile in a thick beam, being the relativeerror given by a 2-nd order beam theory about 27.8% in theworst case and the one given by theories with π β₯ 3 lowerthan 0.8%.
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EBT [38]PLANE183FEM 2D TL
N=2N=3N=5
0
2
4
6
8
10
β6 β5 β4 β3 β2 β1 0β710 Γ οΌ―οΌ²
Figure 6: Load factor π versus dimensionless displacement οΏ½οΏ½π₯ for a short cantilever beam.
0 1 2 3 4 5 6 7 8 9
EBT [38]PLANE183FEM 2D TL
N=2N=3N=5
10 Γ οΌ―οΌ΄
0
2
4
6
8
10
Figure 7: Load factor π versus dimensionless displacement οΏ½οΏ½π§ for a short cantilever beam.
4.4. Simply Supported Beam. Simply supported beams areconsidered.π’π₯ and οΏ½οΏ½π§ are evaluated at (0, ββ/2) and (π/2, β/2),respectively. The load-displacement curves are presented inFigures 18 and 19, whereas Figures 20β22 show the through-the-thickness profile of the stress components at the fixedload parameter π = 6.03. The same considerations of theprevious sections also apply to this case: the displacementprediction can be improved from an error of 3.7% for a 2nd
order theory to 0.8% for a 5th order theory, as shown inTable 5. Similarly for the stresses given in Table 6, π = 2and B4 elements lead to a relative difference with respect toβFEM 2DTLβ of about 2.3% for ππ₯π₯, 60.5% for οΏ½οΏ½π§π§, and 35.2%for ππ₯π§, whereas the errors given by a theory π = 5 andB4 finite elements reduce to about 0.3%. The capability ofthe proposed formulation for an accurate stress predictionis preserved along the full deformation path, as shown in
Mathematical Problems in Engineering 9
PLANE183FEM 2D TLN=2
N=3N=5
β0.5
β0.25
0
0.25
0.5
οΌ΄
0β1β2 1 2 3β310 Γ οΌ²οΌ²
Figure 8: Dimensionless Cauchy stress οΏ½οΏ½π₯π₯ along the thickness coordinate at π₯ = π/4 and π = 5.20 for a short cantilever beam.
PLANE183FEM 2D TLN=2
N=3N=5
β0.5
β0.25
0
0.25
0.5
οΌ΄
0.5β0.5 1.5β1.5 1β1 2 010 Γ οΌ΄οΌ΄
Figure 9: Dimensionless Cauchy stress οΏ½οΏ½π§π§ along the thickness coordinate at π₯ = π/4 and π = 5.20 for a short cantilever beam.
Figures 23β25, where the stress profile given by the π = 5model and the results obtained by the 2D FEM solution arepresented for different load factors π.5. Conclusions
A family of refined one-dimensional finite elements derivedthrough a Unified Formulation of the displacement field has
been proposed for the geometrically nonlinear analysis ofbeam-like structures. Slender as well as short beams havebeen investigated for different boundary conditions. MITCmethod has been adopted in order to tackle the shear andmembrane locking phenomena and improve the convergenceperformance of the proposed elements. The capability of UF-based structural theories to accurately yet efficiently predictboth the displacement and the stress fields in the nonlinear
10 Mathematical Problems in Engineering
PLANE183FEM 2D TLN=2
N=3N=5
1β1 0.5 2β2 β0.5 1.5β1.5 2.5 010 Γ οΌ²οΌ΄
β0.5
β0.25
0
0.25
0.5
οΌ΄
Figure 10: Dimensionless Cauchy stress οΏ½οΏ½π₯π§ along the thickness coordinate at π₯ = π/4 and π = 5.20 for a short cantilever beam.
PLANE183FEM 2D TLN=2
N=3N=5
β3 β2.5 β2 β1.5 β1 β0.5 0β3.5
104Γ οΌ―οΌ²
0
5
10
15
20
Figure 11: Load factor π versus dimensionless displacement οΏ½οΏ½π₯ for a slender doubly clamped beam.
regime via a one-dimensional model has been demonstratedin this work. As far as computational costs are concerned, theuse of UF-based one-dimensional finite elements can save atleast one order of magnitude in terms of DOFs, with respectto two-dimensional elements. The extension of the proposedformulation based on Taylor polynomials for an equivalentsingle-layer analysis of composite beam structures will bepresented in a future work.
