7- JournalFoam-Core Effect on the Integrity of Tailplane Leading Edge During Bird-Strike Event of Aircraft.pdf

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Foam-Core Effect on the Integrity of Tailplane Leading EdgeDuring Bird-Strike Event

Reza Hedayati and Saeed Ziaei-Rad

Isfahan University of Technology, 84156-83111 Isfahan, IranDOI: 10.2514/1.C031451

Theobjectiveof this paper is to describe theprocedure of optimizing theleading-edge structureof a tailplanein an

industrial environment.The paper also investigatesthe effectof implementinga foam core between aluminum sheets

in a tailplane leading-edge structure. Bird strike against twotypes of leading-edge structures, one with and theother

without a foam core, was investigated and then the optimum design for each case was determined. The results

indicate that if a foam core is embedded between the aluminum sheets instead of increasing the thicknesses of

aluminum sheets, the skin overall weight can be reduced by 32%.

I. Introduction

A

IRCRAFT collisions with birds and other wildlife are an

increasing concern for the aviation industry worldwide [1].Civil and military aviation communities have long recognized thatthe threat to human health and safety from aircraft collisions withwildlife (wildlife strikes) is real and increasing. The number ofstrikes annually reportedmore than quadrupled from1759 in 1990 to7516 in 2008. The increase in the number of strikes may be due toseveral factors: an increase in number of flights, an increase inpopulations of hazardous birds, and more awareness of bird-strikehazards, which results in reporting the bird-strike events morecarefully [2]. Therefore, refinements in existing designs seem to benecessary.

In the early times of designing bird-proof structures, experimentaltest and theoretical calculations were used in order to predict loadsand pressures imposed by birds, and therefore the damage incurred,on different types of aircraft structures. The experimental tests have

the disadvantage of being expensive and time-consuming, whereasthe theoretical methods have the drawbacks of being inaccurate andneeding brainstorming for different solutions for different problems.The advent of efficient computers and advanced numericaltechniques made it possible to use numerical simulations instead ofinaccurate theoretical calculations for modeling real bird-strikeinteractions and loads on different aircraft components since the1980s. Explicit nonlinear finite element (FE) codes, which areavailable in several advanced commercial FE solvers (such as LS-DYNA, PAM-CRASH, ABAQUS, etc.), have been employed tosolve this class of problems.

In 1977, Wilbeck [3] found out that during high-speed impacts,birds behave as afluid. He also showed that the fluid behavior of abird during impact can be explained in terms of a roughly circularcylindrical body composed of gelatin with 10% air porosity. Sincethen, many authors have employed this or similar material andgeometrical properties for bird-impact problems. Many researchers(such as Frischbier [4], Langrand et al. [5], McCarthy et al. [6], andAiroldi and Cacchione [7]) have simplified the bird torso as ahemispherical-ended cylinder. The ellipsoid geometry is also a well-accepted choice, which has been suggested by the International BirdStrike Research Group [8], and has been used by Guan et al. [9]. In

addition to these two configurations, the straight-ended cylinder hasalso been adopted by Stoll and Brockman [10], but its applicationremains somewhat infrequent.

In 2004, McCarthy et al. [6] designed, built, and subjected aircraftwing leading-edge structures with a glass-based fiber/metal laminateskin to bird-strike tests and then modeled them with finite elementanalysis. In 2005, Kermanidis et al. [11] proposed a novel design of afiber-reinforced composite leading edge of a horizontal tail plane. Intheir work, numerical modeling issues and critical parameters of thesimulation were discussed.

