A Numerical and Analytical Solution for the Free Vibration of Laminated Composites Using Different Plate Theories

Document Type: Research Paper

Authors

1 University of Applied Science and Technology

2 University of Semnan

Abstract

An analytical and numerical solution for the free vibration of laminated polymeric composite plates with different layups is studied in this paper. The governing equations of the laminated composite plates are derived from the classical laminated plate theory (CLPT) and the first-order shear deformation plate theory (FSDT). General layups are evaluated by the assumption of cross-ply and angle-ply laminated plates. The solver is coded in MATLAB. As a verification method, a finite element code using ANSYS is also developed. The effects of lamination angle, plate aspect ratio and modulus ratio on the fundamental natural frequencies of a laminated composite are also investigated and good agreement is found between the results evaluated and those available in the open literature. The results show that the fundamental frequency increases with the modular ratio and the bending-stretching coupling lowers the vibration frequencies for both cross-ply and angle-ply laminates with the CLPT. Also it is found that the effect of bending-stretching coupling, transverse shear deformation and rotary inertia is to lower the fundamental frequencies.

Keywords



 

Mechanics of Advanced Composite Structures 4 (2017) 75-87

 

 

 

 

 

Semnan University

Mechanics of Advanced Composite Structures

journal homepage: http://MACS.journals.semnan.ac.ir

 

A Numerical and Analytical Solution for the Free Vibration of Laminated Composites Using Different Plate Theories

M.A. Torabizadeh a*, A. Fereidoon b

a Department of Mechanical Engineering, University of Applied Science and Technology, Mashhad, Iran

b Department of Mechanical Engineering, University of Semnan, Semnan, Iran

 

Paper INFO

 

ABSTRACT

Paper history:

Received 2016-11-27

Revised 2017-01-21

Accepted 2017-02-09

An analytical and numerical solution for the free vibration of laminated polymeric composite plates with different layups is studied in this paper. The governing equations of the laminated composite plates are derived from the classical laminated plate theory (CLPT) and the first-order shear deformation plate theory (FSDT). General layups are evaluated by the assumption of cross-ply and angle-ply laminated plates. The solver is coded in MATLAB. As a verification method, a finite element code using ANSYS is also developed. The effects of lamination angle, plate aspect ratio and modulus ratio on the fundamental natural frequencies of a laminated composite are also investigated and good agreement is found between the results evaluated and those available in the open literature. The results show that the fundamental frequency increases with the modular ratio and the bending-stretching coupling lowers the vibration frequencies for both cross-ply and angle-ply laminates with the CLPT. Also it is found that the effect of bending-stretching coupling, transverse shear deformation and rotary inertia is to lower the fundamental frequencies.

 

 

Keywords:

Free vibration

Laminated composites

Plate theories

Numerical method

Analytical method

DOI: 10.22075/MACS.2017.1768.1090

© 2017 Published by Semnan University Press. All rights reserved.

 

 

  1. Introduction

 

A composite material can be defined as a combination of two or more materials that results in better properties than those of the individual components used alone. In contrast to metallic alloys, each material retains its separate chemical, physical and mechanical properties. The two constituents are reinforcement and a matrix. When composites are compared to bulk materials, the main advantages of composite materials are their high strength and stiffness, combined with low density, allowing for a weight reduction in the finished part. The reinforcing phase provides the strength and stiffness. In most cases, the reinforcement is harder, stronger and stiffer than the matrix. The reinforcement is usually a fiber or a particulate. Particulate composites have dimensions that are approximately equal in all directions. They may be spherical, platelets, or any other regular or irregular geometry. Particulate composites tend to be much weaker and less stiff than continuous-fiber composites, but they are usually much less expensive. Particulate reinforced composites usually contain less reinforcement (up to 40 volume percent to 50 volume percent) due to processing difficulties and brittleness [1].

A fiber’s length is much greater than its diameter. The length-to-diameter (l/d) ratio is known as the aspect ratio and can vary greatly. Continuous fibers have long aspect ratios, whereas discontinuous fibers have short ones. Continuous-fiber composites normally have a preferred orientation, whereas discontinuous fibers generally have a random orientation. Examples of continuous reinforcements include unidirectional, woven cloth and helical winding, whereas examples of discontinuous reinforcements are chopped fibers and random material. Continuous-fiber composites are often made into laminates by stacking single sheets of continuous fibers in different orientations to obtain the desired strength and stiffness properties with fiber volumes as high as 60 percent to 70 percent. Fibers produce high-strength composites because of their small diameter; they contain far fewer defects (normally surface defects) compared to those in the material produced in bulk. As a general rule, the smaller the diameter of the fiber, the higher its strength, but often the cost increases as the diameter becomes smaller. In addition, smaller-diameter/high-strength fibers have greater flexibility and are more amenable to fabrication processes, such as weaving or forming over radius. Typical fibers include glass, aramid and carbon, which may be continuous or discontinuous. The continuous phase is the matrix, which is a polymer, metal or ceramic. Polymers have low strength and stiffness, metals have intermediate strength and stiffness but high ductility, and ceramics have high strength and stiffness but are brittle. The matrix (continuous phase) performs several critical functions, including maintaining the fibers in the proper orientation and spacing and protecting them from abrasion and the environment. In polymer and metal matrix composites that form a strong bond between the fiber and the matrix, the matrix transmits loads from the matrix to the fibers through shear loading at the interface. In ceramics-matrix composites, the objective is often to increase the toughness rather than the strength and stiffness; therefore, a low interfacial strength bond is desirable [1].