Appendix
Fundamental Nuclei of the TangentStiffness Matrix
The components of the linear stiffness matrix Kππππππ are
πΎπππ₯π₯ππππ = π½11πππΌπ,π₯π,π₯ + π½55π,π§π,π§πΌππ + π½15π,π§ππΌππ,π₯ + π½15ππ,π§πΌπ,π₯π
Mathematical Problems in Engineering 11
PLANE183FEM 2D TLN=2
N=3N=5
0
5
10
15
20
1.5 1 0.5 2 2.5 0
102Γ οΌ―z
Figure 12: Load factor π versus dimensionless displacement οΏ½οΏ½π§ for a slender doubly clamped beam.
PLANE183FEM 2D TLN=2
N=3N=5
β0.8 β0.6 β0.4 β0.2 0β1
102Γ οΌ―x
0
2
4
6
8
10
12
14
16
Figure 13: Load factor π versus dimensionless displacement οΏ½οΏ½π₯ for a short doubly clamped beam.
Kπππ₯π§ππππ = π½13ππ,π§πΌπ,π₯π + π½15πππΌπ,π₯π,π₯ + π½35π,π§π,π§πΌππ + π½55π,π§ππΌππ,π₯πΎπππ§π₯ππππ = π½13π,π§ππΌππ,π₯ + π½15πππΌπ,π₯π,π₯ + π½35π,π§π,π§πΌππ + π½55ππ,π§πΌπ,π₯ππΎπππ§π§ππππ = π½33π,π§π,π§πΌππ + π½55πππΌπ,π₯π,π₯ + π½35ππ,π§πΌπ,π₯π + π½35π,π§ππΌππ,π₯
(A.1)
The generic term π½πβπ(,π§)π(,π§) is a cross-section moment:
π½πβπ(,π§)π(,π§) = β«Ξ©π=βπΓππ
ππβπΉπ(,π§)πΉπ(,π§)πΞ© (A.2)
and it is a weighted sum (in the continuum) of each elementalcross-section area where the weight functions account for the
12 Mathematical Problems in Engineering
0
2
4
6
8
10
12
14
16
PLANE183FEM 2D TLN=2
N=3N=5
3 2 5 6 1 4 7 8 0
102Γ οΌ―z
Figure 14: Load factor π versus dimensionless displacement οΏ½οΏ½π§ for a short doubly clamped beam.
PLANE183FEM 2D TLN=2
N=3N=5
β0.5
β0.25
0
0.25
0.5
οΌ΄
9 9.5 10 10.5 11 11.5 12 12.5 13 8.5
103Γ xx
Figure 15: Dimensionless Cauchy stress οΏ½οΏ½π₯π₯ along the thickness coordinate at π₯ = π/4 and π = 11.04 for a short doubly clamped beam.
spatial distribution of the geometry and thematerial. πΌπ(,π₯) π(,π₯) isthe integral along the element axis of the product of the shapefunctions and/or their derivatives:
πΌπ(,π₯)π(,π₯) = β«ππ
ππ(,π₯)ππ(,π₯)ππ₯ (A.3)
These integrals are numerically evaluated through Gaussianquadrature. Kππ‘1ππππ is the initial-displacement, or initial-slope,
contribution to the tangent stiffness matrix. Its componentsare
πΎππ‘1π₯π₯ππππ = ππ’π‘π (2π½11πππ‘πΌπ,π₯π,π₯π,π₯ + π½13ππ,π§π‘,π§πΌπ,π₯ππ + π½13π,π§ππ‘,π§πΌππ,π₯π+ 2π½15πππ‘,π§πΌπ,π₯π,π₯π + 2π½15ππ,π§π‘πΌπ,π₯ππ,π₯ + 2π½15π,π§ππ‘πΌππ,π₯π,π₯+ 2π½35π,π§π,π§π‘,π§πΌπππ + 2π½55π,π§π,π§π‘πΌπππ,π₯ ) + π½55π,π§ππ‘,π§πΌππ,π₯π
Mathematical Problems in Engineering 13
PLANE183FEM 2D TLN=2
N=3N=5
0.5 2 3 1 2.5 1.5 3.5 0
103Γ zz
β0.5
β0.25
0
0.25
0.5
οΌ΄
Figure 16: Dimensionless Cauchy stress οΏ½οΏ½π§π§ along the thickness coordinate at π₯ = π/4 and π = 11.04 for a short doubly clamped beam.