Reyes et al. [12] implemented the material model presented byDeshpande and Fleck [13] as a user subroutine for evaluatingbehavior of aluminum foams under different mechanical conditions.Shortly thereafter, Hanssen et al. [14] carried out a number ofexperimental bird-strike tests on double sandwich panels made fromAlSi7Mg0:5 aluminum foam core and aluminum Al 2024 T3 coverplates. They also created numerical FE models and showed that the

numerical explicit method correlates well with experimental data.Generally, there are three main approaches for modeling a bird inan impact event numerically: Lagrangian approach, Eulerian,arbitrary LagrangianEulerian (ALE) approach, and smoothedparticle hydrodynamics (SPH) approach. LS-DYNA, a high- andlow-impact dynamicsfiniteelement software, has implemented theseformulations to model fluidstructure interaction problems. Themain differences in these four formulations are the referencecoordinate system used for describing the motion and the governingequations for the movement of materials. In the ALE approach, thebird material flows relative to an Eulerian mesh, thereby avoidinglarge mesh distortion. The impacting loads are then transferred to theLagrangian mesh of theimpacted structure through an ALEcouplingalgorithm.

The Lagrangian method uses material coordinates (also known as

Lagrangian coordinates) as the reference. The Lagrangianformulation is used mostly to describe solid materials. However, aLagrangian description of the bird-strike problem may result in lossof birdmass,due tothefluid behavior of the bird, which causeslargedistortions in the bird. Considering that the bird as afluid mass is thecause of large distortions; consequently, these distortions are thecause of large variations about volumetric strain (which measuresthe ratio of change of the bodys volume) in some elements of themodeled bird. This loss of mass may reduce the real loads applied inthe impact.

Solvers based on the SPH method have recently been developedand implemented in the framework of an explicit FE code to analyzeevents characterized by large deformations. SPH formulation is ameshless Lagrangian technique used to model the fluid equations ofmotion using a pseudoparticle interpolation method to computesmooth hydrodynamic variables. In this formulation, the fluid isrepresented as a set of moving particles, each representing aninterpolation point, where all the fluid properties are known. Then,

Received 25 March 2011; revision received 30 May 2011; accepted forpublication 31 May 2011. Copyright 2011 by the American Institute ofAeronautics and Astronautics, Inc. All rights reserved Copies of this papermaybe made forpersonal or internal use, on condition that the copier pay the$10.00 per-copy fee to the Copyright Clearance Center, Inc., 222 RosewoodDrive, Danvers, MA 01923; include the code 0021-8669/11 and $10.00 in

correspondence with the CCC.Department of Mechanical Engineering; [email protected](Corresponding Author).

Department of Mechanical Engineering; [email protected].

JOURNAL OF AIRCRAFTVol. 48, No. 6, NovemberDecember 2011

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with a regular interpolation function called smoothing length, thesolution of the desired quantities can be calculated for all theparticles. In the SPH approach to the bird-strike problem, the bird ismodeled using SPH particles, and the structure is constructed usingLagrangian finite elements, such as the Lagrangian and ALEapproaches.

In this paper, the Lagrangian method has been used to solve thebird-strike problem. Numerical difficulties associated with largedistortion of elements have been overcome by using hourglass and

bulk viscosity controls. The objective of the paper is then to describethe procedure of calculating an optimized leading-edge structuredesign in an industrial environment. The effect of implementing afoam core between aluminum sheets in a leading-edge structure isalso investigated in detail.

In this study, thecreatednumerical model is validatedin twosteps,which are 1) verification of the bird models with differentformulations against a rigid target: the proposed bird model and theLagrangian formulation were verified by comparing the obtainedresults with the experimental results reported by Wilbeck [3], and2) verification of the bid strike against a simple foam plate: thenumerical results of this part are compared with the experimentaldata of Hanssen et al. [14], and the validated model is then used topredict the effect of the foam core on the integrityof tailplane leadingedge during bird-strike event. After validating the numerical model,

it was used to forecast the behavior of tailplane structure with andwithout a foam core. Finally, the optimum thickness for aluminumsheets in a leading-edge structure without a foam core is evaluated.The optimum foam-core density in the leading-edge structure with afoam core is also established.

