Composite Structures 94 (2012) 1736–1744

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Composite Structures

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Static strength of a composite butt joint configuration with different attachments

Gang Li a,⇑, Ji Hua Chen b, Marko Yanishevsky a, Nicholas C. Bellinger a

a Structures and Materials Performance Laboratory, Institute for Aerospace Research, National Research Council Canada, 1200 Montreal Road, Ottawa, Ontario, Canada K1A 0R6b Aerospace Manufacturing Technology Centre, Institute for Aerospace Research, National Research Council Canada, 5145 Decelles Avenue, University of MontrealCampus, Montreal, QC, Canada H3T 2B4

a r t i c l e i n f o

Article history:Available online 24 December 2011

Keywords:Bonded-boltedComposite jointsDamageFailure modeUltimate strength

0263-8223/$ - see front matter Crown Copyright � 2doi:10.1016/j.compstruct.2011.12.008

⇑ Corresponding author. Tel.: +1 613 990 4989; faxE-mail address: Gang.Li@nrc-cnrc.gc.ca (G. Li).

a b s t r a c t

The static failure behaviour of a composite single-strap joint configuration using three different attach-ments was studied experimentally. The attachments used were: adhesive bonding; mechanical bolting;and bonded-bolted joining. The dimensions of the composite butt joint used were determined based onactual joint configurations in aircraft structures, such that the test results would be beneficial to both theacademic exploration and practical engineering application for advanced composite joint design andanalysis. A damage evolution process was presented for the bonded-bolted butt joints based on theobserved stress versus displacement curves and associated failure modes from all the related butt jointswith the three attachments. An approach was then proposed for estimating the ultimate tensile strengthin the bonded-bolted joints.

Crown Copyright � 2011 Published by Elsevier Ltd. All rights reserved.

1. Introduction

Mechanical fastening and adhesive bonding are typically usedto join various load-carrying composite elements [1,2]. Bondedjoints are more attractive as compared to bolted joints [3]. How-ever, the bondline itself is still a potential source of weakness, cre-ating uncertainty in adhesive bond performance [4]. Laminates aretypically poor at resisting loads in the thickness direction andinterlaminar shearing. In addition, through-thickness loads andinterlaminar shearing can greatly reduce the bonded joint strength[5–8]. Fasteners can effectively arrest the damage spread in abonded section and improve the joint static strength and fatigueperformance [3,9–11]. However, single-row bolt configuration forbonded-bolted joints [3,10,11] are not used in the actual aircraftstructures, since this poses a safety issue.

Currently, information is still lacking on the strength and failurebehaviour of specific joint configurations, such as the single-strap(or single-doubler) joint configuration. There is little strength dataavailable in the open literature on the butt joint configurationsaimed at actual aircraft structures. High adhesive peel stress atthe inner bonded overlap end and slow recognition of potentialapplications could be sources of the research gaps [7,12,13]. Thisbutt joint configuration could ensure an excellent aerodynamicallysmooth surface exposed to the air stream. The single-strap jointconfiguration could also act as a baseline to develop other relevant

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joints for aircraft composite fuselage structures, with a bonded-bolted or co-cured-bolted attachment approach being used toensure joint structural integrity.

In this paper, the tensile mechanical failure behaviour of single-strap joints with different attachments was studied experimen-tally. Current joint dimensions ensure that two-row bolts with a25.4 mm spacing and proper edges can be installed in each overlapsection. Usually three-row riveted lap joints are used in the fuse-lage structures and some two-row bolt arrangements are installedat specific locations in other aircraft structures. Thus, the test re-sults of the current study can be beneficial to both the academicexploration and practical aircraft engineering application. Atten-tion was focused on the joint strength and associated failure modeinduced by the attachment method and laminate surface ply orien-tation. The results combined with theoretical analysis help tounderstand the damage evolution and the strength enhancementmechanism in the bonded-bolted joints.

2. Experimental testing

2.1. Material and test coupon preparation

To ensure the quality and repeatability of the compositecoupons, an automated fibre placement (AFP) machine wasemployed to fabricate the composite laminates. The material prop-erties for the 0.14 mm thick lamina of CYCOM 5276-1 T40-800 car-bon fibre slit tapes manufactured by Cytec Engineered Materialswere: E11 = 145 GPa, E22 = 8.9 GPa, m12 = 0.31, and G12 = 4.5 GPa.

rights reserved.

G. Li et al. / Composite Structures 94 (2012) 1736–1744 1737

In order to investigate the influence of fibre alignment on thebond line and joint performance, two stacking sequences of S1,[45/�45/0/90]2s, and S2, [0/90/45/�45]2s, were chosen. The curecycle used was based on the supplied lamina technical datasheet.The nominal thickness of the cured 16-ply laminate was2.24 mm.