Tan and Nie [2] studied free and forced vibration of variable stiffness composite annular thin plates with elastically restrained edges based on the classical plate theory. They found that the transverse mode shapes of the plates with in-plane variable stiffness are different from those with constant stiffness. Zhang et al. [3] analyzed free vibration analysis of triangular CNT-reinforced composite plates subjected to in-plane stresses using the FSDT element-free method. Chakraborty et al. [4] presented a novel approach, referred to as polynomial correlated function expansion (PCFE), for a stochastic free-vibration analysis of a composite laminate. Finally, based on the numerical results, new physical insights had been created on the dynamic behavior of composite laminates. Ganesh et al. [5] studied the free vibration analysis of delaminated composite plates using a finite element method. Mantari and Ore [6] presented a simplified first-order shear deformation theory (FSDT) for a laminated composite and sandwich plates. Their approach had a novel displacement field that includes undetermined integral terms and contains only four unknowns. Su et al. [7] illustrated a modified Fourier series to study the free vibration of a laminated composite and four-parameter functionally graded sector plates with general boundary conditions. Zhang et al. [8] studied the free-vibration analysis of functionally graded carbon nanotube-reinforced composite triangular plates using the FSDT and the element-free IMLS-Ritz method. They also examined the influence of a carbon nanotube volume fraction, plate thickness-to-width ratio, plate-aspect ratio and a boundary condition on the plate’s vibration behavior. Marjanović and Vuksanović [9] illustrated a layerwise solution to free vibrations and the buckling of a laminated composite and sandwich plates with embedded delamination. The effects of plate geometry, lamination scheme, degree of orthotropy and delamination size or position on the dynamic characteristics of the plate were presented. Boscolo [10] presented an analytical closed-form solution for a free-vibration analysis of multilayered plates by using a layer-wise displacement assumption based on Carrera’s Unified Formulation. A wide range of boundary conditions were analyzed by using a Levy-type solution. Ou et al. [11] presented an efficient method for predicting the free and transient vibrations of multilayered composite structures with parallelepiped shapes, including beams, plates and solids. Rafiee et al. [12] analyzed the geometrically nonlinear free vibration of shear deformable piezoelectric carbon nanotube/fiber/polymer multiscale laminated composite plates. Akhras and Li [13] used a spline finite strip with higher-order shear deformation for stability and a free-vibration analysis of piezoelectric composite plates. Grover et al. [14] assessed a new shear deformation theory for free-vibration-response laminated composite and sandwich plates. They compared the results with finite element and analytical solutions. Jafari et al. [15] presented a free-vibration analysis of a generally laminated composite beam (LCB) based on the Timoshenko beam theory using the method of Lagrange multipliers. They examined some parameters, such as the slenderness ratio, the rotary inertia, the shear deformation, material anisotropy, ply configuration and boundary conditions on the natural frequency and mode shape. Tai and Kim [16] illustrated the free vibration of laminated composite plates using two variable refined plate theories. They applied the Navier technique to obtain the closed-form solutions of anti-symmetric cross-ply and angle-ply laminates. Srinivasa et al. [17] and Ramu and Mohanty [18] used finite element results as a verification method with those obtained from experimental tests on the free vibration of composite plates. Chandrashekhara [19] presented an exact solution for the free vibration of symmetrically laminated composite beams. Ke et al. [20] investigated the nonlinear free vibration of functionally graded nanocomposite beams reinforced by single-walled carbon nanotubes (SWCNTs) based on the Timoshenko beam theory and von Kármán geometric nonlinearity. Also, the free vibration of anisotropic thin-walled composite beams and delaminated composite beams were performed by Song [21] and Lee [22], respectively.

Based on papers reviewed in the literature, few investigations were found that compared the analytical and numerical analyses of different theories and lamination layups. Therefore, in this paper, the analytical and numerical solutions for the free vibration of laminated polymeric composite plates with different layups are compared. Two different theories and layups are selected. Also, finite-element analysis is performed using ANSYS to validate results obtained by analytical methods. The solver is coded in MATLAB. Also investigated are the effects of different parameters, such as the lamination angle, the plate aspect ratio and the modulus ratio on the fundamental natural frequencies of laminated composite. The main objective of this paper is to compare different theories and lamination schemes on the vibration response of laminated composites.

 

  1. Theoretical Formulation

2.1. Classical lamination plate theory (CLPT)

2.1.1 Displacement and strains

A rectangular plate of sides a and b with thickness h is shown in Fig. 1. Based on the classical lamination plate theory, the following displacement field can be assumed [23]:

 

(1)

 

(2)

 

(3)

where  are the displacements along the coordinate lines of a material point on xy-plane.

The von Karman strains associated with the displacement field in static loading can be computed using the strain-displacement relations for small strains:

 

(4)

Note that the transverse strains are identically zero in classical plate theory. The first three strains have the form

 

(5)

 

2.1.2 Equilibrium equations

By using Eqs. (4) and (5), the constitutive equations are obtained. Equations of equilibrium can be derived using the variational principle, which is not explained in detail here (see [23]). The Euler-Lagrange equations of the theory are obtained as follows,

 

(6)

 

(7)

 

(8)

 

(9)

where, the quantities Nij are called the in-plane force resultants and Mij are called the moment resultants and (I0, I1, I2) are the mass moments of inertia.