II. Leading-Edge Modeling

A. Leading-Edge Component Discretization

In Fig. 1 the setup for a bird model impacting the tailplane leadingedge is shown. The leading-edge structure consists of seven parts:

inner andouter skins, spar, ribs, back frame, backbeams, anda foam-core layer in some cases. In most of the leading-edge structurecomponents, elements with sizes of 5 to 7 mm have been used. Thenumerical model of leading-edge components is shown in Fig. 2.

The beam structure was specified by a large elasticity modulus inorder to model the solidity of the support structure in back of theleading edge. The inner and outer aluminum skin sheets weremodeled using four-noded shell elements with thicknesses rangedfrom 0.4 to 1.4 mm for different simulations. The inner and outer

skins consisted of 14,775 and 14,184 elements, respectively. Insimulations in which the structure has a foam core between thealuminum sheets, the foam core consisted of 32,340 eight-nodedsolid elements with an average size of 8 mm. The spar that providesgood resistance against bird impact was created with 3229 shellelements. Since the ribs are thin, they were also discretized usingshell elements. The ribs consisted of 5719 shell elements with0.8 mm thicknesses. In simulations with or without foam cores, thedistancesbetween the center of the inner and outer skins were set to 4and 6.42 mm, respectively. The total thickness of the foam core wasset to 5 mm.

B. Leading-Edge Material Modeling

In most leading-edge components (such as skins, ribs, spar, and

frame), aluminum 2024 has been used as the material. Theconstitutive model assumed for aluminum parts is isotropic anelasticplastic material model. A CowperSymonds law has alsobeen included to model the strain-rate sensitivity of the yield stress:

ny

1

_"

H

1p

(1)

where n is the dynamic stress, y is the yield stress in quasi-staticconditions, parameters p and H are the constants of CowperSymonds law, and _" is the strain rate of the material. The propertiesused for the aluminum alloy 2024 are listed in Table 1 [15].

In some simulations, aluminum foam has been embedded betweenaluminum skins. Here, aluminum foam has been modeled using theDeshpandeFleck material model [12,13]. In this material model,

yield function is defined by Y, where yield stress Ycan bestated as

Y p R" (2)

where R" represents the strain hardening, and " and p are theequivalent strain and foam plateau stress, respectively. Theequivalent stress is given by [13]

2 1

1 322VM

22m (3)Fig. 1 Setup of the bird model impacting the tailplane leading edge.

Fig. 2 Exploded view of the FE model.

HEDAYATI AND ZIAEI-RAD 2081
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where VM is the effective von Mises stress that is given by

VM

32dev:dev

q. Parameters dev and m are representatives of

deviatoricstress tensorand mean stress, respectively, whichare givenby formulas m tr and dev mI. The parameter defines the shape of the yield surface and can be linked tocompression plastic coefficientp by

2 91 2p

21 p! true (4)

Yield stress at any instance is given by

Y p "

"D 2 ln

1

1 ""D

(5)

where 2, , and are the hardening parameters. In Eq. (5), theequivalent strain " can be given by

" 2

1

3

2

"2e 1

2"2m

(6)

where "e and "m are the von Mises effective strain and hydrostaticstrain, respectively, and are given by

"m

2"

1 32

m

(7)

"e "

1 32

e

(8)

where e is the von Mises effective stress. From uniaxialcompression, the compaction strain "D is given by

"D 9 2

32ln

f

f0

(9)

where f and f0 are the foam density and base material density,respectively. In this finite element analysis, failure is modeled bydeleting or eroding elements in which the hydrostatic strain passes

the critical hydrostatic strain. In other words,

If "m "crm ! element deletion (10)

Hanssen etal. [14] carriedout severalvalidationtestson aluminumfoams and compared their results with material models available inLS-DYNA. In the current study, three types of foam used by theHanssen et al. experimental tests are employed inside the sandwichpanel in the leading-edge structure. The material properties of foamsare presented in Table 2.