The details of the joint coupon fabrication are documented else-where [14]. Prior to bonding, the necessary surface treatmentswere conducted. The joint coupon consisted of identical laminatesfor the adherends and doubler, as shown in Fig. 1. The adhesive fil-let profile was naturally formed during the cure cycle without fur-ther machining and its length on the adherend was approximately3.2 mm. Tapered end tabs were made from 3.18 mm thick FR4Fibreglass sheet. A total of 12 joints were tested. They were bondedS1 and S2 joints, bolted S1 joints, and bonded-bolted S1 joints. Dueto the limitation of the available laminates, there were no S2 jointsin the bolted and bonded-bolted attachment conditions for furthertests. According to the preliminary tensile tests for the laminatecoupons, the strength scatter obtained was small with a standarddeviation less than 3.8% of the average strength. Therefore, onlythree joints in each condition were tested. The film adhesive,FM300K from Cytec Industries, was used for joint bonding andits thickness was controlled to be 0.17 mm. An elastic modulusof Ea = 3 GPa and a Poisson’s ratio of va = 0.32 for the adhesiveparameters were assumed.

For bolted joints, countersunk holes with 25.4 mm spacing atthe joint central axis position were drilled. An ultrasonic pulseecho method was used for detecting potential damage induced;there was no damage detected in the laminates using the

(a)

(b)

(c)

12.725.412.7 12

125.4 12.712.7

Flush head side

7.5

H

101.6

Width: 25.4

30045

2554.2 3.2

74.2

Fig. 1. Schematic configuration for the composite single-strap joints using three attachmin mm).

selected drilling parameters after preliminary trials. The edgewas 12.7 mm from the hole centre to its nearby overlap end.The average hole diameter of the protruding part was approxi-mately 4.19 mm based on 30 optical measurements of 15 holes.After completion of the protruding holes, hole countersinkingwas carried out to a 0.75 mm depth in the adherends. The holediameter was approximately 7.51 mm on the laminate counter-sunk head surface. Blind titanium bolts with a 130� countersunkhead, MBF2113S-5-200 Composi-Lok� II fasteners, were installedusing a pneumatic pistol. The bolt shank diameter was 4.15 mm,and the grip range was 3.81 to 5.08 mm. These fasteners are com-pletely compatible with graphite or carbon fibres without causinggalvanic corrosion. The deformed large foot prints at the butt sidecan avoid introducing damage to the composite during the fas-tener installation stage.

For bonded-bolted joints, holes were drilled after joint bonding.As shown in Fig. 1c, the cap diameter was approximately 7.5 mmon the joint outer surface. The measured average diameter, D, ofthe large foot print on the doubler surface was approximately7.24 mm based on 32 optical measurements of 16 bolts. The aver-age bolt protruding height, H, was approximately 5.58 mm basedon 16 measurements on 16 bolts.

2.2. Testing condition

An MTS servo-hydraulic load frame equipped with a 220 kNload cell was used for tensile testing. All the coupons were loadedto failure at a speed of 1 mm/min. Cross-head displacement wasused to plot the stress versus displacement curves.

25.4.7 12.7

25.42.7 12.7

Fastener large foot print side

D

Doubler Adherend

54.2

ent methods: (a) bonded, (b) bolted, and (c) bonded-bolted (not to scale, dimensions

1738 G. Li et al. / Composite Structures 94 (2012) 1736–1744

3. Experimental results and discussion

3.1. Stress versus displacement curves and the associated failuremodes

3.1.1. Bonded single-strap joints3.1.1.1. Analysis of experimental results. The bonded joint stressversus displacement curves are plotted in Fig. 2. A quasi-linear var-iation behaviour with small displacement was obtained. Identicalstiffness was exhibited in the two bonded joints. Before the cata-strophic failure, two minor stress discontinuities occurred in theS1 joints and only one in the S2 joints. The first discontinuity positionin Fig. 2a was at the exact same position for the only discontinuity inFig. 2b. The reason for the stress discontinuity was explored whileassessing the observed failure mode subsequently in this section.

The average joint tensile strength tested was 246.2 MPa with a3.2 MPa standard deviation for the bonded S1 joints, and343.3 MPa with a 20.3 MPa standard deviation for the S2 joints.The corresponding tensile displacement was 1.2 mm with a0.1 mm standard deviation for the S1 joints, and 1.5 mm with a0.1 mm standard deviation for the S2 joints. The joint tensilestrength was the ultimate applied tensile stress in the joint adher-end. The S2 joint results were approximately 39% higher than theS1 stacked joints due to the ply orientation at the bonding surface,as well as the associated bending stiffness. This measured strengthranking was consistent with the theoretical analysis in the latter ofthis section conducted based on the adhesive stress profile. Com-pared to the previously obtained laminate tensile strength, thetwo bonded joints only achieved approximately 28% and 40% ofthe strengths of their adherends.