PLANE183FEM 2D TLN=2
N=3N=5
β0.5
β0.25
0
0.25
0.5
οΌ΄
0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 0.1
102Γ xz
Figure 17: Dimensionless Cauchy stress οΏ½οΏ½π₯π§ along the thickness coordinate at π₯ = π/4 and π = 11.04 for a short doubly clamped beam.
+ π½55ππ,π§π‘,π§πΌπ,π₯ππ) + ππ’π‘πππ’π π (π½11πππ‘π πΌπ,π₯π,π₯π,π₯π,π₯+ π½13ππ,π§π‘π ,π§πΌπ,π₯ππ,π₯π + π½13π,π§ππ‘,π§π πΌππ,π₯ππ,π₯ + π½15πππ‘π ,π§πΌπ,π₯π,π₯π,π₯π+ π½15ππ,π§π‘π πΌπ,π₯ππ,π₯π,π₯ + π½15πππ‘,π§π πΌπ,π₯π,π₯ππ,π₯ + π½15π,π§ππ‘π πΌππ,π₯π,π₯π,π₯+ π½35ππ,π§π‘,π§π ,π§πΌπ,π₯πππ + π½35π,π§ππ‘,π§π ,π§πΌππ,π₯ππ + π½35π,π§π,π§π‘,π§π πΌππππ,π₯+ π½35π,π§π,π§π‘π ,π§πΌπππ,π₯π + π½33π,π§π,π§π‘,π§π ,π§πΌππππ + π½55πππ‘,π§π ,π§πΌπ,π₯π,π₯ππ
+ π½55ππ,π§π‘,π§π πΌπ,π₯πππ,π₯ + π½55π,π§ππ‘π ,π§πΌππ,π₯π,π₯π + π½55π,π§π,π§π‘π πΌπππ,π₯π,π₯)(A.4)
πΎππ‘1π₯π§ππππ = ππ’π‘π (π½13ππ,π§π‘πΌπ,π₯ππ,π₯ + π½15πππ‘πΌπ,π₯π,π₯π,π₯ + π½35ππ,π§π‘,π§πΌπ,π₯ππ+ π½55πππ‘,π§πΌπ,π₯π,π₯π + π½33π,π§π,π§π‘,π§πΌπππ + π½35π,π§ππ‘,π§πΌππ,π₯π + π½35π,π§π,π§π‘πΌπππ,π₯+ π½55π,π§ππ‘πΌππ,π₯π,π₯) + ππ€π‘π (π½11πππ‘πΌπ,π₯π,π₯π,π₯ + π½13ππ,π§π‘,π§πΌπ,π₯ππ
14 Mathematical Problems in Engineering
0
1
2
3
4
5
6
7
8
PLANE183FEM 2D TLN=2
N=3N=5
1.5 1 2 0 2.5 3 3.5 4 0.5
102Γ οΌ―x
Figure 18: Load factor π versus dimensionless displacement οΏ½οΏ½π₯ for a short simply supported beam.
0
1
2
3
4
5
6
7
8
PLANE183FEM 2D TLN=2
N=3N=5
0.2 0.4 0.6 0.8 1 1.2 1.4 010 Γ οΌ―οΌ΄
Figure 19: Load factor π versus dimensionless displacement οΏ½οΏ½π§ for a short simply supported beam.