In the leading-edge structure, rivets have been used to connectinner skin to ribs, ribs to spar, and inner/outer skins to frame. Formodeling the rivets, a mesh-independent method, i.e., tie-break nodeto surface contact type in LS-DYNA, has been used. This contacttype makes it possible to mesh two connected plates freely withoutthe necessity of defining two riveted nodes on two connected platesin the same place. In this study, aluminum rivets with diameters of 4or 5 mm have been used. Nonlinear behavior and failure criterion forrivets can be stated as

jFNj

FNF

jFSj

FSF

1 (11)

where FN and FNF are the normal force and critical normal force inrivets, respectively, and FS and FSF are the shear force and criticalshear force in rivets, respectively. For Eq.(11), constants 1:5 and 2:1 have been used according to [16]. In this study in eachsimulation, first rivets with a diameter of 4 mm were used, and forcases in which many rivets were failed, they were replaced by 5 mmrivets. The yield criterion for each rivet type is given in Table 3.

For controlling thehourglass effect inthe simulations,a FlanaganBelytschko viscous-form hourglass type with an hourglasscoefficient of 0.14 was implemented. For the bulk viscosity control,thestandard type (type 1) wasused with quadraticcoefficient (Q1)

and linear coefficient (Q2) of 2 and 0.25, respectively. Themagnitudes of hourglassand total energyhave calculated versustimefor all the case studies (see Sec. IV.B). The results indicate that thehourglassenergy remains lowerthan 10%of total energyduring bird-strike simulation.

III. Bird Modeling

At high pressures, the bird material behaves as a hydrodynamicmaterial, for which an equation of state relating the thermodynamicproperties of pressure p and volume is adopted. So far, a number ofequations of state (such as tabulated, linear polynomial, andGrneisen) have been implemented by different authors. In thiswork, the linear polynomial equation of state is chosen because of itsprevalent usage by different researchers. The linear polynomial EOSrelates pressure and volume by means of the following equation:

p C0 C1 C22 C3

3 C4 C5 C62U (12)

where =0 1, and U is the internal energy per volume.Parameters C0 to C6 are material constants, and and o are,respectively, the current and initial densities of the material. The birdmaterial was identified for this model using the material propertiesspecified by Brockman and Held [17], whereby

C0 0; C1 2323 MPa

C2 5026 MPa; C3 15; 180 MPa (13)

The total massof the bird was set to1.8 kg, as itis the required bird

mass according to the Federal Aviation Regulations [18]. A densityof938 kg=m3 is used for the bird models, as it has been previouslyused in many investigations.

As a real bird body consists of several internal cavities, bonestructures, etc., and has a complex geometry, it is not rational tomodel the bird with a shape exactly the same as the real one. So far,researchers have presented and used a number of substitute birdgeometries, such as sphere, hemispherical-ended cylinder, straight-ended cylinder, and ellipsoid. Several studies have been carried outon the appropriate shape of the bird, and the hemispherical-ended

Table 1 Properties used for the aluminumalloy 2024

Property Value

Density 2700 kg=m3

Poisons ratio 0.27Elasticity modulus 72 GPa Yield stress 280 MPa Ultimate stress 385 MPa Failure strain 0.18

H 1:28e5 s1

p 4

Table 2 Material properties of foams

f, kg=m3 E, MPa p , MPa "D 2, MPa p, MPa "cr

150 300 0.05 1.19 2.89 52.1 3.26 0.93 0.1300 1500 0.05 6.1 2.2 38.1 3.1 4.41 0.1510 5562 0.05 5.37 1.67 66.9 2.99 14.82 0.1

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cylinder with a length-to-diameter ratio of 2 has been suggested[19,20]. Therefore, in this study, a bird model with the shape of ahemispherical-ended cylinder and a length-to-diameter ratio oftwo is used. From geometrical relationships the diameter D

8m=43 23

qis obtained, where D, m, and are the diameter,

mass, and density of the hemispherical-ended cylinder, respectively.By replacing the mass of 1.8 kg and density of 938 kg=m3, adiameter ofD 0:1136 m is evaluated.