(a)

(b)

0

50

100

150

200

250

300

App

lied

tens

ile s

tres

s (M

Pa)

Cross-head displacement (mm)

S1-J1S1-J2S1-J3

1st discontinuity

2nd discontinuity

0

50

100

150

200

250

300

350

400

0 0.25 0.5 0.75 1 1.25 1.5 1.75

0 0.25 0.5 0.75 1 1.25 1.5 1.75

App

lied

tens

ile s

tres

s (M

Pa)

Cross-head displacement (mm)

S2-J1

S2-J2

S2-J3

The only discontinuity

Fig. 2. Variation in the stress versus displacement curves obtained from tensiletests on the adhesively bonded single-strap joints with the: (a) S1 laminate and (b)S2 laminate.

If using the adhesive shear strength for comparison, as in refer-ence [15], the associated average shear strength, joint load over thebonding area, was 10.8 MPa for the S1 joint and 15.1 MPa for the S2joint, respectively. Due to data paucity, strength comparison withother researchers’ data of the butt joints could not be conducted.Since the single-strap joint actually consists of two end-to-end sin-gle-lap joints [12–14], an average 11.4 MPa shear strength wastested in a single-lap joint configuration with the same bondingarea 50.8 � 25.4 mm2 and similar laminates [15]. Equivalentstrength was illustrated between the current bonded single-strapjoints and those single-lap joints.

Completely different failure modes can be found from the frac-ture surfaces of the typical failed bonded joints, as shown in Fig. 3.In either case, no fibre breakage was found. For the S1 joints, delam-ination seemed to be present only inside the doubler laminate; minoradhesive failure was found at the inner overlap end region. The del-aminated plies from the doubler were still bonded on the adherendsurfaces. Typical failure modes were: (i) extensive delamination be-tween the 0� and either +45� or �45� plies; (ii) some delaminationbetween the ±45� plies; and (iii) minor adhesive failure at the adhe-sive–doubler interface in the doubler centre area. The observationssuggested that: (i) the S1 joint strength was dominated by laminateinterlaminar strength; and (ii) the discontinuity of the bonding sur-face ply ±45� fibres at the free edge adversely affected the interlam-inar condition (they were too weak to resist the delamination). Forthe bonded joints using S2 laminates, the only adhesive failurewas found at the: (i) adhesive–doubler interface located at the dou-bler central area; and (ii) adhesive–adherend interface in the outeroverlap area. The S2 joint strength was completely determined bythe bondline condition for the laminate with the 0� surface ply atthe bonding surface. The 0� orientation surface ply could effectivelyavoid the delamination.

The following relationships could be discerned: (i) the adhesivefailure at the inner overlap edge region between adherend-to-dou-bler and adherend-to-adherend ends was likely the source of thefirst discontinuity in all curves in Fig. 2; (ii) the subsequent delam-ination, combined with other failure modes was the source of thesecond discontinuity in Fig. 2a; and (iii) the final joint separationwas caused by extensive failure, delamination or adhesive debond-ing, a totally different failure situation.

(a)

(b)

Doubler

Adherend Adherend

Plies delaminated from doubler

Plies delaminated from doubler

Adherend

Doubler Adherend

Fig. 3. The failure mode patterns observed on a: (a) bonded S1 joint and (b) bondedS2 joint.

G. Li et al. / Composite Structures 94 (2012) 1736–1744 1739

3.1.1.2. Theoretical analysis of the strength ranking of the testedbonded joints. To understand the tested joint strength ranking, thefollowing theoretical explanation is provided. Since the adhesivestress magnitude and profile are directly related to the bondedjoint strength, a theoretical analysis of the adhesive stresses shouldprovide guidance on this issue. The single-strap butt joint is actu-ally fabricated by attaching two single-lap joints end-to-end. Thesame adhesive stress equilibrium equations exist between the sin-gle-strap and single-lap joints [13,16–19]. The elastic analysis ofthe bonded joints can be first traced back to the 1930s by Volker-son using a shear lag model [20]. To date, elastic closed-form adhe-sive stress solution of the balanced single-lap joint with secondarybending has been well established [21–27]. Recently, completeclosed-form adhesive stress solutions in the bonded butt joint withdifferent adherend and doubler laminates have been obtained bythe author [13,18,19]. The joint analytical advancements can beused to study the bonded one-sided patch repair applications forcracked or local damaged structures in aged aircrafts [28,29]. Forthe current bonded butt joints made with identical adherend anddoubler laminates, the coupling relationship between the adhesivepeel and shear stresses vanishes. The associated stress equationsunder the cylindrical bent laminate condition in the elastic rangecan be simplified as:

d3sa

dx3 þ a1dsa

dx¼ 0 ð1aÞ

d4ra

dx4 þ b1ra ¼ 0 ð1bÞ

where sa is the adhesive shear stress, ra is the adhesive peel stress,and the two parameters, a1 and b1 are determined by the laminatelayup condition [13] as:

a1 ¼ �Ga

g2

A11þ tðt þ gÞ

2D11

� �ð2aÞ

b1 ¼2Ea

gD11ð2bÞ

A11 and D11 are elements in the stiffness (ABD) matrix of the unit-width laminate constitutive equation; Ea: adhesive Young’s modu-lus; Ga: adhesive shear modulus; g: adhesive thickness; and t:thickness for the identical adherend and doubler laminates. Thegeneral solutions for the adhesive stresses are:

sa ¼ C0S þ C1S coshðxffiffiffiffiffiffiffiffiffi�a1p

Þ þ C2S sinhðxffiffiffiffiffiffiffiffiffi�a1p

Þ ð3aÞ

ra¼C3S coshx

ffiffiffiffiffib1

44

r !cosx

ffiffiffiffiffib1

44

r !þC4S sinhx

ffiffiffiffiffib1

44

r !cosx

ffiffiffiffiffib1

44

r !