 

2.1.3 Navier solution methodology

The displacement fields are assumed by the following form:

 

(10)

 

(11)

 

(12)

where Umn, Vmn and Wmn are the coefficients that should be determined and  and

The consideration of Eqs. (10) – (12), shows that the mechanical transverse load q should also be expanded in a double sine series. Thus,

 

(13)

 

(14)

 

 

Figure 1. The geometry of simply supported rectangular laminated plates used in the analytical solutions.

 

 

 

 

Substituting expansions (10–12) into expressions given in Eqs. (6-8) without thermal loads yields

 

(15)

where Aij, Dij and Bij are called extensional, bending and bending-extensional coupling stiffness, respectively [23]. Also,  and . Note that the edge shear force is necessarily zero.

Substituting the expansion (13) into (15), we obtain expressions of the form

 

 

(16)

where amn, bmn and cmn are coefficients whose explicit form will be given shortly. Since Eq. (16) must hold for any m, n, x and y, it follows that amn=0, bmn=0 and cmn=0 for every m and n. The explicit forms of these coefficients are given by:

 

(17)

or in matrix form

 

(18)

where is

 

(19)

Eqs. (18) provide three second-order differential equations among the three variables Umn, Vmn and Wmn for any fixed values of m and n.

For free vibration, all applied loads and the in-plane forces are set to zero, and we assume a periodic solution of the form:

 

(20)

where  and is the frequency of natural vibration. Then Eq. (18) reduces to the eigenvalue problem:

(

(21)

For a nontrivial solution, the determinant of the coefficient matrix in (21) should be zero, which yields the characteristic polynomial

 ,

(22)

where  is the eigenvalue and

 

(23)

The real positive roots of this cubic equation give the square of the natural frequency  associated with mode (m,n). The smallest of the frequencies is called the fundamental frequency. In general,  is not the fundamental frequency; the smallest frequency might occur for values other than m = n = 1.

If the in-plane inertias are neglected (i.e., ), and irrespective of whether the rotary inertia is zero, Eq. (22) will be

 

(24)

Note that if the in-plane inertias are not neglected, the eigenvalue problem cannot be simplified to a single equation, even if the rotary inertia is zero.

 

2.2. First-order shear deformation theory (FSDT)

2.2.1 Displacement and strains

Under the same assumptions and restrictions as in the classical laminate theory, the displacement field of the first-order theory is of the form:

 

 

 

(25)

Where

 

(26)

which indicate that and are the rotations of a transverse normal about the y- and x- axes, respectively.

The nonlinear strains associated with the displacement field (25) are obtained as

 

 

 

 

 

 

 

 

 

 

 

 

 

(27)

Note that the strains are linear through the laminate thickness, whereas the transverse shear strains are constant through the thickness of the laminate in the first-order laminated theory. These strains have the fo

 

(28)

 

2.2.2 Equilibrium equations

The governing equations of the first-order theory will be derived using the dynamic version of the principle of virtual displacements. The Euler-Lagrange equations are obtained as follows

 

 

 

 

,

(29)

where Qx and Qy are called transverse force resultants and

 

(30)

Parameter K is called a shear correction coefficient and is used because of a discrepancy between the actual stress state and the constant stress state predicted by the first-order theory.

2.2.2 Boundary condition

The natural boundary conditions are obtained by setting the coefficients of and  to zero separately:

 

(31)

where

 

(32)

Thus, the primary and secondary variables of the theory are

primary variables:  

secondary variables :  

(33)

The initial conditions of the theory involve specifying the values of the displacements and their first derivatives with respect to time at t = 0.

 

2.2.3 Equations of motion

The boundary conditions are satisfied by the following expansions

 

(34)

The Navier solution can be calculated from

 

(35)

where

 

 

 

 

 

 

 

 

 

(36)

For free vibration, all thermal and mechanical loads are set to zero and substitute to Eq. (29) and obtain

 

(37)

where

,

 

(38)

and . When rotary inertia is omitted, Eq. (37) can be simplified by eliminating Xmn and Ymn (i.e., using the static condensation method) as follows

 

(39)

where

 

(40)

If the in-plane and rotary inertias are omitted (i.e., ), we have

 

(41)

 

3. Finite Element Method

 

     The finite element method (FEM), known as a powerful tool for many engineering problems, has been used to compute such matters as elastic-plastic, residual and thermal stresses, and buckling and vibration analysis. Because of this, ANSYS software that is a commercial FEM program was preferred for the vibration analysis of the laminated composite plates. The Shell 99 element type was selected for the 2D modeling of solid structures in ANSYS. Initially, the plates are to get an initial estimate of the undamped natural frequencies ωn and mode shape n. The element type of Shell 99 may be used for layered applications of a structural shell model. The element has six degrees of freedom at each node; and translations in the nodal x and y directions and rotations about the nodal z-axis. This element is constituted by layers designated by numbers (LN-layer number), increasing from the bottom to the top of the laminate; the last number quantifies the existing total number of layers in the laminate (NL-total number of layers).     The boundary conditions have been applied to the nodes, i.e., the dimensions in the x and y are 400 mm for 2D, and the displacements and rotations of all nodes about the yz plane are also taken as zero. The model of the laminated plate is generated with a different number of layers (based on different side-to-thickness ratio). The boundary conditions and mesh shape are shown in Fig. 2. It is mentioned that for a free- vibration analysis, the subspace method is applied. The subspace iteration method was described in detail by Bathe [24].