The initial velocity of the bird is set to 124 m=s for all thecalculations afterward. The bird is discretized using 8800 uniformeight-noded fully integrated solid elements in LS-DYNA. The birdfinite element model is shown from two views in Fig. 3.

IV. Numerical Model Validations

A. Validation of the Model for Perpendicular Impact

A steel plate, 60 cm in diameter and 6 cm in thickness, is used astarget for bird-impact simulations. The large thickness of the targetplate allows it to be considered appropriately as a rigid plate. Sincethe target is assumed to be rigid and not allowed to have largedeformations, dimensions and specifications used for the target arenot crucial. In fact, the pressures captured by the sensors are ofinterest and not the target deformation itself.

A number of bird geometries (such as straight-ended cylinder,hemispherical-ended cylinder, and ellipsoid) have been used as asubstitute for the real bird geometry. Several studies have beencarried out on the appropriate shape of birds, and a hemispherical-

ended cylinder with length-to-diameter ratio of 2 has been suggested[19]. As a result, in this section the bird was modeled as a waterprojectile with the shape of a hemispherical-ended cylinder and alength-to-diameter ratio of 2. The impact velocity was set to116 m=s. In numerical simulations, authors have used densitiesranging from 900 to 950 kg=m3 for the bird material. In the currentstudy, a density of 938 kg=m3 was used for the traditional birdmodel.

Before selecting the appropriate formulation for the bird-strikesimulations, a comprehensive set of simulations was carried out.Three types of bird formulations (i.e., Lagrangian, ALE, and SPH)were considered. The hemispherical-ended cylinder bird modelswere created and then impacted to perpendicular rigid target plates.The pressure profiles were compared with Wilbecks [3]experimental test. The deformation of the bird model obtained from

different formulations with respect to time can be seen in Fig. 4. Thepressure profilesat thecenterof theimpact for three formulationsandfor different element sizes are depicted in Figs. 5a5c.

As can be seen from the figures, all three methods predict closeresults forfine mesh in comparison with each other, as well as withthe experimental tests. Since the Lagrangian formulation needs lesscomputational effort due to its simplicity, and since the impactedtailplane structure consists of numerous elements, the Lagrangian

formulation was chosen for the study to decrease the time ofnumerical solution. In addition, Lagrangian formulation has alsobeen used and verified in other bird-strike studies, such as [2124].

B. Validation of the Model for Foam Sandwich Panels

Experimental data provided by Hanssen et al. [14] have been usedin order to validate our numerical model. In Hanssen et al.sexperimental work, the impact of a 1.81 kg bird with two sandwichpanels placed on top of each other has been investigated(Fig. 6). The

inner and outer sandwich panels each consist of two 0.8-mm-thickaluminum sheets and foam cores of 20 and 33 mm thicknesses,respectively. In the adopted test, foam-core densities for twosandwich panels are the same and equal to 300 kg=m3. The initialimpact velocity of the bird has been chosen to be 139 m=s. A fewstraingaugeshavebeen installedon thebackplatein order to measurethe strain in different positions and orientations. Signals recorded bystrain gauges provided essential data for comparison betweenexperimental and numerical results. The setup and numbering ofstrain gauges on the backplate are shown in Fig. 6.

Thefinite element model of thebird and sandwich panels is shownin Fig. 7. The numerical model consists of four parts:bird (7600 solidelements), aluminum sheets (14,000 shell elements), two foam cores(21,000 solid elements), and the back sheet (1986 elements). For the

foam cores and aluminum sheets, respectively, the Deshpande

Fleckand isotropic elasticplastic material models have been used.The deformation of the bird and sandwich panels versus time is

shown in Fig. 8. In Figs. 9a9d, the strains recorded by strain gaugeshave been compared with strains of corresponding elements fromnumerical model. In Fig. 9athe experimental curve is an average ofresults of gauges 1, 2 and 3. The sharp drop at t 0:6 ms in Fig. 9dcan be attributed to the short-time malfunction of carbon gauges dueto the high speed of deformations.