þC5S coshx

ffiffiffiffiffib1

44

r !sinx

ffiffiffiffiffib1

44

r !

þC6S sinhx

ffiffiffiffiffib1

44

r !sinx

ffiffiffiffiffib1

44

r !ð3bÞ

where the corresponding boundary conditions and the determinedseven constants CiS (i = 0 to 6) in the above expressions are repre-sented in the Appendix from Refs. [13,19].

At the inner bonded overlap edge, the adhesive stresses willreach their peak levels as:

ðsaÞmax ¼ sajx¼�c ð3cÞðraÞmax ¼ ðraÞjx¼�c ð3dÞ

where the original point x = 0 of the coordinate system xow, Fig. 4,for the adhesive stress analysis is located at the centre of eachbonded overlap section (using left one for example), and the lengthof each bonded overlap is 2c.

As shown in Fig. 4, a simplified bonded joint configuration with-out the adhesive fillets was used for theoretical analysis. The appliedloads are: V0, and M0 = unit-width shear force and bending momentat the outer bonded overlap edge; V1, and M1 = unit-width shearforce and bending moment at the inner bonded overlap edge. Theseedge loads can be determined under the given joint remote tensileload, T, as explained elsewhere [13]. The unit-width tensile load, T,is applicable for both edges. Then, according to the laminate layupcondition and material properties, the two coefficients, a1 and b1

can be calculated. Finally, the adhesive stresses can be determined.The dimensions of the simplified joint used for theoretical anal-

ysis were that: each outer unbonded adhernend was set to be54 mm long (excluding the tensile loading sections), the unbondedinner doubler section was set to be 0.5 mm long, and the remainingdimensions were the same as in tested bonded joint. For instance,when the joint was loaded at a remote tensile stress of 80 MPa, thecorresponding adhesive stress profiles are presented in Fig. 5. It canbe seen from this figure that both the adhesive peel and shearstresses were highly localized near the bonded overlap edge areas,with much higher stress magnitudes at the inner bonded overlapedge. It is evident that the laminate layup sequence does have cer-tain influence on the peak stress magnitude. For the 16-ply layupS1 and S2 laminates, the effective Young’s moduli was identical,while the bending stiffness of the S1 laminates was approximately80% of the S2. As shown in Fig. 5a, the peak peel stress in the S1joint was approximately 11% at the inner overlap edge and 8% atthe outer overlap edge, higher than that of S2 the joint. Fig. 5bshows that the peak shear stress in the S1 joint was approximately10% higher than the S2 joint at the inner bonded overlap edge. Peakshear stresses were nearly the same at the outer overlap edge forboth the S1 and S2 joints. The test results were verified by theadhesive stress profiles in two aspects: (i) the bonded S1 joint suf-fered higher stresses and its strength should be lower than the S2joint; and (ii) the damage evolution occurred first at the innerbonded overlap edge area and then propagated towards the outeredge direction. It is proven that this theoretical analysis is a prac-tical approach to predict the strength trend of the bonded buttjoints under the influence of laminate layup condition.

3.1.2. Bolted single-strap jointsFig. 6 shows the stress versus displacement curve and the asso-

ciated failure mode. As shown in Fig. 6a, the average maximum dis-placement was approximately 8.1 mm. The curves displayedhighly non-linear behaviour with a progressive failure process. Arelatively large stress discontinuity phenomenon occurred at theultimate load position. However, the bolted joint was able to con-tinue sustaining load and underwent extra displacement. Compar-ing the bonded joint information presented in Fig. 2, the stiffness inthe bolted joint was less than that of the bonded joint. The testedjoint average tensile strength was 164.3 MPa (with a 5.7 MPa stan-dard deviation), accounting for 67% of the strength of the corre-sponding bonded joints. The lower strength of a single-rowbolted joints versus the bonded joints was also reported in [3]for a double-lap joint configuration.