 

 

Figure 2: View of a laminated composite plate with boundary conditions and mesh shape.

 

     After the mesh generation process, a laminated composite plate with six layers has 500 elements and 1,488 nodes. By increasing the number of layers, the numbers of elements and the nodes of the plates increase. The normal penalty stiffnesses of the contact element are chosen between 104 and 109.

 

4. Results and Discussions

 

     The nondimensionalized frequencies , of specially orthotropic and antisymmetric cross-ply square laminates are presented in Table 1 for modulus ratios  and 20 (G12 = G13 = 0.5E2, G23 = 0.2E2, ʋ12 = 0.25). All layers are of equal thickness. Results are presented for m,n = 1,2,3 and for when the rotary inertia is neglected.

     The fundamental frequency increases with the modular ratio. The effect of including rotary inertia is to decrease the frequency of vibration. Note that the first four frequencies for an antisymmetric cross-ply plates are (m,n) = (1,1), (1,2), (2,1) and (2,2) and  for antisymmetric laminates. Also, good agreement was found between the analytical solution and the FEM analysis.

     Fig. 3 shows a plot of fundamental frequency  versus aspect ratio a/b for symmetric (0/90)s cross-ply and antisymmetric (0/90)2 cross-ply laminates. The material properties used are , G12 = G13= 0.6E2, G23 = 0.2E2, ʋ12 = 0.25). Fig. 4 shows the effect of coupling between bending and extension on the fundamental frequencies of antisymmetric cross-ply laminates. The material properties used are , G12 = G13 = 0.5E2, G23 = 0.2E2, ʋ12 = 0.25). With an increase in the number of layers, the frequencies approach those of the orthotropic plate. The bending-stretching coupling lowers the vibration frequencies. For example, the two-layer plate has vibration frequencies about 40 percent lower than those of an eight-layer antisymmetric laminate or orthotropic plate with the same total thickness.

 

 

    

Figure 3: The nondimensionalized fundamental frequency versus the plate aspect ratio (a/b) for cross-ply laminates.

 

 

Table 1. The nondimensionalized frequencies of cross-ply laminates, according to the classical plate theory

m

n

Layup

 

(0/90)

 

(0/90)2

 

(0/90)3

Method

 

 

 

 

 

 

 

 

 

 

 

 

1

1

Analytical

 

1.066

 

0.977

 

1.359

 

1.257

 

1.445

 

1.321

FEM

 

1.254

 

1.098

 

1.536

 

1.478

 

1.613

 

1.528

Reddy [23]

 

1.183

 

0.990

 

1.479

 

1.386

 

1.545

 

1.469

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

1

2

Analytical

 

3.090

 

2.697

 

3.987

 

3.851

 

4.136

 

4.087

FEM

 

3.265

 

2.861

 

4.182

 

4.023

 

4.351

 

4.295

Reddy [23]

 

3.174

 

2.719

 

4.077

 

3.913

 

4.274

 

4.158

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

2

1

Analytical

 

3.090

 

2.697

 

3.987

 

3.851

 

4.136

 

4.087

FEM

 

3.265

 

2.861

 

4.182

 

4.023

 

4.351

 

4.295

Reddy [23]

 

3.174

 

2.719

 

4.077

 

3.913

 

4.274

 

4.158

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

2

2

Analytical

 

4.266

 

3.911

 

5.812

 

5.449

 

5.985

 

5.774

FEM

 

4.882

 

4.078

 

6.081

 

5.631

 

6.227

 

5.937

Reddy [23]

 

4.733

 

3.959

 

5.918

 

5.547

 

6.179

 

5.877

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

3

1

Analytical

 

6.542

 

5.747

 

8.537

 

8.329

 

8.989

 

8.778

FEM

 

6.741

 

5.829

 

8.773

 

8.546

 

9.231

 

9.097

Reddy [23]

 

6.666

 

5.789

 

8.698

 

8.456

 

9.136

 

8.998

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

3

2

Analytical

 

7.386

 

6.014

 

9.894

 

9.423

 

10.398

 

9.997

FEM

 

8.011

 

6.271

 

10.125

 

9.696

 

10.553

 

10.201

Reddy [23]

 

7.927

 

6.193

 

10.034

 

9.507

 

10.494

 

10.088

 

 

 

 

 

Figure 4: The nondimensionalized fundamental frequency versus the plate aspect ratio (a/b) for antisymmetric cross-ply laminates.

 

Fig. 5 shows a plot of fundamental frequency  versus the modulus ratio E1/E2 for antisymmetric (0/90) cross-ply laminates for various values of plate aspect ratios. The plate aspect ratio lowers the vibration frequencies. The rectangle plate has vibration frequencies about 50 percent lower than those of a square plate with the same total thickness.

 

 

Figure 5: The nondimensionalized fundamental frequency versus the modulus ratio for antisymmetric cross-ply laminates for various plate aspect ratios.

