2.9 REFERENCES

Clayton, G. and Jones, D.P. (1976), Report SON (P) 149, British Aircraft Corp., Warton,
Preston, U.K.

Collings, T.A. (1977), “The Strength of Bolted Joints in Multi-Directional CFRP Laminates,”
Composites, Jan., vol. 8, no.1, pp. 43-55.

Collings, T.A. (1982), “On the Bearing Strength of Carbon Fibre-Reinforced Plastic Laminates,”
Composites, July, vol. 13, no. 3, pp. 241-252.

Collings, T.A. (1987), “Experimentally Determined Strength of Mechanically Fastened Joints,”
in Joining of Fibre-reinforced Plastics, ed.. Matthews, F. L., Elsevier.

Garbo, S.P. and Ogonowski, J.M., AFWAL-TR-81-3041, Vol.1.

Godwin, E.W. and Matthews, F.L. (1980), “A Review of the Strength of Joints in Fibre-
Reinforced Plastics - Mechanically-Fastened Joints,” Composites, July, vol. 11, no. 3, pp. 150-
160.

Godwin, E.W., Matthews, F.L. and Kilty, P.F. (1982), “Strength of Multi-Bolt Joints in GRP,”
Composites, July, vol. 13, no. 3, pp. 268-272.

Hart-Smith, L. J. (1977), “Bolted Joints in Graphite-Epoxy Composites;” NASA Langley CR
144899, Jan.

Hart-Smith, L. J. (1980), “Mechanically-Fastened Joints for Advanced Composites -
Phenomenological Considerations and Simple Analyses,” Fibrous Composites in Structural
Design, ed. E. M. Lenoe, D. W. Oplinger and J. J. Burke, Plenum Press, New York.

Hart-Smith, L. J. (1987), “Design and Empirical Analysis of Bolted or Riveted Joints,” in
Joining of Fibre-Reinforced Plastics, ed. Matthews, F. L., Elsevier.

Heath-Smith, J.R. (1979), Tech. Report TR 79085, Royal Aircraft Establishment, Farnborough,
UK.

Kim, R.Y. and Whitney, J.M. (1976), “Effect of Temperature and Moisture on Pin-Bearing
Strength of Composite Laminates,” J. Composite Materials, Apr. vol. 10.

Kretsis, G. and Matthews, F.L. (1985), “The Strength of Bolted Joints in Glass Fibre / Epoxy
Laminates,” Composites, Apr., vol. 16, no. 2, pp. 92-102.

Matthews, F.L. and Kalkanis, P. (1985), Proc. 5" Int. Conf. on Composite Materials, San Diego,
USA.

2-6


15 -
Effect of varying
joint width

zs 10/--
=
nod
8s
2g
2 Effect of varying
tg 5 joint overlap fength*

0

1 1 J
i) 10 20 30

Joint length or width (mm)

Figure 3.8 Effect of overlap length and width on single
lap joint strength (from Hutchinson, 1997).

FULLY PLASTIC ZONE (P&L) JOINT CROSS SECTION
P 8 € P— +p
TRANSI-
TION | MAXIMUM (CONSTANT) Lc _/ *
ZONE STRENGTH ZONE Tl IconsT

L
EFFECT OF LAP LENGTH ON
ADHESIVE-BONDED JOINT STRENGTH

——
| .
A. SHORT OVERLAP fee -7* |

C. LONG OVERLAP

B. INTERMEDIATE OVERLAP

4

Figure 3.9 Influence of lap length on bond stress
distribution. *Adhesive bond stress distribution
(from Hart-Smith, 1987)

3-9


I L
LAMINATED JOINT
AXIAL CONNECTIONS
l - =n 1
en

LAMINATED SCARF JOINT

— pp)

BONDED SCARF JOINT
BONDED INSERT JOINT
BASEMENT/SECTION CONNECTIONS
eo
a Fe
ae
wa
atte
Oar
CAST-IN JOINT BONDED INSERT/BASE PLATE CONNECTION

Figure 3.1 Joining techniques for pultruded sections.
(from Clarke, 1996)

3-5


Volume

‘ROOM TEMPERATURE

SHEAR STRESS

NTLONEPOXY ADHESWE (120°C CURE
° 1 2 3
SHEAR STRAIN

Loss in Joint strength (%/e)

50 L
° 2000 4000
Time of immersion in water ot 40°C (h)

Figure 3.17 (Upper) Effect of temperature on adhesive shear behavior
(from Matthews, 1987). (Lower) Effect of immersion in water of FRP / two-part
acrylic adhesive / joints on tensile lap shear strength (from Kinloch, 1987).

