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Composite Structures 81 (2007) 450–462

Analysis of a hybrid composite/metal bolted connection subjectedto flexural loading

Vincent Caccese *, Jean-Paul Kabche, Keith A. Berube

Department of Mechanical Engineering, 5711 Boardman Hall, University of Maine, Orono, ME 04469, USA

Available online 3 November 2006

Abstract

This paper presents the structural response of hybrid composite/metal bolted joints using an experimental investigation and finite ele-ment analysis. The primary motivation for this work is to provide alternatives to conventional hull construction techniques using mod-ular hybrid construction methods. The main objective is to implement a watertight hybrid connection concept that provides access usingremovable panels on a ship hull. The hybrid joint developed is used to connect fiberglass reinforced composite panels to a metal sub-structure. An experimental study was conducted to quantify the performance of various hybrid joints with various geometries, loadedin flexure. The test results show that, for resisting bending loads, joints with metal cap plates can be made stronger and rotationally stifferthan standard bolted joints, while also mitigating opening of the joint, thereby improving the ability to seal the connection for watertightintegrity. Local shell and detailed solid finite element analyses of the connections are presented. Simplified shell finite element modelswere developed for implementation into global models of the structure and good correlation with the test results was observed withrespect to connection stiffness. Nonlinear response of the connection was predicted using detailed plane strain contact models to capturethe three-dimensional effects. The results of this study were used to select a connection geometry, which was subsequently incorporatedinto the hydrostatic testing of a full-scale four-panel assembly.� 2006 Elsevier Ltd. All rights reserved.

Keywords: Hybrid connections; Cyclic testing; Nonlinear finite element analysis; Contact analysis

1. Introduction

The implementation of composite materials in conjunc-tion with metals into hybrid structural systems is currentlybeing developed in several key applications such as ships,aircraft, buildings, bridges, automobiles and other trans-portation vehicles. A hybrid structure can potentially bemore advantageous than a single material system whencost, maintenance, weight and structural performance areconsidered simultaneously. Hybrid systems, where metalsare used as the backbone of the structure and compositesare used for the outer shell of the system, are of currentinterest in ship construction.

Traditional hull-forms are typically constructed usingsteel for large vessels and composite or aluminum for smal-

0263-8223/$ - see front matter � 2006 Elsevier Ltd. All rights reserved.

doi:10.1016/j.compstruct.2006.09.009

* Corresponding author. Tel.: +1 207 581 2131; fax: +1 207 581 2379.E-mail address: vince_caccese@umit.maine.edu (V. Caccese).

ler vessels. Using steel for hull-form construction has madeit difficult and costly to achieve complex hydrodynamicshapes, which in turn has led to higher structural weight.In naval vessels, concerns over cost, weight, stealth andcorrosion have led designers to look at advanced materialsfor hull-form construction. Although very good in fatigue,carbon steel, for instance, is magnetic, dense, and highlysusceptible to corrosion when at sea, which translates intohigher maintenance costs. Stainless steels solve the mag-netic signature and corrosion issues; however, they are alsoheavy and are quite costly. Aluminum, although light-weight, corrosion resistant and non-magnetic, is prone tofatigue failures. In light of these concerns, advanced com-posite materials have emerged as a viable alternative tothe conventional hull construction methods. Accordingly,fiber reinforced composite systems are of particular interestfor mid-size to larger ship structures, provided that theycan be cost-effective and lead to superior hydrodynamic

V. Caccese et al. / Composite Structures 81 (2007) 450–462 451

shapes, weight reduction, higher speeds, and adequatestrength and stiffness. Additionally, using composites forthe bulk of the structure will help achieve a more stealthyand corrosion-resistant structure.

The current investigation was motivated by the ModularAdvanced Composite Hull-forms (MACH) project, whichis aimed at developing watertight, hybrid, composite/metalstructural joint concepts for naval ship hull applications.This project was funded by the Office of Naval Research(ONR) and conducted as a joint effort by the Universityof Maine, Navatek Ltd., of Honolulu, Hawaii, AppliedThermal Sciences, Inc. (ATS) of Sanford, Maine, and theNaval Surface Warfare Center at Carderock, Maryland(NSWCCD). Navatek Ltd. is known for their constructionof innovative hull-forms using underwater bodies. Onesuch example is shown in Fig. 1, where a reconditioned sur-face effect ship (SES-200) was retrofit with a large liftingbody. Control surfaces are used on this structure to main-tain desired vessel characteristics and equipment is housedwithin the shell of the lifting body. The lifting body shownin this figure was constructed of an all-metal shell and sub-structure. The proposed hybrid construction method forthis and similar lifting bodies is to use an all-metal frame-work with composite panel skins. The composite panels are

Fig. 1. Reconditioned surface effect ship (S

attached to a metallic sub-frame using a hybrid joint, asshown schematically in Fig. 2. A bolted connectionbetween the shell and the framework will allow for removalof the panels, therefore making maintenance an easier task.Also, the stiffer, metallic sub-frame will facilitate theattachment of propulsion equipment when compared tothe use of an all-composite system. The design of the skinpanels and connections is dependent upon the geometry,load, and environmental requirements. The panels mustbe able to withstand the applied loads, while maintainingwatertight integrity.