 

The nondimensionalized fundamental frequencies  of graphite-epoxy composites with , G12/E2 = 0.5, ʋ12 = 0.25 and a/b = 1 are shown as a function of the lamination angle in Fig. 6. The bending-stretching coupling due to the presence of B16 and B26 lowers the frequencies. The coupling is the maximum for two-layer plates, and it rapidly decreases with increasing number of layers. At θ = 45°, the fundamental frequency of the two-layer plate is about 40 percent lower than that of the eight-layer laminate.

The nondimensionalized fundamental frequencies of graphite-epoxy composites with , G12/E2 = 0.5, ʋ12 = 0.25 are shown as a function of plate aspect ratios in Fig. 7. With an increase in the number of layers, the frequencies approach those of the orthotropic plate. The bending-stretching coupling lowers the vibration frequencies. For example, the two-layer plate has vibration frequencies about 40 percent lower than those of a four-layer antisymmetric angle-ply laminate or orthotropic plate with the same total thickness and aspect ratio. Also, effects of an aspect ratio on the fundamental frequencies of a laminated composite with same total thickness are more significant for values less than 1.

 

 

Figure 6: The nondimensionalized fundamental frequency versus the lamination angle of antisymmetric angle-ply square laminates.

 

 

Figure 7: The nondimensionalized fundamental frequency versus the plate aspect ratio of antisymmetric angle-ply laminates.

 

Figure 8: The nondimensionalized fundamental frequency versus the modulus ratio of antisymmetric angle-ply square laminates.

 

Fig. 8 shows a nondimensionalized fundamental frequency versus a modulus ratio of antisymmetric angle-ply square laminates. The effect of coupling is significant for all modulus ratios and the difference between the two-layer solution and orthotropic solution increases with the modulus ratio.

Table 2 and 3 contain the nondimensionalized fundamental frequencies  for symmetric cross-ply laminates using the FSDT. The effect of the shear correction factor is to decrease the frequencies. The smaller the K, the smaller the frequencies are. The rotary inertia (RI) also decreases the frequencies.

Fig. 9 shows the effect of transverse shear deformation and rotary inertia on the fundamental natural frequencies of orthotropic and symmetric cross-ply (0/90/90/0) square plates with the following lamina properties:

, G12 = G13 = 0.5E2, G23 = 0.2E2, ʋ12 = 0.25.

The symmetric cross-ply plate behaves much like an orthotropic plate. The effect of rotary inertia is negligible in the FSDT and, therefore, is not shown in the figure.

Fig. 10 shows the effect of transverse shear deformation, bending-extensional coupling and rotary inertia on the fundamental natural frequencies of two-layer and eight-layer antisymmetric cross-ply laminates ( , G12 = G13 = 0.5E2, G23 = 0.2E2, ʋ12 = 0.25).

 

 

 

 

 

 


 

Table 2. The effect of shear deformation on the dimensionless natural frequencies of simply supported symmetric cross-ply plates.

a/h

 

Theory

 

0o

 

Three-ply

 

Five-ply

 

Seven-ply

 

Nine-ply

5

 

FSDT

 

 8.388­

 

8.094

 

8.569

 

8.673

 

8.713

 

 

 

9.019

 

8.698

 

9.197

 

9.312

 

9.357

 

 

 

9.534

 

9.196

 

9.706

 

9.829

 

9.877

 

FEM

 

9.643

 

9.234

 

9.857

 

9.997

 

10.017

 

CLPT

 

14.750

 

14.750

 

14.750

 

14.750

 

14.750

10

 

FSDT

 

12.067

 

11.730

 

12.167

 

12.290

 

12.342

 

 

 

12.540

 

12.223

 

12.621

 

12.735

 

12.783

 

 

 

12.890

 

12.592

 

12.956

 

13.062

 

13.107

 

FEM

 

12.901

 

12.668

 

13.078

 

13.215

 

13.295

 

CLPT

 

15.104

 

15.104

 

15.104

 

15.104

 

15.104

20

 

FSDT

 

14.220

 

14.042

 

14.229

 

14.288

 

14.312

 

 

 

14.411

 

14.254

 

14.412

 

14.461

 

14.461

 

 

 

14.542

 

14.402

 

14.538

 

14.580

 

14.598

 

FEM

 

14.568

 

14.523

 

14.638

 

14.705

 

14.712

 

CLPT

 

15.197

 

15.197

 

15.197

 

15.197

 

15.197

25

 

FSDT

 

14.569

 

14.433

 

14.563

 

14.604

 

14.621

 

 

 

14.700

 

14.582

 

14.688

 

14.722

 

14.737

 

 

 

14.789

 

14.682

 

14.774

 

14.803

 

14.815

 

FEM

 

14.812

 

14.723

 

14.835

 

14.907

 

14.918

 

CLPT

 

15.208

 

15.208

 

15.208

 

15.208

 

15.208

50

 

FSDT

 

15.079

 

15.015

 

15.052

 

15.063

 

15.068

 

 

 

15.115

 

15.057

 

15.086

 

15.096

 

15.100

 

 

 

15.139

 

15.085

 

15.110

 

15.117

 

15.121

 

FEM

 

15.238

 

15.128

 

15.262

 

15.236

 

15.240

 

CLPT

 

15.223

 

15.223

 

15.223

 

15.223

 

15.223

100

 

FSDT

 

15.215

 

15.173

 

15.183

 

15.186

 

15.187

 

 

 

15.225

 

15.184

 

15.192

 

15.194

 

15.195

 

 

 

15.231

 

15.191

 

15.198

 

15.200

 

15.200

 

 

 

15.312

 

15.284

 

15.293

 

15.301

 

15.302

 

CLPT

 

15.227

 

15.227

 

15.227

 

15.227

 

15.227

 

­The first line corresponds to the shear correction coefficient of K=2/3 and the second and third lines correspond to the shear correction coefficient of K=5/6 and K=1.0, respectively.