Fig. 6b shows the failure mode presented on a typical failedbolted joint. The identified failure modes were: bearing, bolt rota-tion, sharp edge of the countersunk cap cutting into the adherendlaminate, bolt countersunk cap shear-off, bolt pull-through, andlocalized delamination at the big foot edge (due to the bolt rota-tion) on the doubler free inner surface. Bearing failure was foundat the fastener hole edge area due to the high stress concentration.The rotated bolts applied an additional compressive load to thejoint laminate through-thickness direction. As the joint tensile loadwas increased, causing further secondary bending and bolt rota-tion, this damage development led to the occurrence of other fail-ure modes, such as the cap sharp edge cutting into the laminate,

Inner section

1V TV0M0

adherend2c c

w

x adhesive

doubler

T 1M

c

t t o

Fig. 4. Schematic diagram of the loads at the (left) bonded overlap edges and the coordinate frame for the adhesive stress analysis.

(a) Adhesive peel stress profile as well as the peak stress at the inner overlap edge region

(b) Adhesive shear stress profile as well as the peak stress at the inner overlap edge region

-40

-20

0

20

40

60

80

100

120

140

Adh

esiv

e pe

el s

tres

s (M

Pa)

Overlap position (mm)

S1 joint

S2 joint

0

20

40

60

80

100

120

140

Adh

esiv

e pe

el s

tres

s (M

Pa)

Overlap position (mm)

S1 joint

S2 joint

Left overlap section

Right overlap section Right inner overlap edge

-100

-80

-60

-40

-20

0

20

40

60

80

100

Adh

esiv

e sh

ear s

tres

s (M

Pa)

Overlap position (mm)

S1 joint

S2 joint

Left overlap section

Right overlap section

0

20

40

60

80

100

-60 -40 -20 0 20 40 60

0 0.5 1 1.5 2

-60 -40 -20 0 20 40 60

0 1 2 3 4 5

Adh

esiv

e sh

ear s

tres

s (M

Pa)

Overlap position (mm)

S1 joint

S2 joint

Right inner overlap edge

Fig. 5. Profiles for the adhesive: (a) peel and (b) shear stresses in the bonded composite butt S1 [45/�45/0/90]2s and S2 [0/90/45/�45]2s joints at a remote tensile stress of80 MPa.

1740 G. Li et al. / Composite Structures 94 (2012) 1736–1744

countersunk head shear-off, bolt pull-through, delamination at thebolt big foot edge, etc. The development of these failure modes wasreflected by discontinuities in the stress versus displacementcurves. The large tensile displacement could be the consequenceof the significant bolt rigid rotation and pull-through.

3.1.3. Bonded-bolted single-strap jointsThe stress variation curve and the associated failure for the

bonded-bolted joint are shown in Fig. 7. A much higher tensilestrength than either the bonded or bolted joints was exhibited inthe bonded-bolted joints, as shown in Fig. 7a. Following severalsmall stress discontinuities, a significant strength drop was ob-served at the ultimate stress level. Large displacement and pro-gressive failure behaviour were exhibited. Two distinct stageswere displayed: nearly linear variation was observed up to the ulti-mate tensile stress level; followed by highly non-linear behaviour

in the second stage until failure. This nonlinear stage accounted foralmost two thirds of the displacement range. Table 1 summariesthe test data. The joint tensile strength averaged 373.5 MPa,approximately 52% higher than the corresponding bonded jointsand 127% higher than the bolted joints. At the ultimate stress level,the corresponding displacement was approximately 50% higherthan for the corresponding bonded joints. Comparing the jointstress curves in Figs. 2a and 7a, identical initial stiffness was foundfor the bonded and bonded-bolted joints. Even though the average‘‘significant tensile strength drop’’ was up to 259.3 MPa, the jointresidual tensile strength, 186.1 MPa, was still higher than that ofthe corresponding bolted joint. The average overall displacement,5.7 mm, was greater than that of the bonded joints.

Similar strength change trends were also reported in references[10,11] in a single-lap joint configuration by comparing thebonded-bolted (with only one bolt) and bonded methods. Despite

(a)

(b)

0

20

40

60

80

100

120

140

160

180

0 2 4 6 8 10

Tens

ile s

tres

s (M

Pa)

Elongation (mm) of bolted joint

S1-BltJ2

S1-BltJ3

S1-BltJ4

Doubler

Adherend

Shear-off cap bolt

Adherend

Fig. 6. Test information of the bolted single-strap joints with the S1 laminate for:(a) variation in the stress versus displacement curves and (b) the associated failuremode pattern.

(a)

(b)

0

100

200

300

400

0 2 4 6 8

Tens

ile s

tres

s (M

Pa)

Joint elongation (mm)

S1-BBJ1

S1-BBJ2

S1-BBJ3Significant strength drop

Bolts

Adherend

Doubler

Adherend

Inner overlap end

Delaminated surface of adherend

Fig. 7. Test information of the bonded-bolted single-strap joints with the S1laminate for: (a) variation in the stress versus displacement curves and (b) theassociated failure mode.

G. Li et al. / Composite Structures 94 (2012) 1736–1744 1741

efforts, a practical quantitative analysis of the increment mecha-nism correlated with fastener deformation in the bonded-boltedattachment could not be established.