1.1. Objectives

The goal of the work presented herein is to develop, testand analyze hybrid connection concepts that will allow forremovable panels, maintain watertight integrity, and haveadequate strength. Techniques to manufacture connectionspecimens using standard construction methods in a cost-effective manner were considered. Sub-component sectionsof bolted joint concepts were fabricated and tested undercyclic loading to quantify the structural response and deg-radation mechanisms in the connection. Two finite elementmodeling approaches are used to study the connection

ES-200) with underwater lifting body.

Hybrid Connection to Structural Frame

Metal Framework

Composite panel

Protective coating

Fig. 2. Schematic of the hybrid concept with composite panels over metalframework.

452 V. Caccese et al. / Composite Structures 81 (2007) 450–462

behavior: (1) a simplified shell element model used toaccount for the connection stiffness in a global sense; and(2) a detailed plane strain model used to predict the connec-tion strength. This work led toward the selection of a con-nection configuration for implementation into ahydrostatic test of a full-scale four-panel assembly. Thispaper emphasizes the finite element modeling of theselected connection.

2. Connection concepts and test articles

Berube and Caccese [1] studied numerous connectionconcepts in an attempt to narrow down the possibilitiesto the most sound and cost-effective details. The leastexpensive connection that allows removal of the panels isthe standard bolted joint. This connection was used asthe baseline for comparison of the strength, fabricationcomplexity, and cost. Two types of this joint include a pro-truding head version and a tapered head version, where thebolts on the exterior are countersunk. Concerns overwatertight integrity of the standard bolted connection ledtoward a connection called a doubler-type joint, where acontinuous metal plate acting as a cap-plate is used to

lc

wcbt

e bl

wl

wu

tl

tu

dw tw

tc

wd

Ø

td

P2

P2

Fig. 3. Geometry of the bolted hybrid joints.

cover the joint region. Fig. 3 shows the doubler concept,where countersunk bolts are used with a metal cap-plate.Versions with one and two rows of bolts per side weretested. Fig. 4 shows photographs of the joint response atthe interface for standard bolted joints (Fig. 4a) and boltedjoints with doubler plates (Fig. 4b). The addition of a dou-bler plate helps to keep the joint closed when acted upon bytransverse forces that tend to open up the joint.

A test matrix of the connections studied is presented inTable 1. Specific details of the testing are provided by Kab-che et al. [2]. Joints without doublers are designated aseither BP or BT, indicating protruding head or taperedhead bolts, respectively. A gusset was used at the end ofthe steel section in several cases to study the effect of a verystiff flange boundary. Joints with doubler plates are indi-cated by DS, DM or DL, indicating a short, medium orlong length doubler. Cost and complexity will be depen-dent on the size, thickness, and number of bolt rows usedper side and these parameters were varied as indicated inthe table. Foam inserts were added in several cases to cre-

Fig. 4. Response at connection interface. (a) Baseline standard boltedjoint. (b) Hybrid joint with doubler plate.

Table 1Geometric configuration of hybrid joint specimens

Specimen Bolt type Bolt diameter Ø (mm) No. of bolt rows per side Doubler width wd (mm) Doubler thickness td (mm) Remarks

BP-1 Protruding 12.7 2 – – –BT-2 Tapered 12.7 2 – – –BP-3 Protruding 19 2 – – –BT-4 Tapered 19 2 – – –BP-5 Protruding 19 2 – – GussetsBT-6 Tapered 19 2 – – GussetsBT-7 Tapered 19 2 – – SIA2119 AdhesiveDM-1 Tapered 12.7 2 267 12.7 Foam insertDL-2 Tapered 12.7 2 324 6.35 –DL-3 Tapered 12.7 2 324 6.35 SIA2119 AdhesiveDS-4 Tapered 12.7 1 146 6.35 Foam insertDL-5 Tapered 19 2 324 12.7 –DM-6 Tapered 19 2 267 12.7 –DL-7 Tapered 19 2 324 12.7 Foam insertBD-1 None – – – – SIA2119 Adhesive

V. Caccese et al. / Composite Structures 81 (2007) 450–462 453

ate a geometric condition where the composite at the end ofthe joint extended to the edge of the metal flange. One con-nection was tested with adhesive bonding only.

The materials used for the composite test article fabrica-tion were DOW Derakane 8084 Vinyl Ester resin and E-Glass fibers. The reinforcing fibers were 680-g 0–90� and680-g ± 45� knit fabric, manufactured by Brunswick Tech-nologies, Inc. (BTI.) The 0 and 90� fibers are stitchedtogether, as are the +45� and �45� fibers. Specimens werefabricated using a vacuum-assisted resin transfer moldingprocess (VARTM), which resulted in a 51% fiber volumecontent on the average, determined from burn-off tests. Aforty-eight layer, quasi-isotropic, [(±45f, 0/90f)6]s configu-ration was used as the basic laminate architecture. Thenominal thickness of each layer was 0.397 mm, resultingin a total laminate thickness of 19 mm. The material prop-erties for the composite laminate and the foam insert arepresented in Table 2. The steel components were madefrom a 19-mm thick, A36 grade, plate. The steel doublerplates were made from 12.7-mm and 6.35-mm thick, A36grade steel.