+ π½15πππ‘,π§πΌπ,π₯π,π₯π + π½15ππ,π§π‘πΌπ,π₯ππ,π₯ + π½15π,π§ππ‘πΌππ,π₯π,π₯ + π½35π,π§π,π§π‘,π§πΌπππ+ π½55π,π§ππ‘,π§πΌππ,π₯π + π½55π,π§π,π§π‘πΌπππ,π₯) + ππ’π‘πππ€π π (π½11πππ‘π πΌπ,π₯π,π₯π,π₯π,π₯+ π½13ππ,π§π‘π ,π§πΌπ,π₯ππ,π₯π + π½15πππ‘π ,π§πΌπ,π₯π,π₯π,π₯π + π½15ππ,π§π‘π πΌπ,π₯ππ,π₯π,π₯+ π½15πππ‘,π§π πΌπ,π₯π,π₯ππ,π₯ + π½35ππ,π§π‘,π§π ,π§πΌπ,π₯πππ + π½55πππ‘,π§π ,π§πΌπ,π₯π,π₯ππ
+ π½55ππ,π§π‘,π§π πΌπ,π₯πππ,π₯ + π½13π,π§ππ‘,π§π πΌππ,π₯ππ,π₯ + π½33π,π§π,π§π‘,π§π ,π§πΌππππ+ π½35π,π§ππ‘,π§π ,π§πΌππ,π₯ππ + π½35π,π§π,π§π‘,π§π πΌππππ,π₯ + π½15π,π§ππ‘π πΌππ,π₯π,π₯π,π₯+ π½35π,π§π,π§π‘π ,π§πΌπππ,π₯π + π½55π,π§ππ‘π ,π§πΌππ,π₯π,π₯π + π½55π,π§π,π§π‘π πΌπππ,π₯π,π₯)
(A.5)
Mathematical Problems in Engineering 15
β0.5
β0.25
0
0.25
0.5
PLANE183FEM 2D TLN=2
N=3N=5
οΌ΄
β8 β6 β4 β2 0 2 4 6 8 10β10
102Γ xx
Figure 20: Dimensionless Cauchy stress οΏ½οΏ½π₯π₯ along the thickness coordinate at π₯ = π/4 and π = 6.03 for a short simply supported beam.
PLANE183FEM 2D TLN=2
N=3N=5
β0.5
β0.25
0
0.25
0.5
οΌ΄
2β2 6β6 4β4 0 8
103Γ zz
Figure 21: Dimensionless Cauchy stress οΏ½οΏ½π§π§ along the thickness coordinate at π₯ = π/4 and π = 6.03 for a short simply supported beam.
πΎππ‘1π§π₯ππππ = ππ’π‘π (π½15πππ‘πΌπ,π₯π,π₯π,π₯ + π½35ππ,π§π‘,π§πΌπ,π₯ππ + π½55πππ‘,π§πΌπ,π₯π,π₯π+ π½55ππ,π§π‘πΌπ,π₯ππ,π₯ + π½13π,π§ππ‘πΌππ,π₯π,π₯ + π½33π,π§π,π§π‘,π§πΌπππ + π½35π,π§ππ‘,π§πΌππ,π₯π+ π½35π,π§π,π§π‘πΌπππ,π₯) + ππ€π‘π (π½11πππ‘πΌπ,π₯π,π₯π,π₯ + π½15ππ,π§π‘πΌπ,π₯ππ,π₯+ π½15πππ‘,π§πΌπ,π₯π,π₯π + π½55ππ,π§π‘,π§πΌπ,π₯ππ + π½13π,π§ππ‘,π§πΌππ,π₯π + π½35π,π§π,π§π‘,π§πΌπππ
+ π½15π,π§ππ‘πΌππ,π₯π,π₯ + π½55π,π§π,π§π‘πΌπππ,π₯) + ππ€π‘πππ’π π (π½11πππ‘π πΌπ,π₯π,π₯π,π₯π,π₯+ π½13ππ,π§π‘π ,π§πΌπ,π₯ππ,π₯π + π½15πππ‘π ,π§πΌπ,π₯π,π₯π,π₯π + π½15ππ,π§π‘π πΌπ,π₯ππ,π₯π,π₯+ π½15πππ‘,π§π πΌπ,π₯π,π₯ππ,π₯ + π½35ππ,π§π‘,π§π ,π§πΌπ,π₯πππ + π½55πππ‘,π§π ,π§πΌπ,π₯π,π₯ππ+ π½55ππ,π§π‘,π§π πΌπ,π₯πππ,π₯ + π½13π,π§ππ‘,π§π πΌππ,π₯ππ,π₯ + π½33π,π§π,π§π‘,π§π ,π§πΌππππ
16 Mathematical Problems in Engineering
β0.5
β0.25
0
0.25
0.5
PLANE183FEM 2D TLN=2
N=3N=5
οΌ΄
β2.5 β1.5 β1 β0.5 0 0.5 1 1.5 2 2.5 3β2
102Γ xz
Figure 22: Dimensionless Cauchy stress οΏ½οΏ½π₯π§ along the thickness coordinate at π₯ = π/4 and π = 6.03 for a short simply supported beam.