A photo of a typical failed bonded-bolted joint is shown inFig. 7b. Again, the rigid bolt rotation was observed as in the boltedjoints. In addition to the typical failure modes in the bolted joints,two extra failure modes were observed: (i) extensive delaminationin the adherend laminate (minor in doubler) near the bonding sur-face; and (ii) adhesive failure at the adhesive–adherend interface ina small area. Comparing the test information presented in Figs. 2a,3a, 6, and 7, the correlations between the curve discontinuities andfailure modes in the S1 bonded-bolted joint could be: (i) the initialsmall stress discontinuities in Fig. 7a were caused by early failureof the adhesive bond and delamination at the inner overlap end re-gion; (ii) the significant joint strength drop was caused by theextensive delamination combined with bolt rotation and hole bear-ing damage; and (iii) the second stage was similar to that of abolted joint under the influence of the bolts and the interlaminarconditions. The bonded-bolted joint failure occurred when thetwo countersunk caps were sheared off, followed by the immediatepulling out of the bolt from one adherend.

The highest strength in the joints could be attributed to the estab-lishment of an integrated section between the two bolts of thebonded-bolted attachment. This section has a self-locking andstrengthening capability from a sort of ‘‘feedback loop’’ betweenthe adhesive bonding and the two bolts installed to this overlap sec-tion. The peel stress at the crack tip decreased as a result of the intro-duced bolt clamping compressive load, while at the same time theadhesive bond increased the overlap bending stiffness and de-creased the relative shear displacement between the adherendand doubler, which improved the bolt capability to resist pull-through. Shearing deformation in adhesive layer was also reduced

due to the bonded-bolted attachment. Before the occurrence ofextensive delamination as well as the bolt rotation, the stiffness ofthe overlap section was very high and behaved like a single laminate.

3.2. Damage evolution in the bonded-bolted joints

For the bonded-bolted joint, the fasteners changed the damageevolution previously experienced in the bonded-only joint. A dam-age evolution sequence was proposed and presented in Fig. 8. Itwas observed that: (a) adhesive debond first developed at the inneroverlap edge between the adherend-to-adherend end position. Thendelamination damage in both adherend and doubler developed atthe inner bonded overlap end region. A single major delaminationthen grew and stopped at the inner bolt position; (b) new adhesiveand/or adherend delamination damage was created at the two outeroverlap end areas, followed by the adherend delamination growinguntil the nearby outer bolt position and the delamination in the dou-bler grew simultaneously; (c) delaminations in the adherend, com-bined with bolt rotation and other minor failure modes thenfollowed and the adherend delamination grew extensively in theoverlap section between the two bolts, leading to a significant lossof joint strength due to the linkage of the damage sites inside andoutside of the bolted overlap section; and (d) localized damage inthe hole bore and bolt developing rapidly until complete separationoccurred between the adherend and doubler, leading to the caps ofthe two bolts in one overlap section shearing off. The most dangerousdamage was found to have been created in stage (c).

3.3. An approach for estimating the ultimate strength in the bonded-bolted joint

Since the joint tensile strength consists of two contributions, onefrom the adhesive and the other from the installed bolts, the

Table 1Tensile test data for the bonded-bolted composite joints.

Joint coupon S1-BBJ1 S1-BBJ2 S1-BBJ3 Average Std. dev.

Tensile strength (MPa) 369.8 365.6 385 373.5 10.2Significant strength drop (MPa)a 271.5 263.8 242.7 259.3 14.9Residual tensile strength in the second stage (MPa)b 189.6 193.3 175.3 186.1 9.5Cross-head displacement (mm) at ultimate load 1.9 1.7 1.9 1.8 0.1Cross-head displacement (mm) at final failure 5.4 5.9 5.7 5.7 0.3

a The magnitude of the joint tensile strength drop at the ultimate stress position.b The maximum tensile strength in the second stage.

(a) Adhesive failure and delaminations in adherend and doubler in the inner overlap region

(b) New damages in adhesive and adherend at the outer overlap end regions

(c) Adherend delamination in one bolted overlap section with evident bolt rotation

(d) One adherend separated from the joint

Outer overlap end Adhesive failure

Bolted overlap section

Cap shear-off Separated adherend with delamination

Delaminated plies from adherend

Delamination

Inner overlap end

Fig. 8. Schematic representation showing the proposed damage evolution sequence in S1 bonded-bolted joint.

1742 G. Li et al. / Composite Structures 94 (2012) 1736–1744

ultimate tensile strength induced by the bonded-bolted attachmentcould be estimated based on the test data from the bonded-only andbolted-only joints, with the aid of theoretical stress/strain analysisdescribed in Section 3.1.1. Since the bonded-only joint has muchless elongation than that of the bolted-only joint, it was reasonablyconcluded that the adhesive bond reached its peak capability whenthe bonded-bolted joint reached the ultimate strength. The strengthcomponent from the adhesive might be approximated based on thetested strength of the bonded-only joint. Thus, the ultimate jointtensile strength could be expressed as:

ðrbbÞult ¼Abonded � 2Ahole

AbondedðrbondedÞult þ rbolted ð4Þ

where (rbb)ult is the ultimate tensile strength of bonded-boltedjoint; (rbonded)ult is the ultimate tensile strength of bonded-onlyjoint; Abonded = 25.4 � 25.4 mm2 is the bonded area in one overlapsection in bonded-only joint; Aholde = pr2, area of a hole (r � 2.1 mm)drilling-out for bolt installation; rbolted, is the strength or joint ten-sile stress contributed from bolt, which would be the enhancementin bonded-bolted joint strength provided Abonded�2Ahole

Abonded� 1.