3. Joint testing summary

Details of the testing are presented in Kabche et al. [2,3].A general overview is presented herein for better under-standing of the finite element analyses. Tests wereperformed in displacement control mode using a 245-KN, ± 254-mm MTS� actuator that was mounted to a

Table 2Material properties

E11 (GPa) E22 = E33 (GPa) G12 = G13 (GPa)

Composite laminate: E-glass/8084 vinyl ester system, quasi-isotropic lay-up19.1 11.8 2.79

Ecomp (MPa) Etens (MPa) G (MPa)

Foam insert: DIAB Divinycell H-8085 80 31

1335-KN modular reaction frame. The end of the actuatorwas attached to the upper surface of a guided load beam.The test article was loaded transversely, in a cyclic fashion,by pushing (towards the steel support) and pulling (awayfrom the steel support) at the free ends of the compositebeams, as indicated by the double arrows in Fig. 3. Theload application point was at ±381 mm from the symmetryline of the joint. Testing was typically carried out overtwelve incrementally increasing sets of displacement cycles,where each set was comprised of three repeated hysteresisloops for a total of 36 fully reversed cycles.

Load capacities of the test articles are summarizedgraphically in Fig. 5. The damage load, Pdmg, is definedas the force at which the load-displacement hysteresis loopbecomes unstable and/or macroscopic damage is firstobserved. The ultimate load, Pult, was taken to be the max-imum force that the connection was able to withstand dur-ing the tests. Values of load at the initial damage state,Pdmg, and peak load, Pult, for both downward and upwardloading, are summarized in Table 3. For the bolted jointswithout doubler plates, the gusseted versions, specimensBP-5 and BT-6, were the stiffest and strongest. This wasparticularly true when displaced in the downward direc-tion, where the stiffer sub-structure (I-beam) had the largesteffect. When displaced in the upward direction, specimensBP-1 and BT-4 showed the peak capacities. In a generalsense, the response of standard bolted joints was tightlygrouped for the set of parameters investigated. On theother hand, the response of joints with doubler plates

G23 (GPa) m12 = m13 (GPa) m23 (GPa)

7.16 0.30 0.33

Scomp (MPa) Stens (MPa) Sshear (MPa)

1.20 2.20 1.00

0

5

10

15

20

25

30

35

40

BP-5 BT-6 BT-4 BP-3 BP-1 BT-7 BT-2 DL-2 DS-4 DL-3 DM-1 DL-7 DM-6 DL-5 BD-1Specimen

Lo

ad,

KN

Damage load - Load down

Ultimate load- Load down

Damage load - Load up

Ultimate load- Load up

Fig. 5. Summary of the damage and ultimate loads of the connections tested.

Table 3Summary of the damage load, ultimate load and failure modes

Specimen Pdmg (KN) Pult (KN) Failure modes

Load down Load up Load down Load up

BP-1 13.26 17.13 17.93 19.92 BLSC #, NSBL "BT-2 11.73 14.10 16.99 16.18 NSBL "BP-3 11.92 13.08 18.15 19.48 BLSC #, NSBL "BT-4 13.15 8.10 18.82 20.38 NSBL "BP-5 15.45 13.34 23.74 14.67 BLSC #, NSBL "BT-6 15.67 13.20 20.94 14.87 BLSC #, NSBL "BT-7 9.95 14.33 17.65 19.10 NSBL "DM-1 22.23 23.61 23.05 24.79 BLSC #, DLEC ", FFIB "DL-2 30.36 25.00 34.00 30.27 BUDE #, BLSC #, DLEC ", DUEC #, GSDE #DL-3 23.18 20.38 30.50 27.97 DLEC ", BLSC #, GSDE #DS-4 23.06 18.62 31.81 24.17 DUEC ", FFIB #, BLSC #, DLEC "DL-5 7.71 9.27 9.91 12.52 BLSC #, BUDE ", DLEC "DM-6 6.88 10.20 10.30 12.78 BLSC #, DLEC #, GSDE #DL-7 13.99 9.38 17.43 19.84 BLSC #, DLEC #, FFIB #, BUDE ", DUEC #, GSDE "BD-1 5.60 2.77 4.89 2.46 DACB "#

-40

-30

-20

-10

0

10

20

30

40

-80 -60 -40 -20 0 20 40 60 80Displacement, mm

Lo

ad

, KN

BP-1

BT-2

DS-4

DL-2

Load down Load up

Fig. 6. Load versus displacement comparison of standard and doublertype joints.

454 V. Caccese et al. / Composite Structures 81 (2007) 450–462

was much more widespread. The capacity of these jointswas demonstrated to be relatively independent of thegeometry of the doubler, but more dependent upon designfeatures such as foam inserts. The joint specimens withlong doublers, DL-2 and DL-3, and short doubler, DS-4,showed the highest capacities.

Load versus displacement envelopes are plotted usingthe peak load and displacement values for each cycle setfor specimens BP-1, BT-2, DS-4, and DL-2, as shown inFig. 6. This figure shows a comparison of the responsebetween standard bolted joints (BP-1, BT-2) and jointswith doublers (DS-4, DL-2). The doubler type connectiontypically shows an increase in capacity when compared tostandard bolted joints. Multiple failure modes wereobserved during testing and designations of the failuremodes are presented in Table 4 . The failure modesobserved are also indicated in Table 3 for each test article,where the downward and upward arrows indicate the dis-

Table 4Failure mode designations

Designation Description

BLSC Bearing at the lower surface of the composite beamBUDE Bearing at the upper surface

of composite beam under doubler edgeNSBL Net section along bolt-lineGSDE Gross section along doubler edge (BLSC + BUDE)FFIB Fracture along foam insert boundaryDUEC Delamination at the upper edge of the composite beamDLEC Delamination at the lower edge of the composite beamDACB Debonding at the adhesive and composite beam interface

V. Caccese et al. / Composite Structures 81 (2007) 450–462 455

placement direction at which the failure mode occurred.Macroscopic damage was monitored throughout thecourse of the test by using the results of the three repetitivecycles of the load versus displacement curves and by visualinspection.