Stiffness

Ty Ty
Temperature Temperature

Figure 3.18 The glass transition temperature
(from Hutchinson, 1997).

3-14




Unloading
© First failure
Wmj_bt+bd

40

1

i) 20
mrad

Figure 4.8 Moment-rotation curves of isolated pinned connections

(Wmj = web cleat, bending about major axis, bt = bolting, bd = bonding)
(from Mottram and Zheng, 1996)

IS

75
100

75

Figure 4.9 Molded FRP “universal
connector” (from Mosallam, 1994)

4-8


Double thickness
Column top cheat (1 inch)

stiffener
‘Lower top cheat

M16 bolts
grade 4.6

No adhesive bond

L Packing

Column A M20 bolts
stiffener, grade 8.8
x
Sl 4
‘ WF 8x8x3/8 inch
10 m mm gap
a
- 1
4 M16 bolts
ap grade 4.6
:
ye 100x100x8
steel leg-angle
1
= =H
hal o -

Figure 4.5 Detail of FRP connection DTLmj (top) and steel connection
Stmj (bottom two). (from Mottram and Zheng, 1996)

4-6


system has been expanded further to include members that would be found in a multi-story
building (Figs. 5.10 and 5.11).

5.4 CONCLUSION
Interlocking connections use the light weight and ease of FRP to be fabricated into custom-made

shape to advantage. Although they are adaptations of historical designs with timber and cast
iron, they represent a significant and promising departure from steel-like connections.

5.5 REFERENCES

Goldsworthy, W.B. and Hiel, C. (1998), “Composite Structures are a SNAP,” Proc. Second
International Conference on Composites in Construction, ICCI 98, Tucson, Arizona., USA, 2,
pp. 382-396, Jan.

Head, P.R. and Churchman, Q.E. (1989), “Design, Specification and Manufacture of a Pultruded
Composite Construction System.” Proc., BPF Symposium on Mass Production Composites,
Imperial College, London, Sep., pp. 117-162.

Hutchinson, A.R. (1997), “Joining of Fibre-reinforced Polymer Composite Materials,” CIRIA
Report 46, Nov.

5-2


@ (b)

Figure 5.7 (a,b) “Snap” joint concept
(Designed by W.B. Goldsworthy & Associates, Inc.)
(from Goldsworthy and Hiel, 1998)

Figure 8 (a, b, c) Hardware for fasteners “Snap” joint
(Designed by W.B. Goldsworthy & Associates, Inc.)
(from Goldsworthy and Hiel, 1998)


Uniformly distributed loading

$= 13 mrad

ss. panins0

Bending
moments

Pinned Semi-rigid Rigid

Figure 4.1 Influence of end-restraint on bending moment distribution
and deflection of a beam (from Mottram and Zheng, 1996)

aa

fl
[

Principal fibre directions

Figure 4.2 Web-cleated connection
(from Huchinson, 1997)

4-4


d=6-35mm (torque 3-4 Nm)

|

Bearing failure

10004

Tension failure

‘Shear failure

Bearing stress (MN/m?)

oO T T T T 700
° + 45*plies (%)

Figure 2.5 Influence of CFRP fiber proportion (0° /+45°)
on failure mode (from Garbo and Ogonowski, 1981)

Torque tight Ge=22 mN/m?
1000-4 Nim
"
2
= ry
rs
2 Light Fingerctight
Fy
:
g
£
5
3
es 4 .
Free pin

500} ——1

U 50 rn

Proportion of £45"plies (%)

Figure 2.6 Effect of clamping pressure on the bearing
strength of HTS/914 0°/445° laminate (from Collins, 1982)

2-10


Adhesive -
normal stress tension

compression

ee

2P

(Se

Figure 3.10 Bending moments and induced stresses in the
outer adherends of a double lap joint (from Adams, 1987).