In Fig. 9b the experimental curve is an averageof results of gauges4 and 5. As can be seen from Fig. 9, the numerical and experimentalresults correlate well with each other, except for case of Fig. 9a. Thiscanbe explained by the fact, as Hanssen et al. [14] also stated in theirpaper, that the bird model used for simulation is more cigar-shapedthan the real bird used in the tests. This may offer an explanation for

why higher strains are observed in the center of target where gauge 1is installed. In other words, the real bird may spread the load over alarger area, thus reducing the local strains in the central impact area.

V. Results and Analysis

A. Bird Impact Against Leading Edge Without Foam Core

After validation of our numerical model with experimental data, itwas decided to use the model for the bird impact against the tailplaneleading edge. For the first case there is no foam core between the

Table 3 Yield criteria for rivets

River type Diameter, mm Yield criterion

1 4jFNj

7200

1:5

jFS j

4500

2:1

1

2 5jFNj

12;000

1:5

jFS j

7000

2:1

1

Fig. 3 Lagrangian bird model.Fig. 4 Deformation of different bird models in the perpendicularimpact.

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leading-edge sheets. The deformation of the bird and the leading-edge structure are shown in Fig. 10. It is notable that the anglebetween the bird velocity and the leading-edge front line is 102 deg.This is due to the fact that for the sake of better aerodynamicperformances, the front lines of most airplane leading-edge struc-tures (shown by line AB in Fig. 1) are not exactly perpendicular tothe aircraft symmetry plane. For the type of airplane considered inthis study, the angle is not 90 deg, but it is 102 deg.

The deformed shape of the leading-edge structure after 5 ms fromthe beginning of impact for different configurations (i.e., differentinner and outer surface thicknesses) is shown in Fig. 11. As can beseen, after the impact process was finished, a complete penetrationoccurs in simulations 1 and 2, partial penetration occurs in simul-ations 3 and 4, and no penetration is observable in simulations 5 and6. It is remarkable that in simulation numbers 3 and4, after theimpact

duration the leading-edge deformation for both cases is almostidentical. The same trend is apparent for simulation numbers5 and6.It is noticeablethat thesums of inner andouterthicknesses forcases 3and 4 are equal. This is also true for simulations 5 and 6.

The displacement in the central node of the leading-edge structurehas been compared for different configurations in Fig. 12. As can beseen from the figure, the displacement of the outer-skin central nodedecreases when the overall thicknesses of inner and outer skinsincrease. Whereas in the simulation numbers 1 and 2, the outer-skincentral node passes through the leading-edge back beams, insimulation numbers 3, 4, 5, and6, theouter-skin central node remainsinside the leading-edge structure. In other words, in simulationnumbers 1 and 2, penetration occurs, and in the other simulationspenetration does not occur. It is noteworthy that the central-nodedisplacement profiles for simulation numbers 3 and 4, as well as

simulationnumbers 5 and 6,are close toeachother. Thiscan be againattributed to thesame overall thicknessin thesimulations(i.e.,sum of

Fig. 5 Pressure profile at the center of impact for a) Lagrangian formulation, b) SPH formulation, and c) ALE formulation.

Fig. 6 Setup and numbering of strain gauges on the backplate(Hanssen et al. [14]). Fig. 7 Numerical model of bird and sandwich panels.

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inner and outer thicknesses). The absorbed energy by the leading-edge structure is shown in Fig. 13. Increasing the overall aluminum

skins thickness, more energyis absorbed by theleading edge throughthe work of plastic deformation. This means that by increasing theoverall thickness, less bird mass can pass through the leading edge ormore bird kinetic energy is convertedto theplastic deformation of thestructure.