3.1. Selection of joint for implementation into a full-scaleassembly

The long-term objective of this research effort is to dem-onstrate the implementation of a hybrid connection withwatertight integrity in a full-scale multi-panel assemblytest. As described by Kabche et al. [2], the DS-4 connectionconfiguration was selected for this purpose. This joint wasmade using a short cap plate with one bolt row on each sideof the connection, thereby reducing the installation labor

Cycle set 1

-4

-3

-2

-1

0

1

2

3

4

-0.60 -0.40 -0.20 0.00 0.20 0.40 0.60Displacement, mm

Load

, kN

Cycle set 8

-25-20-15-10

-505

1015202530

-5 0 5 10Displacement, mm

Load

, kN

Fig. 7. Cyclic response curves for connection type DS-4. (a) First set of disdisplacement cycles. (d) Complete set of displacement cycles.

when compared to joints with two rows of bolts. This con-nection resulted in a joint that was at least 80% of the peakcapacity of all other joints, including those with long dou-blers and two bolt rows on each side. Although more com-plex than simple bolting, the doubler-type hybrid joint wasadequate in keeping the joint closed during deformation,and the cap plate distributes the bolt load more evenly overthe composite, which will mitigate bolt load loss due to vis-coelastic creep of the composite material. The cyclicresponse of the DS-4 joint is shown in Fig. 7. The connec-tion behavior is characterized by hysteresis loops thatrepeat themselves up to the eighth set of cycles (Fig. 7c),where a load drop is observed in successive loops, indicat-ing the onset of permanent damage. A significant amountof damage occurred during the twelfth cycle set, after apeak load of 31.8 KN was achieved when pushing down,and 24.2 KN when pulling up.

4. Finite element Analysis of hybrid connections

Finite element analysis was conducted considering theimplementation of the modeling methodology into a globalmodel of the lifting body hull structure. In the global struc-ture, the mechanical response of joints is coupled to thestructural elements that they connect, and the geometry,materials and loading of the global structure will have adirect influence on the forces being transferred to the joint,and vice versa. Understanding the relationship between theresponse of both local and global systems is imperative in

Cycle set 6

-20

-15

-10

-5

0

5

10

15

20

-4 -2 0 2 4 6Displacement, mm

Load

, kN

Complete hysteresis loop

-30

-20

-10

0

10

20

30

40

-10 -5 0 5 10 15Displacement, mm

Load

, kN

Load down Load up

Loaddrop

Loaddrop

placement cycles. (b) Sixth set of displacement cycles. (c) Eighth set of

x

38.10 34.93 60.33 35.56 212.09

19.05

19.05

247.65Bolt line

A B C D E

F

381

Symmetry plane

6.35

9.525

6.35

z

Fig. 8. Geometry of the DS-4 joint (in mm) indicating shell modeldivisions.

456 V. Caccese et al. / Composite Structures 81 (2007) 450–462

the design of structures being joined with hybrid connec-tions. In general, when modeling large structures, it is notfeasible to include all of the small-scale features of thejoint. As described by Bonnani [4], the usual approachfor analyzing mechanical joints is to perform a global anal-ysis of the large-scale structure using an approximate jointstiffness, in order to determine the loads being transferredto the joint. A detailed joint configuration is then devel-oped and its stiffness can be calculated and input back intothe global model to verify design criteria. This process maybe repeated if further refinement of the joint is needed. Fora new joint design, experimental validation of the connec-tion behavior is required, as demonstrated in this work.

One simple option that can be used in developing a glo-bal model of the hybrid structure is to assume ideal condi-tions at the boundaries of the panels and sub-structure. Theideal case is such that the panel ends either have full conti-nuity or are free to rotate. The case of full continuity is asimple modeling process, as the finite element method nat-urally assumes equal displacements and rotations betweenconnected elements. The ideal pinned condition needs torelieve the equal rotation restriction along the panel ends.This can be accomplished using coincident nodes and con-straint equations.

It may be difficult to treat the true response of a hybridconnection as one of these two ideal cases since the jointfixity is such that it is partially restrained. Modeling of eachfeature of the joint may be feasible in a local study of thejoint response using 3D elements. However, this type ofmodel is currently not feasible for analysis of a large-scalestructure, such as a ship or a ship lifting body. Otheroptions for implementation of the hybrid connection prop-erties into the global structure include: (1) the use of con-nection elements between the end of the panels and thestructural backbone; and (2) modifying the properties ofthe elements making the connection. Implementation ofconnection elements requires coincident or nearly coinci-dent nodes and interconnecting elements that describe theconnection stiffness for all six degrees of freedom used ina shell element model. Modifying the properties of the ele-ments in the connection region requires that the mesh bedivided up methodically into regions that can be given dif-ferent material and section properties, which may includemanual input of the A, B, and D matrices of the shellelement.