Table 3. The effect of shear deformation, rotary inertia and the shear correction coefficient on the dimensionless natural frequencies of simply supported symmetric cross-ply (0/90/0) plates

 

a/h

 

m

 

n

 

CLPT w/o RI

 

CLPT with RI

 

FSDT w/o RI

 

FSDT with RI

10

 

1

 

1

 

15.228

 

1.104

 

12.593

 

 12.573­

 

 

 

 

 

 

 

 

 

 

12.223

 

12.163

 

 

1

 

2

 

22.877

 

22.421

 

19.440

 

19.203

 

 

 

 

 

 

 

 

 

 

18.942

 

18.729

 

 

1

 

3

 

40.229

 

38.738

 

32.496

 

31.921

 

 

 

 

 

 

 

 

 

 

31.421

 

30.932

 

 

2

 

1

 

56.885

 

55.751

 

33.097

 

32.931

 

 

 

 

 

 

 

 

 

 

31.131

 

30.991

 

 

2

 

2

 

60.911

 

59.001

 

36.786

 

36.362

 

 

 

 

 

 

 

 

 

 

34.794

 

34.434

 

 

1

 

4

 

66.754

 

62.526

 

48.837

 

47.854

 

 

 

 

 

 

 

 

 

 

46.714

 

45.923

 

 

2

 

3

 

71.522

 

67.980

 

45.484

 

44.720

 

 

 

 

 

 

 

 

 

 

43.212

 

42.585

 

 

 

 

 

 

 

 

 

 

 

 

 

100

 

1

 

1

 

15.228

 

15.227

 

15.192

 

15.191

 

 

 

 

 

 

 

 

 

 

15.185

 

15.183

 

 

1

 

2

 

22.877

 

22.873

 

22.831

 

22.827

 

 

 

 

 

 

 

 

 

 

22.822

 

22.817

 

 

1

 

3

 

40.299

 

40.283

 

40.190

 

40.147

 

 

 

 

 

 

 

 

 

 

40.169

 

40.153

 

 

2

 

1

 

56.885

 

56.874

 

56.330

 

56.319

 

 

 

 

 

 

 

 

 

 

56.221

 

56.210

 

 

2

 

2

 

60.911

 

60.891

 

60.342

 

60.322

 

 

 

 

 

 

 

 

 

 

60.230

 

60.211

 

 

1

 

4

 

66.754

 

66.708

 

66.466

 

66.421

 

 

 

 

 

 

 

 

 

 

66.409

 

66.364

 

 

2

 

3

 

71.522

 

71.484

 

70.919

 

70.882

 

 

 

 

 

 

 

 

 

 

70.801

 

70.764

­The first line corresponds to the shear correction coefficient of K=1.0 and the second line corresponds to the shear correction coefficient of K=5/6.

 

 

 

Figure 9: The nondimensionalized fundamental frequency versus the side-to-thickness ratio for simply supported orthotropic and symmetric cross-ply (0/90/90/0) laminates.

 

. The eight-layer antisymmetric cross-ply plate behaves much like an orthotropic plate. The effect of rotary inertia is negligible in the FSDT and, therefore, is not shown in the figure.

 

 

 

 

Table 4 contains numerical values of the fundamental frequencies of antisymmetric cross-ply laminated plates for various modular ratios. The results for both two-layer and eight-layer laminated plates for square and rectangular geometries are presented.

 

 

Figure 10: The nondimensionalized fundamental frequency versus the side-to-thickness ratio for simply supported orthotropic and antisymmetric cross-ply (0/90) laminates.

 

 

 

Table 4. The effect of shear deformation on the nondimensionalized fundamental frequencies of simply supported antisymmetric cross-ply plates.

 

 

 

 

 

 

 

 

 

b/h

 

Theory

 

 

 

 

 

 

 

 

 

 

 

 

Square plate (a/b=1)

 

 

 

 

 

 

 

 

 

 

 

10

 

FSDT

 

7.530

 

9.507

 

8.990

 

12.683

 

10.122

 

14.611

 

 

CLPT

 

7.832

 

10.268

 

9.566

 

14.816

 

11.011

 

18.265

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

100

 

FSDT

 

7.927

 

10.345

 

9.688

 

14.913

 

11.152

 

18.366

 

 

CLPT

 

7.931

 

10.354

 

9.695

 

14.941

 

11.163

 

18.419

Rectangular plate (a/b=3)

 

 

 

 

 

 

 

 

 

 

 

10

 

FSDT

 

4.780

 

6.341

 

5.988

 

8.824

 

6.884

 

10.290

 

 

CLPT

 

4.930

 

6.772

 

6.324

 

10.201

 

7.437

 

12.738

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

100

 

FSDT

 

4.962

 

6.798

 

6.367

 

10.231

 

7.487

 

12.763

 

 

CLPT

 

4.964

 