Table 1: Displacements for a short cantilever beam, π = 10.10 Γ βοΏ½οΏ½π₯ 10 Γ οΏ½οΏ½π§
Plane183 6.1163 8.7165Beam3 TBT 6.4377 8.2409Beam3 EBT 6.3579 8.1661EBT [1] 6.2652 8.1061N B2 B3 B4 B2 B3 B45 6.1579 6.1585 6.1585 8.6817 8.6820 8.68204 6.1575 6.1580 6.1581 8.6809 8.6812 8.68123 6.1596 6.1601 6.1601 8.6760 8.6764 8.67642 6.1521 6.1527 6.1527 8.6671 8.6674 8.6674
= 12.06 = 6.03 = 3.54
FEM 2D TLN=5
β0.5
β0.25
0
0.25
0.5
οΌ΄
β6 0 12β12 6
102Γ xx
Figure 23: Through-the-thickness profile of οΏ½οΏ½π₯π₯ for different loadfactors.
+ π½35π,π§ππ‘,π§π ,π§πΌππ,π₯ππ + π½35π,π§π,π§π‘,π§π πΌππππ,π₯ + π½15π,π§ππ‘π πΌππ,π₯π,π₯π,π₯+ π½35π,π§π,π§π‘π ,π§πΌπππ,π₯π + π½55π,π§ππ‘π ,π§πΌππ,π₯π,π₯π + π½55π,π§π,π§π‘π πΌπππ,π₯π,π₯)
(A.6)
πΎππ‘1π§π§ππππ = ππ€π‘π (π½13ππ,π§π‘πΌπ,π₯ππ,π₯ + π½13π,π§ππ‘πΌππ,π₯π,π₯ + 2π½15πππ‘πΌπ,π₯π,π₯π,π₯+ 2π½35ππ,π§π‘,π§πΌπ,π₯ππ + 2π½35π,π§ππ‘,π§πΌπ,π₯ππ + 2π½35π,π§π,π§π‘πΌπππ,π₯+ 2π½55πππ‘,π§πΌπ,π₯π,π₯π + π½55ππ,π§π‘πΌπ,π₯ππ,π₯ ) + π½55π,π§ππ‘πΌππ,π₯π,π₯+ 2π½33π,π§π,π§π‘,π§πΌπππ) + ππ€π‘πππ€π π (π½11πππ‘π πΌπ,π₯π,π₯π,π₯π,π₯+ π½33π,π§π,π§π‘,π§π ,π§πΌππππ + π½55πππ‘,π§π ,π§πΌπ,π₯π,π₯ππ + π½55ππ,π§π‘,π§π πΌπ,π₯πππ,π₯+ π½55π,π§ππ‘π ,π§πΌππ,π₯π,π₯π + π½55π,π§π,π§π‘π πΌπππ,π₯π,π₯ + π½13ππ,π§π‘π ,π§πΌπ,π₯ππ,π₯π+ π½13π,π§ππ‘,π§π πΌππ,π₯ππ,π₯ + J15πππ‘π ,π§ πΌπ,π₯π,π₯π,π₯π + π½15ππ,π§π‘π πΌπ,π₯ππ,π₯π,π₯
Mathematical Problems in Engineering 17
β0.5
β0.25
0
0.25
0.5
β10 β8 β6 β4 β2 0 2 4 6 8 10 12
= 12.06 = 6.03 = 3.54
FEM 2D TLN=5
οΌ΄
103Γ zz
Figure 24: Through-the-thickness profile of οΏ½οΏ½π§π§ for different load factors.