4.1. Hybrid connection model using shell elements

The purpose of the hybrid connection shell elementmodel is to provide a simplified approach for including apartially restrained joint stiffness using shell elements. Thisanalysis focuses on a case study of the hybrid connectionwith a short doubler joint configuration (DS-4), whichwas selected for implementation into a full-scale four-panelassembly test. The experimental data for this connectionscheme is used as the baseline for validation of the model-ing techniques. All finite element models are developed

using the commercial software package ANSYS, Version10.0.

The geometry of the DS-4 sub-component joint speci-men is shown in Fig. 8. This connection scheme consistsof a composite beam sandwiched between a steel flangeand a steel doubler plate. A foam insert is used to alignthe tapered region of the composite with the edge of thesteel flange. For purposes of property designation in theanalysis, the joint has been sub-divided into six distinct sec-tions: A, B, C, D, E, and F. Section A represents the regioncomposed of the steel doubler, the composite beam and thesteel flange, from the symmetry plane to the bolt-line; sec-tion B represents the steel doubler, the composite beam andthe steel flange section, from the bolt-line to the edge of thedoubler plate; section C represents the composite beam andsteel flange in the region where a foam insert is present; sec-tion D represents the tapered composite region; section E isthe straight composite beam; and section F is the steel web.Fig. 9 shows the geometry of the top and side views of theshell model representation of this joint configuration. Thesymmetry plane is located at the center of the steel web.Element type SHELL181, which is a 4-node element withsix degrees of freedom at each node and includes the effectsof shear deformation, was chosen for this study. This ele-ment is suitable for modeling thin to moderately thick shellstructures. Based on a preliminary convergence study, amesh of 1785 elements was found appropriate. To estimateinitial stiffness, a nominal load of 4.45-KN was applied atthe free ends of the composite beams, which representsthe force applied at the location of the load heads dur-ing the experiment. Model verification was conducted atthis load magnitude, since linear response of the experi-mental load versus displacement and load versus straincurves was still observed. To simulate a fixed base condi-tion, all degrees of freedom at the bottom edge of sectionF (steel web) were constrained.

Upon prescribing element properties, the followingassumptions are made regarding the behavior of each sec-tion. In Section A, the steel flange, composite beam andthe steel doubler are assumed to act together (full continu-

Deflected shape, DS-4 (P = 4.45 KN)

-10

-8

-6

-4

-2

0

2

0 100 200 300 400

x, mm

Dis

plac

emen

t, m

m

Test EG/VESteel Hybrid

A B C D E

Fig. 10. Deflected shape of DS-4 shell element model at a 4.45 KN load.

Fig. 9. Shell model representation of the DS-4 hybrid joint.

Table 5Vertical displacement of hybrid joint model at P = 4.45 KN, load down

Position x

(mm)Experiment(mm)

Hybrid joint(mm)

Percentdifference

133 0.340 0.311 9.23203 1.407 1.359 3.41381 7.437 7.282 2.08

V. Caccese et al. / Composite Structures 81 (2007) 450–462 457

ity) due to the clamping action of the bolt. Section proper-ties are computed about the x-axis, which is assumed to belocated through the centerline of the panel acreage. In Sec-tion B, the extensional resistance is assumed to be providedonly by the composite, but not the steel doubler or steelflange, given that the section begins past the bolt-line.The flexural rigidity of this section is dependent upon theloading direction. When loaded upward, only the doublerplate and the composite resist the bending moment. Whenloaded downward, the steel flange and composite resist thebending moment. Flexural properties of each section arecomputed about each component’s own centroidal axis.The composite and steel are assumed to act independently(no continuity) where the total stiffness is the sum of theindividual parts. In Section C, it is assumed that only thecomposite region provides the extensional resistance. Whenloaded upward, the composite (with the foam insert) resiststhe bending moment, while the steel flange and compositeresist a downward acting movement. The flexural responseof each section is again assumed to be independent and theproperties are computed about each components own cent-roidal axis, as the total stiffness is the sum of the individualparts. Sections D and E are input as composite layeredshell sections, D being a sandwich and E being a laminate.Section F is input as isotropic steel.

4.2. Local shell element model results

The local shell modeling approach was validated usingthe results obtained from the DS-4 joint specimen testing.

Models with an all-steel and all-composite outer plate atregions A, B, and C were also created with fully rigid con-nectivity. These models represent the stiffest case and themost flexible case for a given joint configuration, respec-tively, and are used to determine the upper and lowerbounds of the response of the joint. Analysis at 4.45-KNresults in a maximum displacement of 7.28 mm at the freeend of the composite beam. A peak strain of 2881 micro-strain was recorded at 203 mm from the joint centerline,which lies in the straight composite beam section (E). Ver-tical displacement values (Uz) were recorded along thelength of the FE model and compared with the experimen-tal values recorded at the three LVDT locations. The strainvalues (ex) recorded at these locations correspond to straingages installed during testing. Fig. 10 shows the deflectedshape for the DS-4 joint for the three modeling cases andthe experimental results. Table 5 summarizes the discretevertical displacement values recorded and gives the percentdifference between the FEA and experimental results.Using hybrid joint properties yields reasonably accuratedisplacement predictions, within 10% difference of the testresults. Fig. 11 shows the strain profile for the DS-4 jointfor the three modeling cases and the experimental results.Overall, the local shell model in this section was found ade-quate to estimate of the joint stiffness. However, the strainresults show that this model is not feasible for providingestimates of the strains and hence the strength at the jointregion. A detailed 2D plane strain or 3D solid model,which includes local effects such as bolt pre-load and con-tact interactions, is needed to predict the strength.