The strength component contribution from the bolt could beestimated based on the correlation of the bolt rotation angle withthe tensile stress in the bolted-only joint. The bolt rotation anglecould be obtained based on the adhesive shear strain at the innerbolt position. Under the ultimate strength of the bonded-onlyjoint condition, the adhesive shear strain at the inner bolt posi-tion can be obtained from the presented theoretical solutionand the work in references [13,18] by assuming an adhesivedelamination or crack generated between the two inner bolts.For the current condition, the calculated adhesive shear stressat the inner bolt position was approximately 270 MPa, and theassociated shear strain was 0.238 radians, i.e. 13.6�. This angleshould be the same as the bolt rotation angle. In the theoreticalanalysis, linear elastic behaviour was assumed. In reality, theshear stress might not be that high, and the corresponding shearstrain could be high, since the adhesive was behaving nonlinearlyin an elasto-plastic sense. For the sake of simplicity and approx-imation, the linear analysis for estimating the shear strain at highstress level could be acceptable for the preliminary engineeringanalysis and application.

0

20

40

60

80

100

120

140

160

180

0 10 20 30 40 50 60 70

Tens

ile s

tres

s (M

Pa)

Bolt rotation (degree)

S1_BltJ1

S1_BltJ2

S1_BltJ3

Fig. 9. Correlation of the bolt rotation with the applied remote tensile stress in thebolted-only S1 joints.

G. Li et al. / Composite Structures 94 (2012) 1736–1744 1743

The rotation-stress correlation could be obtained from thestress versus elongation curve of the bolted-only joint. The totaltensile elongation in bolted joint could be expressed as:

Dltotal ¼ 2Dl1 þ Dl2 þ 2Dl3T þ 2Dl3R ð5aÞ

where Dl1 ¼ rL1E , the elongation of an outer adherend with the length

of L1 = 74.2 + 12.7 = 86.9 mm; Dl2 ¼ rL2E , the elongation of the inner

unbolted doubler with the length of L2 = 25.4 mm; Dl3T ¼ rL32E , the

elongation of each L3 = 25.4 mm long overlap section withinthe two bolts; E = the laminate effective Young’s modulus; andr = the applied joint remote tensile stress.

The relative displacement between the mid-planes of theadherend and doubler in each bolted overlap section caused bythe bolt rotation can be approximated as:

Dl3R ¼ ðDltotal � 2Dl1 � Dl2 � 2Dl3TÞ=2 ð5bÞ

The bolt rotation angle b in radians could be determined as:

b ¼ artanDl3R

2ðt1 þ t2Þ=2

� �¼ artan

Dl3R2t

� �ð5cÞ

where t1 and t2 are the adherend and doubler thicknesses. For thecurrent joints, they were the same and expressed as t.

The variation in the joint tensile stress versus the bolt rotationcurve can be determined and is shown in Fig. 9. The evident dis-continuities and drops in the stress variation were likely causedby the considerable damage in the laminate hole bearing andbolt cap tearing. The bolted joint tensile stress was approxi-mately 138 MP at b = 13.6�. Using Eq. (4), the estimated averageultimate tensile stress of the bonded-bolted S1 joint was373.7 MPa, while the tested average strength was 373.5 MPa.Very good agreement was achieved. However, applicability ofthe conjecture needs further validation through more test datafor this kind of butt joint configuration, which will be carriedout in the near future.

4. Conclusions

The following conclusions have been made:

(i) The bonded single-strap joint could have equivalent staticstrength to a single-lap joint. At the coupon level, the sur-face 0� ply orientation in the S2 joints avoided delamina-tion in laminates. The higher tensile strength was testedfrom the bonded S2 joints than the S1 joints, which canbe theoretically explained based on the adhesive stressprofile. Progressive failure behaviour was exhibited by bothbolted and bonded-bolted joints.

(ii) The bonded-bolted joints exhibited an increased strength ofapproximately 52% with respect to the correspondingbonded joints and 127% higher than the bolted joints. Spe-cific failure modes were identified for the qualitative inves-tigation of the joint strength under different attachment andlaminate; and

(iii) The obtained results suggest that an integrated sectionbetween the two bolts installed in that overlap section wasachieved by using the bonded-bolted attachment method,which led to a significant increase in the joint strength. Adamage evolution process and an approach for estimatingthe ultimate joint strength were proposed for the bonded-bolted joints, which would help the future work for jointstrength analysis and design using the bonded-boltedattachments.