β2.5 β2 β1.5 β1 β0.5 0 0.5 1 1.5 2 2.5
= 12.06
= 6.03 = 3.54
FEM 2D TLN=5
102Γ xz
β0.5
β0.25
0
0.25
0.5
οΌ΄
Figure 25: Through-the-thickness profile of οΏ½οΏ½π₯π§ for different load factors.
Table 2: Cauchy stresses evaluated at π₯ = π/4 and π§ = ββ/2 for a short cantilever beam, π = 5.20.10 Γ οΏ½οΏ½π₯π₯ 10 Γ οΏ½οΏ½π§π§ 10 Γ οΏ½οΏ½π₯π§
Plane183 2.4166 1.3372 1.7978FEM 2D TL 2.7403 1.5302 2.0480N B2 B3 B4 B2 B3 B4 B2 B3 B45 2.5622 2.7426 2.7414 1.6556 1.5336 1.5291 2.0730 2.0496 2.04744 2.5632 2.7435 2.7424 1.6552 1.5332 1.5287 2.0730 2.0494 2.04723 2.5554 2.7355 2.7344 1.6468 1.5248 1.5203 2.0779 2.0546 2.05222 2.4216 2.5995 2.5984 1.7133 1.5896 1.5851 2.1354 2.1114 2.1090
18 Mathematical Problems in Engineering
0-0.05-0.1-0.15-0.2-0.25-0.3-0.35(a) FEM 2D TL
0-0.05-0.1-0.15-0.2-0.25-0.3-0.35(b) N=5
Figure 26: Axial displacement οΏ½οΏ½π₯, cantilever beam, π = 3.79, and π/π = 10.
0.10 0.2 0.3 0.50.4 0.6(a) FEM 2D TL
0.10 0.2 0.3 0.50.4 0.6(b) N=5
Figure 27: Transverse displacement οΏ½οΏ½π§, cantilever beam, π = 3.79, and π/π = 10.
Table 3: Displacements for a short doubly-clamped beam, π = 15.89.103 Γ βοΏ½οΏ½π₯ 102 Γ οΏ½οΏ½π§
Plane183 9.4375 7.1822FEM 2D TL 8.8841 6.7892Beam3 TBT 9.2102 7.1003Beam3 EBT 9.5180 6.4598N B2 B3 B4 B2 B3 B45 8.8863 8.8848 8.8848 6.7911 6.7918 6.79184 8.8857 8.8841 8.8841 6.7879 6.7887 6.78873 8.8829 8.8814 8.8814 6.7725 6.7733 6.77332 8.8476 8.8461 8.8460 6.7109 6.7117 6.7117
Mathematical Problems in Engineering 19
-0.8 -0.6 -0.4 -0.2 0.80.60.40.20(a) FEM 2D TL
-0.8 -0.6 -0.4 -0.2 0.80.60.40.20(b) N=5
Figure 28: Piola-Kirchoff stress ππ₯π₯, cantilever beam, π = 3.79, and π/π = 10.
-0.2 0.2 0.3-0.1 0.10(a) FEM 2D TL
-0.2 0.2 0.3-0.1 0.10(b) N=5
Figure 29: Piola-Kirchoff stress ππ§π§, cantilever beam, π = 3.79, and π/π = 10.