Principal tensile
N stresses in adhesive

S
VAAN SS YAS
VAN ANN XN

~
\ vs Paty SANS

Probable failure path
(a) (b)

Figure 3.11 Lap joint formed with
spew fillet (from Hutchinson, 1997).


ABSTRACT

This review of the state of the art of connections of fiber-reinforced polymer (FRP) members
covers primary structural connections (whose integrity is crucial to the functioning and safety of
a structure), in which at least one of the members is made of FRP. The connections consist of
mechanical joints, bonded joints, interlocking joints, or a combination of those, such as frame
connections. After an earlier period of benefiting from aerospace research and imitation of steel
connections, FRP connection design in civil infrastructure is beginning to develop in new,
original directions

Keywords: adhesive, bolt, bond, composite, connection, fiber-reinforced polymer, frame, FRP,
interlocking, joint.

ii


mm (0.001 in) causes a stress concentration of 25 % in the elastic adhesive shear stress
distribution (Thamm, 1976). Stiffness balance of the adherents is an important consideration
here as well. Figures 3.15 and 3.16 show a variety of bonded joint configurations and their
relative strength.

3.5. ENVIRONMENTAL EFFECTS

The performance of adhesives is, in general, highly affected by temperature and humidity. As
with all polymers, the stiffness and strength of an adhesive drop sharply beyond the glass
transition temperature (Figs. 3.17 and 3.18). The material is rubbery above the glass transition
temperature, and glass-like below it. Absorption of water plasticizes adhesives and lowers their
mechanical performance (Figs. 3.19 and 3.20). For joints sustaining permanent loads, the visco-
elastic properties of polymers must be taken into account.

3.6 SURFACE PREPARATION

For surfaces to bond properly, they must be free of grease, dust, moisture or other release agents.
For most applications, surface preparation simply involves solvent degreasing, abrasion and dust
removal. FRP surfaces should be abraded until they lose their gloss or luster and become dull.
However, one should be careful not to break reinforcing fibers near the surface.

A peel ply is sometimes used if a very clean and reproducible surface is needed for bonding on
site. It is a sacrificial layer of fabric laid up on the outermost surface and co-cured with the FRP.
This layer is peeled off prior to bonding, leaving behind a clean, rough surface for bonding.

Corona discharge treatment is a factory technique ideal for GFRP skins of large sandwich panels.
It makes use of a high frequency electrical discharge, which ionizes atmospheric oxygen and
produces a highly oxidized, polar surface with increased surface free energy.

3.7 SUMMARY ’

Adhesive joints can be highly effective when large surfaces are available for joining and the
members are thin. Careful design is necessary to minimize stress concentrations and peeling
stresses. A trade-off must usually be made between strength and deformation requirements.

Adhesive joints have been used in structural applications, but usually under tightly controlled
factory conditions, as found in the aerospace industry. The bonded joints must be designed to
withstand environmental conditions found in their operational life, and frequent inspections must
be scheduled to ensure their integrity.

In civil infrastructure, adhesive joints are usually not relied upon as the sole means of
connection, mostly because of the thickness of the members to be joined, and the difficulty in
controlling quality in field assembly. In cases where members are fastened together
simultaneously by bolts and adhesive, joint strength is calculated based on the bolts only,
whereas the bond provides additional stiffness and safety. For thin and moderately loaded
structures, adhesive bonding provides a fail-safe load path to protect the structure in the vicinity
of the fastener holes. On the other hand, thick, highly loaded, adhesively bonded structures need
mechanical fasteners to protect the bond from unzipping catastrophically from any small initial
damage.

3-3



Potter, R.T. (1978), “On the Mechanism of Tensile Fracture in Notched FRP,” Proc. R. Soc.
London, vol. A361, pp. 325-341.

Quinn, W.J. and Matthews, F.L. (1977), “Effect of Stacking Sequence on the Pin-Bearing
Strength in Glass Fibre-Reinforced Plastics,” J. Composite Materials, Apr., vol. 11, pp. 139-45.