Acknowledgements

The full support received from IAR is acknowledged and greatlyappreciated. A special appreciation is extended to Dr. AndrewJohnston for his valuable discussions and suggestions in the re-search. Many thanks to the staff who helped perform the experi-mental work in this program.

Appendix A. Expressions for relevant constants in adhesivestresses

Detailed explanations can be found in references [13,19], andonly basic information is being presented here. The three boundaryconditions for determining the three constants, CiS (i = 0 to 2) in theadhesive shear stress sa, are:Z c

�csadx ¼ �T;

dsa

dx

����x¼�c

¼ Ga

gT

A11þ t

2M0

D11

� �;dsa

dx

����x¼c

¼ Ga

g� T

A11þ t

2M1

D11

� �

where A11 and D11 are elements in the stiffness (ABD) matrix of theunit-width laminate constitutive equation; Ga: the adhesive shearmodulus; and g: the joint adhesive thickness. Referring to Fig. 4,2C is the length for both bonded overlap sections; t: the thicknessfor the identical adherend and doubler laminates; T: the remoteunit-width tensile force; M0 and M1: unit-width bending momentsat the outer and inner bonded overlap ends, respectively. The firstboundary condition is the equilibrium relationship in the adherendbetween the applied tensile load and the integral of the resultingshear stress in the adhesive layer. The second and third boundaryconditions relate the first derivative of the adhesive shear stressto the loads at two overlap edges.

The obtained constants are:

C0S ¼1c� T

2þ

TA11þ t

4M0�M1

D11

2A11þ t2

2D11

!C1S ¼

Gag � T

A11þ t

4M1�M0

D11

� �ffiffiffiffiffiffiffiffiffi�a1p

sinhðc ffiffiffiffiffiffiffiffiffi�a1p Þ C2S

¼Gag

t4

M0þM1D11ffiffiffiffiffiffiffiffiffi�a1

pcoshðc ffiffiffiffiffiffiffiffiffi�a1

p Þ

The boundary conditions for determining the four constants, CiS

(i = 3 to 6) in the adhesive peel stress ra are:

d2ra

dx2

�����x¼�c

¼ Ea

gM0

D11;

d2ra

dx2

�����x¼c

¼ � Ea

gM1

D11;

d3ra

dx3

�����x¼�c

¼ Ea

gV0

D11;

d3ra

dx3

�����x¼c

¼ � Ea

gV1

D11:

1744 G. Li et al. / Composite Structures 94 (2012) 1736–1744

The above four boundary conditions relate the derivatives of theadhesive peel stress with the applied loads at the outer and inneroverlap edges of the balanced butt joint.

The obtained constants are:

C3S¼

Ea2gD11

coshcffiffiffiffib14

4q� �

coscffiffiffiffib14

4q� �

V0þV1ffiffiffiffiffiffiffiffiffiffib14

� 34q þM0�M1ffiffiffiffi

b14

p tanhcffiffiffiffib14

4q� �

�tancffiffiffiffib14

4q� �� �0

@1A

sinh2cffiffiffiffib14

4q� �

þsin2cffiffiffiffib14

4q� �

C4S¼

Ea2gD11

sinhcffiffiffiffib14

4q� �

coscffiffiffiffib14

4q� �

V0�V1ffiffiffiffiffiffiffiffiffiffib14

� 34q þM0þM1ffiffiffiffi

b14

p 1

tanhcðffiffiffiffib14

4p

Þ� tanc

ffiffiffiffib14

4q� � !0

@1A

sin2cffiffiffiffib14

4q� �

�sinh2cffiffiffiffib14

4q� �

C5S¼

Ea2gD11

coshcffiffiffiffib14

4q� �

sincffiffiffiffib14

4q� �

V0�V1ffiffiffiffiffiffiffiffiffiffib14

� 34q þM0þM1ffiffiffiffi

b14

p tanhcffiffiffiffib14

4q� �

þ 1

tancffiffiffiffib14

4p� �

0@

1A

0@

1A

sin2cffiffiffiffib14

4q� �

�sinh2cffiffiffiffib14

4q� �

C6S¼

Ea2gD11

sinhcffiffiffiffib14

4q� �

sincffiffiffiffib14

4q� �

V0þV1ffiffiffiffiffiffiffiffiðb1

4 Þ34

p þM0�M1ffiffiffiffib14

p 1

tanhcðffiffiffiffib14

4p

Þþ 1

tancffiffiffiffib14

4p� �

0@

1A

0@

1A

sin2cffiffiffiffib14

4q� �

þsinh2cffiffiffiffib14

4q� �

where c = is the half length in one bonded overlap section; Ea = theadhesive Young’s modulus; Ga = the adhesive shear modulus;T = the unit-width tensile load; V0, and M0 = the unit-width shearforce and bending moment at the outer bonded overlap edge; V1,and M1 = the shear force and bending moment at the inner bondedoverlap edge.

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