+ π½15πππ‘,π§π πΌπ,π₯π,π₯ππ,π₯ + π½15π,π§ππ‘π πΌππ,π₯π,π₯π,π₯ + π½35ππ,π§π‘,π§π ,π§πΌπ,π₯πππ+ π½35π,π§ππ‘,π§π ,π§πΌππ,π₯ππ + π½35π,π§π,π§π‘,π§π πΌππππ,π₯ + π½35π,π§π,π§π‘π ,π§πΌπππ,π₯π)
(A.7)
Kππ‘2ππππ is the initial stress, or geometric, contribution to thetangent stiffness matrix. Its components are
πΎππ‘2π₯π₯ππππ = πΎππ‘2π§π§ππππ = ππ’π‘π (π½11πππ‘πΌπ,π₯π,π₯π,π₯ + π½15πππ‘,π§πΌπ,π₯π,π₯π+ π½15ππ,π§π‘πΌπ,π₯ππ,π₯ + π½55ππ,π§π‘,π§πΌπ,π₯ππ + π½15π,π§ππ‘πΌππ,π₯π,π₯ + π½55π,π§ππ‘,π§πΌππ,π₯π+ π½13π,π§π,π§π‘πΌπππ,π₯ + π½35π,π§π,π§π‘,π§πΌπππ) ππ€π‘π (π½13πππ‘,π§πΌπ,π₯π,π₯π+ π½15πππ‘πΌπ,π₯π,π₯π,π₯ + π½35ππ,π§π‘,π§πΌπ,π₯ππ + π½55ππ,π§π‘πΌπ,π₯ππ,π₯ + π½35π,π§ππ‘,π§πΌππ,π₯π
+ π½55π,π§ππ‘πΌππ,π₯π,π₯ + π½33π,π§π,π§π‘,π§πΌπππ + π½35π,π§π,π§π‘πΌπππ,π₯) 12 (ππ’π‘πππ’π π+ ππ€π‘πππ€π π) (π½11πππ‘π πΌπ,π₯π,π₯π,π₯π,π₯ + π½13πππ‘,π§π ,π§πΌπ,π₯π,π₯ππ+ π½15πππ‘,π§π πΌπ,π₯π,π₯ππ,π₯ + π½15πππ‘π ,π§πΌπ,π₯π,π₯π,π₯π + π½15ππ,π§π‘π πΌπ,π₯ππ,π₯π,π₯+ π½35ππ,π§π‘,π§π ,π§πΌπ,π₯πππ + π½55ππ,π§π‘,π§π πΌπ,π₯πππ,π₯ + π½55ππ,π§π‘π ,π§πΌπ,π₯ππ,π₯π+ π½15π,π§ππ‘π πΌππ,π₯π,π₯π,π₯ + π½35π,π§ππ‘,π§π ,π§πΌππ,π₯ππ + π½55π,π§ππ‘,π§π πΌππ,π₯ππ,π₯+ π½55π,π§ππ‘π ,π§πΌππ,π₯π,π₯π + π½13π,π§π,π§π‘π πΌπππ,π₯π,π₯ + π½33π,π§π,π§π‘,π§π ,π§πΌππππ+ π½35π,π§π,π§π‘,π§π πΌππππ,π₯ + π½35π,π§π,π§π‘π ,π§πΌπππ,π₯π)
(A.8)
20 Mathematical Problems in Engineering
0 0.05 0.1 0.15 0.2(a) FEM 2D TL
0 0.05 0.1 0.15 0.2(b) N=5
Figure 30: Piola-Kirchoff stress ππ₯π§, cantilever beam, π = 3.79, and π/π = 10.