6.804

 

6.372

 

10.249

 

7.493

 

12.798

Rectangular plate (a/b=5)

 

 

 

 

 

 

 

 

 

 

 

10

 

FSDT

 

4.631

 

6.209

 

5.863

 

8.707

 

6.769

 

10.170

 

 

CLPT

 

4.825

 

6.612

 

6.332

 

10.200

 

7.582

 

12.729

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

100

 

FSDT

 

4.806

 

6.657

 

6.236

 

10.108

 

7.364

 

12.640

 

 

CLPT

 

4.809

 

6.667

 

6.243

 

10.117

 

7.371

 

12.643

 

 

 

 

 

5. Conclusions

 

     Analytical and numerical solutions for the free vibration of laminated polymeric composite plates with different layups are compaired based on different plate theories. Also, the effects of some parameters on the fundamental frequencies of laminated plate were performed. As a verification method, an FEM was applied with ANSYS to compare the results with those obtained from a closed-form solution. Based on the results observed, the following comments are as such:

  • The fundamental frequency increases with the modular ratio. The effect of including rotary inertia is to decrease the frequency of vibration.
  • The bending-stretching coupling lowers the vibration frequencies.
    • The plate aspect ratio lowers the vibration frequencies. The rectangle plate has vibration frequencies about 50 percent lower than those of a square plate with the same total thickness.
    • The effect of the shear correction factor is to decrease the frequencies. The smaller the K, the smaller the frequencies. The rotary inertia (RI) also decreases frequencies.
    • In all cases, results obtained from the FEM are in good agreement with analytical outputs. Also, it is shown that the FEM predicted higher values, which is reported in the literature review presented [14, 18-19].
    • The finite element model presented has an acceptable accuracy for utilizing this model for the analysis of more complicated cases.

 

References

 

[1]      Campbell FC. Structural composite materials, ASM international; 2011.

[2]      Tan p, Nie JG. Free and forced vibration of variable stiffness composite annular thin plates with elastically restrained edges. Compos Struct 2016; 149: 398-407.

[3]      Zhang LW, Zhang Y, Zou GL, Liew KM. Free vibration analysis of triangular CNT-reinforced composite plates subjected to in-plane stresses using FSDT element-free method. Compos Struct 2016; 149: 247-260.

[4]      Chakraborty S, Mandal B, Chowdhury R, Chakrabarti A. Stochastic free vibration analysis of laminated composite plates using polynomial correlated function expansion. Compos Struct 2016; 135: 236-249.

[5]      Ganesh S, Kumar KS, Mahato PK. Free Vibration Analysis of Delaminated Composite Plates Using Finite Element Method. Int Conf Vib Problems; 2015. 1067-1075.

[6]      Mantari JL, Ore M. Free vibration of single and sandwich laminated composite plates by using a simplified FSDT. Compos Struct 2015; 132: 952-959.

[7]      Su Z, Jin G, Wang X. Free vibration analysis of laminated composite and functionally graded sector plates with general boundary conditions. Compos Struct 2015; 132: 720-736.

[8]      Zhang LW, Lei ZX, Liew KM. Free vibration analysis of functionally graded carbon nanotube-reinforced composite triangular plates using the FSDT and element-free IMLS-Ritz method. Compos Struct 2015; 120: 189-199.

[9]      Marjanivic M, Vuksanovic D. Layerwise solution of free vibrations and buckling of laminated composite and sandwich plates with embedded delamination. Compos Struct 2014; 108: 9-20.

[10]   Boscolo M. Analytical solution for free vibration analysis of composite plates with layer-wise displacement assumptions. Compos Struct 2013; 100: 493-510.

[11]   Qu Y, Wu S, Li H, Meng G. Three-dimensional free and transient vibration analysis of composite laminated and sandwich rectangular parallelepipeds: Beams, plates and solids. Compos Part B: Eng 2015; 73: 96-110.

[12]   Rafiee M, Liu XF, He XQ, Kitipornchai S. Geometrically nonlinear free vibration of shear deformable piezoelectric carbon nanotube/fiber/polymer multiscale laminated composite plates. J Sound Vib 2014; 333 (14): 3236-3251.

[13]   Akhras G, Li W. Stability and free vibration analysis of thick piezoelectric composite plates using spline finite strip method. Int J Mech Sci 2011; 53 (8): 575-584.

[14]   Grover N, Singh BN, Maiti DK. Analytical and finite element modeling of laminated composite and sandwich plates: An assessment of a new shear deformation theory for free vibration response. Int J Mech Sci 2013; 67: 89-99.

[15]   Jafari RA, Abedi M, Kargarnovin MH, Ahmadian MT. An analytical approach for the free vibration analysis of generally laminated composite beams with shear effect and rotary inertia. Int J Mech Sci 2012; 65 (1): 97-104.

[16]   Tai H, Kim S. Free vibration of laminated composite plates using two variable refined plate theories. Int J Mech Sci 2010; 52 (4): 626-633.

[17]   Srinivasa CV, Suresh JY, kumar WP. Experimental and finite element studies on free vibration of skew plates. Int J Adv Struct Eng 2014; 48 (6): 123-129.

 

 

[18]   Ramu I, Mohanty SC. Study on Free Vibration Analysis of Rectangular Plate Structures Using Finite Element Method. Proc Eng2012; 38: 2758-2766.