Strain profile,DS-4 (P = 4.45 KN)

0

500

1000

1500

2000

2500

3000

3500

0 50 100 150 200 250x, mm

Str

ain,

mic

rost

rain

Steel EG/VE

Hybrid Test

A B C D E

Fig. 11. Longitudinal strain profile (ex) of DS-4 joint at a 4.45 KN load.

458 V. Caccese et al. / Composite Structures 81 (2007) 450–462

4.3. Nonlinear hybrid connection model

The detailed finite element model presented in this sec-tion is used to estimate the nonlinear local joint responseby modeling the contact interactions between the joint con-stituents and the effect of the bolt. A plane strain modelwas chosen to represent the connection, as implementedon a lifting body hybrid structure, since this connection isrelatively deep, resulting in near plane strain conditions

Fig. 12. Nonlinear plane strain model with contact. (a) Symm

of a section away from the corners of a panel. This modelwill not represent the response near a panel corner. Previ-ous experience with full 3D solid models of connectionshas shown that inclusion of nonlinear material response,contact and pre-load makes the models computationallyintensive and currently unfeasible for global structuralmodeling. Hence, the choice was made to use the simplerplane strain model. Model verification is conducted forthe case of the DS-4 joint configuration.

Fig. 12 presents a schematic of a plane strain model of arepresentative section of the DS-4 connection. In eachanalysis, the composite laminate consists of orthotropiclayers that are perfectly bonded together. A perfect bondis also assumed between the composite and foam insert.The orthotropic material properties and strength for a sin-gle EG/VE lamina (51% v.f.) in the principal material coor-dinate system were obtained analytically and are presentedin Table 6. Element type PLANE183 is used, which is ahigher order 2-D, 8-node element with quadratic displace-ment behavior. To save computational time, one half of thejoint was modeled with a mesh of approximately 18,500elements, using three elements through the thickness ofeach layer. All degrees of freedom at the bottom of the steelflange are constrained to simulate a fixed base conditionand symmetry boundary conditions are applied about the

etric model. (b) Exploded view showing contact surfaces.

0

2

4

6

8

10

12

14

0 100 200 300 400x, mm

Dis

plac

emen

t, m

m

Test

FE No Contact

FE Nodal ConstraintsFE Spring

-10

-8

-6

-4

-2

0

2

0 100 200 300 400

x, mm

Dis

plac

emen

t, m

m

TestFE No Contact

FE Nodal ConstraintsFE Spring

Fig. 13. Deflected shape of DS-4 plane strain model at 4.45 KN. (a) Loaddown. (b) Load up.

Table 6Orthotropic lamina properties for an EG/VE system (51% v.f.)

E11 (GPa) E22 (GPa) E33 (GPa) G12 (GPa) G13 (GPa) G23 (GPa) t12 t13 t23

38.0 10.6 10.6 3.19 3.19 2.26 0.29 0.29 0.44

V. Caccese et al. / Composite Structures 81 (2007) 450–462 459

vertical joint centerline. Nodal forces are input per unit ofdepth at the end of the joint. The steel doubler plate andsteel I-beam members are assigned isotropic propertieswith an elastic modulus of 203 GPa, a Poisson’s ratio of0.30 and yield strength of 248 MPa. The foam insert ismodeled as an isotropic material, having an elastic modu-lus of 80 MPa and a Poisson’s ratio of 0.29.

Nonlinear structural behavior is mainly sub-divided intothree types: geometric, material, and changing status non-linearity. Analyses involving contact interactions representthe class of changing status nonlinear problems. The con-tact surfaces modeling the DS-4 joint configuration aretreated as rigid-to-flexible, where the steel elements are cho-sen as the ‘‘target’’ surface and the composites as the ‘‘con-tact’’ surface. Fig. 12b shows an exploded view of the twocontact pairs (dashed lines) used to model the interactionsbetween the joint constituents, noting that the contact sur-faces will overlap geometrically in the model. TARGET169elements are used to define target contact surfaces andCONTAC172 elements are used to represent contact andsliding between 2-D target surfaces and a deformable sur-face. The TARGET169 and CONTAC172 are elementswith mid-side nodes and are hence compatible with elementPLANE183.

Several modeling approaches are used to envelop thestructural response. The initial approach is a model ofthe joint as a ‘‘glued’’ entity (no contact), where all thecomponents act together with full continuity and contactinteractions are not included. The second approachemployed contact modeling, by specifying surfaces that willinteract as the joint is loaded in different directions (nodalconstraints). The model of the bolt consists of coupling thenodal displacements of the different joint components(doubler, composite, and steel flange) along the bolt-linein the y-direction. Nodal constraints are also enforced inthe x-direction at the coincident nodes between the steeland composite interfaces, to prevent rigid body motion ofthe composite beam. Lastly, a model including the effectof the bolt is presented, where nodal constraints and aspring element along the bolt-line are implemented(spring). The spring element is attached to the upper nodeof the steel doubler and the lower node of the steel flange.In prescribing the spring element stiffness, the axial stiffnessof the bolt (AE/L) was smeared per unit depth of the con-nection. Analyses by Kabche et al. [2] show that the con-nection response is relatively insensitive to the springstiffness value in the range representing the actual boltstiffness.