B. Bird Impact Against Leading Edge with Foam Core

In this subsection, the effect of embedding aluminum foam coreswith densities of 150, 300, and 500 kg=m3 between the inner andouter aluminum sheets is investigated. In four simulations conducted

Fig. 8 Deformation of bird and sandwich panels over time.

Fig. 9 Comparison of strain recorded in experimental tests and simulations for strain gauges numbered a) 1, b) 4 and 5, c) 6, and d) 7.

Fig. 10 Deformation of bird and leading-edge structure.

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here (one with a foam core and three without a foam core), thethicknesses of the inner and outer aluminum sheets are identical andequal to 0.6 mm. Therefore, the only difference between thesimulations is the foam-core density. The deformed shape of theleading edge after 5 ms for four configurations is shown in Fig. 14. Itis remarkable that in the simulation without a foam core, the spar isvisible after the impact process (i.e., bird penetration), whereas it isnot visible in the simulations possessing a foam core.

The central-node displacement has been compared for fourconfigurations in Fig. 15. As could be predicted, the displacement

decreases by increasing the foam-core density. In the simulationshaving no foam core and having low-density foam core, the centralnode passes through the back beams, i.e., the penetration occurs,whereas in the simulations having medium- and high-density foamcores, no penetration occurs. The amount of bird kinetic energyabsorbed by the leading edge for the four configurations is compared

in Fig. 16. As can be seen, for the simulations with no foam core andpossessing low- and medium-density foam cores, the absorbedenergy increases when the foam density increases. However, theabsorbed energy in the simulation with high-density foam hasslightly decreased. The explanation is that in the simulation withhigh-densityfoam, thehigher solidity of thestructurecausesless birdmass toenter theleading-edgestructure and makes thebird pass bythe leading edge, whereby less energy is absorbed by it.

C. Effect of Impact Incidence Angle

Since birds mostly fly in low altitudes, more than 60% of birdstrikes happen at heights lower than 100 m [25]. Thus, most of birdstrikes take place during the takeoff/landing. Most of commercial

and transport airplanes can take a maximum angle of 15 deg with

Fig. 11 Deformed shape of leading-edge structure for different

confi

gurationsafter 5 ms from the beginningof impact(I.T.is inner-skinthickness, O.T. is outer-skin thickness).

Fig. 12 Central-node displacement for impact of a bird with differentleading-edge configurations.

Fig. 13 Energy absorbed by the leading-edge structure from birdkinetic energy.

Fig. 14 Deformed leading-edge structure for configurations with a) no

foam core, b) low-density foam, c) medium-density foam, and d) high-density foam core.

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respect to horizon while taking off. The deformation of the bird andthe leading-edge structure components can be seenversus time for anincident angle of 15 deg in Fig. 17.

The deformed shape of the tailplane leading edge for the impact ofthe bird to the structure with or without a foam core is shown inFig. 18. It is interesting to note that for the same configurations in theinclined impacts, more damage (Fig. 18) in comparison withthe normal impacts can be observed (Figs. 11 and 13). This is due tothe factthat in an inclined impact, the spar moves upward,whereby it

does not contribute much in the resistance against the bird impact.The displacements of the central node of the leading-edge outerskin for structures with or without foam cores, as well as the sparposition with respect to the central node, are plotted in Fig. 19. It isimportant to mention that the spar position is lower for the inclinedimpact in comparison with the normal impact. This is due to the factthat in theinclined impact, thefirst point of the outer skinthatthe birdtouches is closer to the spar. In Fig. 19a, the displacements of thecentral node for the structures without a foam core and with differentouter-skin thicknesses are shown. As can be seen, the central node

Fig. 15 Central-node displacement for structures with and without afoam core.

Fig. 16 Absorbed energy by structures with and without a foam core.

Fig. 17 Deformation of bird and the leading-edge structurecomponents versus time for an impact with 15 deg angle.

Fig. 18 Deformed shape of leading-edge structures for cases with and

without a foam core after 5 ms from thebeginning ofimpact (I.T.is inner-skin thickness, O.T. is outer-skin thickness).