Table 4: Cauchy stresses evaluated at π₯ = π/4 and π§ = 0 for a short doubly-clamped beam, π = 11.04.103 Γ οΏ½οΏ½π₯π₯ 103 Γ οΏ½οΏ½π§π§ 102 Γ οΏ½οΏ½π₯π§
Plane183 8.6194 3.2658 1.1290FEM 2D TL 9.8958 3.1957 1.1281N B2 B3 B4 B2 B3 B4 B2 B3 B45 9.9050 9.8829 9.8987 3.1738 3.2083 3.1858 1.1210 1.1315 1.12504 9.9077 9.8858 9.9014 3.1859 3.2205 3.1979 1.1208 1.1312 1.12483 9.9050 9.8830 9.8987 3.1858 3.2200 3.1976 1.1210 1.1314 1.12492 10.4620 10.4390 10.4550 2.3066 2.3405 2.3182 0.8477 0.8581 0.8516
Table 5: Displacements for a short simply supported beam, π = 8.36.102 Γ οΏ½οΏ½π₯ 10 Γ οΏ½οΏ½π§
Plane183 3.9585 1.3032FEM 2D TL 3.8909 1.2805N B2 B3 B4 B2 B3 B45 3.9239 3.9239 3.9192 1.2843 1.2843 1.28394 3.9011 3.9000 3.8960 1.2820 1.2819 1.28143 3.8267 3.8251 3.8232 1.2739 1.2738 1.27352 3.7501 3.7482 3.7474 1.2583 1.2580 1.2579
Table 6: Cauchy stresses evaluated at π₯ = π/4 and π§ = ββ/2 for a short simply supported beam, π = 6.03.102 Γ βοΏ½οΏ½π₯π₯ 103 Γ βοΏ½οΏ½π§π§ 102 Γ βοΏ½οΏ½π₯π§
Plane183 8.6470 5.2555 2.1314FEM 2D TL 8.2064 4.8990 2.0056N B2 B3 B4 B2 B3 B4 B2 B3 B45 8.2878 8.2030 8.2002 6.9946 4.8699 4.9120 2.2776 2.0010 2.00674 8.2696 8.1840 8.1811 6.9402 4.8190 4.8614 2.2707 1.9942 2.00023 8.2018 8.1142 8.1125 6.7966 4.6917 4.7352 2.2464 1.9717 1.97792 8.4861 8.3928 8.3917 4.0106 1.8893 1.9344 1.5683 1.2941 1.3006
Mathematical Problems in Engineering 21
πΎππ‘2π₯π§ππππ = πΎππ‘2π§π₯ππππ = 0 (A.9)
The integrals π½πβπ(,π§)π(,π§) π‘(,π§) , πΌπ(,π₯)π(,π₯)π(,π₯) , π½πβπ(,π§)π(,π§)π‘(,π§)π (,π§) , andπΌπ(,π₯)π(,π₯)π(,π₯)π(,π₯) in (A.6) and (A.9) are given by
π½πβπ(,π§)π(,π§)π‘(,π§) = β«Ξ©π=βπΓππ
ππβπΉπ(,π§)πΉπ(,π§)πΉπ‘(,π§)πΞ© (A.10)
πΌπ(,π₯)π(,π₯)π(,π₯) = β«ππ
ππ(,π₯)ππ(,π₯)ππ(,π₯)ππ₯ (A.11)
π½πβπ(,π§)π(,π§)π‘(,π§)π (,π§) = β«Ξ©π=βπΓππ
ππβπΉπ(,π§)πΉπ(,π§)πΉπ‘(,π§)πΉπ (,π§)πΞ© (A.12)
πΌπ(,π₯)π(,π₯)π(,π₯)π(,π₯) = β«ππ
ππ(,π₯)ππ(,π₯)ππ(,π₯)ππ(,π₯)ππ₯ (A.13)
If a MITC beam element is considered, the πΌ-integrals in(A.3), (A.11), and (A.13) are replaced, respectively, by thefollowing integrals:
πΌπ(,π₯)π(,π₯) = β«ππ
πππππ(,π₯)πππππ(,π₯)ππ₯ (A.14)
πΌπ(,π₯)π(,π₯)π(,π₯) = β«ππ
πππππ(,π₯)πππππ(,π₯)πππ(,π₯)ππ₯ (A.15)
πΌπ(,π₯)π(,π₯)π(,π₯)π(,π₯) = β«ππ
πππππ(,π₯)πππ(,π₯)πππππ(,π₯)πππ(,π₯)ππ₯ (A.16)
Data Availability
The data used to support the findings of this study areincluded within the article.
Conflicts of Interest
The authors declare that there are no conflicts of interestregarding the publication of this paper.
Acknowledgments
This work has been partially supported by the EuropeanUnion within the Horizon 2020 Research and InnovationProgram under Grant Agreement no. 642121.
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