The various plane strain modeling approaches are com-pared to the experimental results for a net applied end loadof 4.45 KN on a model representing the joint test specimenof 171.45 mm in depth. The displacements are comparedfor the DS-4 joint specimen in Fig. 13, which shows thedeflected shape of the joint loaded downward and upward.In a general sense, the bonded model (no contact) is thestiffest, due to the omission of the contact interfaces andthe bolt flexibility, which does not allow opening of gapsat the interfaces. The results of all FE models are within10% of the experimental displacement at x = 133 mm.The shapes begin to diverge at about x = 200 mm. Themodel with nodal constraints along the bolt-line over pre-dicts the peak displacement by about 20%. The hybrid jointmodel using contact and the spring element representingthe bolt provides a good depiction of the deflected shapeof the joint, with a difference of less than 10%.

The effect of contact pressure at the interfaces betweenthe steel and composite components was investigated using

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the spring element model for an applied load of 4.45 KN atthe free end of the beam. The experiments show that dam-age initiated at the top and bottom layers of the compositedue to the bearing action of the steel doubler and flange onthe composite beam [3]. Hence, the stresses and pressure atthese locations are of interest. Fig. 14a presents the dis-placement contour of the joint loaded downward. Asexpected, a gap opens up at the tapered interface betweenthe steel doubler and the composite beam. Another gap isobserved at the bottom left interface between the compos-ite and the steel flange. This gap occurs due to the action ofthe spring and the symmetry boundary condition appliedto the doubler plate, which allows movement of the com-posite beam in the y-direction. Fig. 14b shows the displace-ment contour of the joint loaded upward. In this case, thegap between the doubler and the composite at the taperedregion is closed. A gap opens up at the steel flange and

Fig. 14. Displacement contours (in mm) of the plane strain

composite beam interface, after the bolt-line (x > 38 mm).The larger displacement at the end of the beam is due tothe lower bending resistance provided by the steel doublerin the upward direction, when compared to the steel flangein the downward direction.

Fig. 15a shows the contact pressure distribution at theinterface between the steel doubler plate and the top ofthe composite beam. When loaded downward, a peak pres-sure of about 4.25 MPa occurs at x = 34 mm, just beforethe bolt-line, and it decreases rapidly to zero, since theregion of the composite located after the bolt-line doesnot interact with the doubler edge. For the case of loadingup, a contact pressure of about 7 MPa is observed atx = 29 mm, then a rapid drop to zero occurs. A peak pres-sure of 8.9 MPa is observed at x = 61 mm, which is wherethe edge of the steel doubler contacts the tapered region ofthe composite. Fig. 15b shows a similar phenomenon at the

model at a 4.45 KN load. (a) Load down. (b) Load up.

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Fig. 15. Contact pressure at composite/metal interfaces at a 4.45 KNload. (a) Doubler/composite interface. (b) Composite/flange interface.

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Fig. 16. Load versus outer displacement envelope for DS-4 joint.

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composite/flange interface. When compared to loading up,significantly higher pressures, exceeding 40 MPa, areobserved at this interface when loading down.

4.4. Failure prediction

In this analysis, the Tsai–Wu general quadratic interac-tion criterion is used to predict failure initiation [5]. Initia-tion of failure is used as the strength estimate for thisconnection. The plane strain model presented in this sec-tion does not account for progressive material nonlinearitydue to the decrease in stiffness with increasing load. Inorder to appropriately predict the strength of the joint, aprogressive damage model is required, where the stiffnessof the failed elements is reduced once the layers containingthese elements reach the failure envelope. Such anapproach is beyond the scope of this work, but it is recom-mended for future analyses. The plane strain model usedcorresponds to the case of a joint section which is deep intothe out-of-plane dimension (i.e., the joint in the panelassembly testing) using the DS-4 joint configuration. Thejoint is loaded in flexure, by applying a concentrated endload in increments of 4.45 KN, up to a peak value of40 KN. Both the loading downward and loading upwardcases are considered. Fig. 16 presents the resulting experi-mental and numerical load versus displacement envelopesof the joint. It is noted that the response of the system isstiffer when loading down than when loading up. For load-ing down, the response of the model closely follows theexperimental results until a load of about 8 KN, after

which slight stiffness degradation is observed in the exper-imental data. It is theorized that this decrease is due to thecyclic loading that causes minor material damage, espe-cially in the regions of high stress concentration. For load-ing up, the experimental and numerical results agree welluntil a load value of about 8 KN.