[19]   Chandrashekhara K. Free vibration of composite beams including rotary inertia and shear deformation. Compos Struct 1990; 14 (4): 269-279.

[20]   Ke L, Yang J, Kitipornchai S. Nonlinear free vibration of functionally graded carbon nanotube-reinforced composite beams. Compos Struct 2010; 92 (3): 676-683.

[21]   Song O. Free Vibration of Anisotropic Composite Thin-Walled Beams of Closed Cross-Section Contour. J Sound Vib 1993; 167 (1): 129-147.

[22]    Lee J. Free vibration analysis of delaminated composite beams. Comput Struct 2000; 74 (2): 121-129.

[23]   Reddy JN. Mechanics of laminated composite plates and shells, theory and analysis. 2nd ed. CRC Press; 2003.

[24]   Bathe KJ. Finite element procedures. Prentie-Hall, Englewood cliffs; 1996.

[1]      Campbell FC. Structural composite materials, ASM international; 2011.

[2]      Tan p, Nie JG. Free and forced vibration of variable stiffness composite annular thin plates with elastically restrained edges. Compos Struct 2016; 149: 398-407.

[3]      Zhang LW, Zhang Y, Zou GL, Liew KM. Free vibration analysis of triangular CNT-reinforced composite plates subjected to in-plane stresses using FSDT element-free method. Compos Struct 2016; 149: 247-260.

[4]      Chakraborty S, Mandal B, Chowdhury R, Chakrabarti A. Stochastic free vibration analysis of laminated composite plates using polynomial correlated function expansion. Compos Struct 2016; 135: 236-249.

[5]      Ganesh S, Kumar KS, Mahato PK. Free Vibration Analysis of Delaminated Composite Plates Using Finite Element Method. Int Conf Vib Problems; 2015. 1067-1075.

[6]      Mantari JL, Ore M. Free vibration of single and sandwich laminated composite plates by using a simplified FSDT. Compos Struct 2015; 132: 952-959.

[7]      Su Z, Jin G, Wang X. Free vibration analysis of laminated composite and functionally graded sector plates with general boundary conditions. Compos Struct 2015; 132: 720-736.

[8]      Zhang LW, Lei ZX, Liew KM. Free vibration analysis of functionally graded carbon nanotube-reinforced composite triangular plates using the FSDT and element-free IMLS-Ritz method. Compos Struct 2015; 120: 189-199.

[9]      Marjanivic M, Vuksanovic D. Layerwise solution of free vibrations and buckling of laminated composite and sandwich plates with embedded delamination. Compos Struct 2014; 108: 9-20.

[10]   Boscolo M. Analytical solution for free vibration analysis of composite plates with layer-wise displacement assumptions. Compos Struct 2013; 100: 493-510.

[11]   Qu Y, Wu S, Li H, Meng G. Three-dimensional free and transient vibration analysis of composite laminated and sandwich rectangular parallelepipeds: Beams, plates and solids. Compos Part B: Eng 2015; 73: 96-110.

[12]   Rafiee M, Liu XF, He XQ, Kitipornchai S. Geometrically nonlinear free vibration of shear deformable piezoelectric carbon nanotube/fiber/polymer multiscale laminated composite plates. J Sound Vib 2014; 333 (14): 3236-3251.

[13]   Akhras G, Li W. Stability and free vibration analysis of thick piezoelectric composite plates using spline finite strip method. Int J Mech Sci 2011; 53 (8): 575-584.

[14]   Grover N, Singh BN, Maiti DK. Analytical and finite element modeling of laminated composite and sandwich plates: An assessment of a new shear deformation theory for free vibration response. Int J Mech Sci 2013; 67: 89-99.

[15]   Jafari RA, Abedi M, Kargarnovin MH, Ahmadian MT. An analytical approach for the free vibration analysis of generally laminated composite beams with shear effect and rotary inertia. Int J Mech Sci 2012; 65 (1): 97-104.

[16]   Tai H, Kim S. Free vibration of laminated composite plates using two variable refined plate theories. Int J Mech Sci 2010; 52 (4): 626-633.

[17]   Srinivasa CV, Suresh JY, kumar WP. Experimental and finite element studies on free vibration of skew plates. Int J Adv Struct Eng 2014; 48 (6): 123-129.

[18]   Ramu I, Mohanty SC. Study on Free Vibration Analysis of Rectangular Plate Structures Using Finite Element Method. Proc Eng2012; 38: 2758-2766.

[19]   Chandrashekhara K. Free vibration of composite beams including rotary inertia and shear deformation. Compos Struct 1990; 14 (4): 269-279.

[20]   Ke L, Yang J, Kitipornchai S. Nonlinear free vibration of functionally graded carbon nanotube-reinforced composite beams. Compos Struct 2010; 92 (3): 676-683.

[21]   Song O. Free Vibration of Anisotropic Composite Thin-Walled Beams of Closed Cross-Section Contour. J Sound Vib 1993; 167 (1): 129-147.

[22]    Lee J. Free vibration analysis of delaminated composite beams. Comput Struct 2000; 74 (2): 121-129.

[23]   Reddy JN. Mechanics of laminated composite plates and shells, theory and analysis. 2nd ed. CRC Press; 2003.

[24]   Bathe KJ. Finite element procedures. Prentie-Hall, Englewood cliffs; 1996.