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passes through the back beams of the leading edge in the structureswith outer thicknesses of 0.6, 0.8, and 1 mm. However, in theleading-edge structure with an outer thickness of 1.4 mm, nopenetration occurs. In Fig. 19b, the displacements of the central nodeforthe structureswith inner andouter thicknesses of 0.6mm andwithdifferent foam-core densities are shown and compared with that ofthe structure with no foam core. As can be seen from the figure, nopenetration occurs in any of the leading-edge configurations thatinclude foam cores, whereas in the leading-edge configurationwithout a foam core, penetration occurs.

D. Effect of Foam Core on Structure Weight

One of the main factors that must be considered in an aircraftstructure design is the weight of its components. The weight of anairplane affects the amount of fuel consumption. Therefore, lightnessof a component design makes it more appealing for being employedin an airplane structure. In the previous sections, two designs that donot permit the bird penetration for leading-edge structures werechosen. One of these structures has a foam core, and the other iswithout any foam core. Two optimum designs for the structures withandwithout a foam core are presentedin Table 4.Ascanbeseenfromthe table, if a foam core is embedded between the aluminum sheetsinstead of increasing the thicknesses of aluminum sheets, the skinoverall weight can be reduced by 32%. Thus, implementing foam

cores between aluminum sheets is recommended. The change in themass of the foam core for the case that includes a foam core is shownin Fig. 20.

VI. Conclusions

In this paper,a numerical model for sandwich panel structureswasdeveloped and then verified by experimental data available in theliterature. Next, bird strikes against two types of leading-edgestructures, one including a foam core and the other without a foam

core, were investigated. In the structure without a foam core, forhaving no penetration, the overall thickness of aluminum sheets waschanged from 1.2 to2 mm. Inthe structure witha foamcore, the innerandouterthicknesses werekept constant to 0.6mm andthe density ofthe foam core was varied. In the first case, the optimum overallaluminum-sheet thicknesswas calculated, andin thesecondcase, theoptimum densities for foam cores for no penetrations werecalculated. It was found that for the case without a foam core, theoverall thickness of 2 mm for aluminum sheets is required to preventbird penetration into the structure, whereas in the case with a foamcore, the medium foam density (i.e., 300 kg=m3) was found to beadequate for not permitting any bird penetration. The results alsoindicate that if the foam core is embedded between the aluminumsheets instead of increasing aluminum-sheet thicknesses, the skinweight can be reduced by 32%. Therefore, in order to increase the

resistance of the leading-edge structure against bird strike, one canembed a foam core between aluminum sheets instead of increasingthe aluminum-sheet thickness; however, this increases the cost ofcomponent manufacturing.

In addition to the main conclusion mentioned above,the followingpoints are also worth mentioning:

1) Inclined impacts cause more damage to the leading edge incomparison with normal impacts. This is due to the fact that in theinclined impact, the spar moves upward, whereby it does notcontribute much in the resistance against bird impact.

2) By increasing the aluminum-sheet thickness and the foam-coredensity, thetotal displacement of the central node decreases, whereasthe absorbed energy by the leading-edge structure increases.

References[1] Dolbeer, R. A., Birds and Aircraft Compete for Space in Crowded

Skies, ICAO Journal, Vol. 61, No. 3, 2006, pp. 2124.

Fig. 19 Displacement of the central node for the impact of a bird to the leading-edge structure a) without a foam core and b) with a foam core.

Table 4 Optimum designs for the structures with and without foam core

Structure type Thicknesses of inner and outer sheets, mm Foam-core density, kg=m3 Overall mass of inner/outer sheets and the foam core

Without foam core 0.61.4 3.57 kgWith foam core 0.60.6 300 2.43 kg

Fig. 20 Change in mass of the foam core for the structure with a foamcore.

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7- JournalFoam-Core Effect on the Integrity of Tailplane Leading Edge During Bird-Strike Event of Aircraft.pdf