Failure initiation is estimated by observation of thestresses at key locations. Selection of these locations wasbased upon experimental observations, which showed thatdamage initiated at the key locations shown in Fig. 12a.When loading down, the points of interest are A (edge ofthe steel flange) and B (end of the tapered compositeregion). At point A, relatively high through the thicknesscompressive stresses exist at this location, which tend toinduce local compressive failure in the composite material.The analysis indicates that failure is reached at this point ata load of approximately 27.41 KN, assuming that all otherlocations are intact. At point B, a shear failure in the com-posite laminate is predicted at a load of 21.40 KN. Thisfailure is of the type that results in material degradation,but not a complete loss in load carrying capacity. For thecase of loading up, the points of interest are B and C. Fail-ure initiation at point B is predicted at a load of 15.80 KN;and a load of 17.76 KN is predicted for point C (edge ofthe doubler contacting the composite laminate.) At thisinstant, the adjacent 0-deg layer has a failure index of0.76. A prudent approach to design of this connection byfinite element analysis would use the load of 15.80 KN asthe joint capacity, which results in a reserve capacity of35 percent compared to the ultimate failure of 24.17 KN,when loading up.

5. Conclusions

An experimental investigation was conducted to assessthe relative structural performance of hybrid composite/metal bolted connections subjected to flexural loading.Two types of hybrid connections were investigated as partof this study, namely, standard bolted joints and boltedjoints with doubler plates. Overall, the goal was to developa hybrid composite/metal connection with watertight

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integrity to primarily resist bending loads. This was accom-plished by using tapered-head bolts in lieu of protruding-head bolts and by implementing a doubler plate, which actsas a cap plate and prevents the joint from opening. In addi-tion, the doubler plate decreases the bolt load loss due tocreep by spreading the bolt pre-load over a wider area, thusreducing the normal stress on the viscoelastic compositematerial. For resisting bending loads, the study showed thatjoints with doubler plates can be made stronger and rota-tionally stiffer than standard bolted joints. For example,when compared to a bolted joint counterpart with two rowsof the same diameter bolts (BT-2), a joint using a short dou-bler plate, a foam insert, and a single row of bolts (DS-4)was found to be at least 46% stronger and 29% stiffer whendisplaced downward, and 33% stronger and 34% stifferwhen displaced upward. Doubler plates proved effective inmitigating opening of the joint, which will improve the abil-ity to seal the joint and maintain watertight integrity. Whencomparing the performance of the joint configurations withdoubler plates to each other, it was observed that a jointwith a short doubler plate and one row of bolts (such asDS-4) can attain capacities comparable to those of jointswith long doubler plates and two rows of bolts (such asDL-2 and DL-3). The DS-4 joint requires roughly half ofthe steel material for fabrication of the doubler plates andhalf the number of bolts than those joints with mediumand long doubler plates. This may be a significant factorwhen considering installation and/or maintenance oflarge-scale structures that require multi-panel assembliesusing bolted connections.

Finite element analysis of the DS-4 joint configurationwas conducted using two approaches: (1) local shellelement models, used to validate connection modelingtechniques in a global analysis of the entire structure;and (2) a detailed local plane strain model, used to inves-tigate the nonlinear response of the joint. The local shellmodeling approach divided the hybrid connection intosix regions. Properties for each region were prescribedaccording to assumptions made with regards to the struc-tural details of each region. Results show that the hybridlocal shell model gives a good prediction of the connec-tion response, since the displacements were within 10%of the experimental values. Predictions of stress, strain,and strength of the connection region were not accom-

plished with this model, and required developing detailedlocal models. The detailed plane strain model includes thecontact interactions between the composite and steel con-stituents, and springs to represent the bolts. Different stiff-ness was demonstrated in pulling and pushing due to thevariation in joint gap opening caused by the connectiongeometry. The detailed model gave a good prediction ofthe DS-4 joint response. Local contact pressures wereshown to be high at the end of the flange, which is wherefailure initiated during the testing. Fabrication details toreduce this stress riser may improve the strength of theconnection.

Acknowledgements

The authors gratefully acknowledge funding for thisproject through the Office of Naval Research (ONR), un-der Grant No. N00014-01-1-0916. Dr. Roshdy G.S. Bar-soum of ONR is the cognizant program officer. Theguidance provided by Drs. Milt Critchfield, Loc Nguyen,and Gene Camponeschi of NSWC-Carderock is gratefullyacknowledged. Also, the support of Steven Loui, ToddPeltzer, Gary Shimozono, and Eric Schiff, of NavatekLtd., is greatly appreciated, along with the guidance andsupport of personnel at Applied Thermal Sciences includ-ing Steve Webber, Kendrick Light, Larry Thompson, andJosh Walls.

References

[1] Berube K, Caccese V. Construction and connection concepts for amodular advanced composite hull-form. Report No. C-9914. Depart-ment of Mechanical Engineering, University of Maine, December 16,1999.

[2] Kabche JP, Caccese V, Berube KA. Testing and analysis of hybridcomposite/metal connections and hull section for the MACH project.University of Maine, Department of Mechanical Engineering, ReportNo. UM-MACH-RPT-01-01, April, 2006. 348 p.

[3] Kabche JP, Caccese V, Berube KA, Bragg R. Experimental charac-terization of hybrid composite-to-metal bolted joints under flexuralloading. Composites-B 2007;38(1):66–78.

[4] Bonanni DL, Caiazzo A, Flanagan G. Design guide for joints inmarine composite structures—part 1, concept development. ReportNo. NSWCCD-65-TR-2000/01. NSWC Carderock, January 2001.

[5] Tsai SW, Wu EM. A general theory of strength for anisotropicmaterials. Journal of Composite Materials 1971;5:58–80.