Shivakumar, K.N. and Crews, J.H. (1982), Tech. Memo 83268, Jan., NASA, Langley, Va., USA.

Stockdale, J.H. and Matthews (1976), F.L., “Effect of Clamping Pressure on Bolt-Bearing Loads
in Glass Fibre-Reinforced Plastics,” Composites, Jan., vol. 7, no. 1, pp. 34-38.


Failing load (kN)
1

rs

?
ZrIN eoo-—e
“N71 cove
7-N\I eee
147\-— ote
—17\  ee—ee
FIN eee
7N|- eee

+ + + t

Stacking sequence

Figure 2.10 Effect of stacking sequence on bearing strength
for GFRP laminates (from Quinn and Matthews, 1977)

Figure 2.11 Terminology used for describing a pin-loaded hole

2-13


7
I oapES LOADED HOLE TESTS + 38 -
? ee

“2sgo27 BEARING FAILURES
more TESTS v (APPARENTLY PREMATURELY

ab YS WITH RESPECT TO

y Y TENSILE STRENGTHS)
STRESS CONCENTRATION us 9 of
Le HEF oR
FACTOR hte 7,2 TENSION FAILURES A REIN, os
ON NET SECTION , “ef re RN Kann
[- “S TRANSITION IN PICKS CON OK
OF FIBROUS COMPOSITE } ae tba Root pict
LAMINATE AT FAILURE S coe
4
LS
2F “ce 3 COMPLETE
1& % STRESS CONCENTRATION
oe TYPICAL TEST RESULTS® RELIEF
>? of CHANGE IN FAILURE MODE
1 ! Ll | zl !
1 2 3 4 5 6 7 8
OPEN HOLES 2 4610 THEORETICAL ELASTIC STRESS CONCENTRATION FACTOR, Kye
APPROXIMATE Sa j j '
w/d VALUES 1 " 1. L
LOADED HOLES 2 3 4 5 6

Figure 2.3 Relation between stress concentration factors observed at

7

failure of fibrous composite laminates and predicted for perfectly
elastic isotropic materials (from Hart-Smith, 1987)

BEARING FAILURE { w/d =8)

LOAD NET TENSION FAILURE ( w/d=3)

DEFLECTION

Figure 2.4 Load-deflection curves for bolted
composite joints (from Hart-Smith, 1987)

2-9



3.8 REFERENCES

Adams, R.D. (1987), “Theoretical Stress Analysis of Adhesively Bonded Joints,” Joining Fibre-
Reinforced Plastics, Matthews, F.L. ed., Elsevier.

Adams, R.D. and Peppiatt, N.A. (1973), “Effect of Poisson’s Ratio Strains in Adherends on
Stresses of an Idealized Lap Joint,” J. Strain Anal., Apr. vol. 8, no. 2, pp. 134-139.

Adams, R.D. and Wake, W.C. (1984), Structural Adhesive Joints in Engineering, Elsevier.

Allman, D. J. (1977), “A Theory for Elastic Stresses in Adhesive Bonded Lap Joints,” Quarterly
J. Mechanics Appl. Maths, vol. 30, pp.415-436.

Clark, J.L., ed. (1996), “Structural Design of Polymer Composites,” EUROCOMP Design Guide
and Handbook, E&FN SPON, 751 pp.

Goland, M. and Reissner, E. (1944), “The Stresses in Cemented Joints,” J. Appl. Mech., Trans.,
ASME, vol. 66, A17-A27.

Hart-Smith, L.J. (1987), “Design of Adhesively Bonded Joints,” Joining Fibre-Reinforced
Plastics, Matthews, F.L. ed., Elsevier.

Hutchinson, A.R. (1997), “Joining of Fibre-Reinforced Polymer Composite Materials,” CIRIA
Project Report 46, Nov.

Kinloch, A.J. (1987), Adhesion and Adhesives, Chapman and Hall.

Kutscha, D. and Hofer, K.E. (1969), “Feasibility of Joining Advanced Composite Flight Vehicle
Structures,” AD 690 616, IIT Research Institute, Chicago, Il.

Matthews, F.L., Kilty, P.F. and Godwin, E.W. (1982), “A Review of the Strength of Joints in
Fiber-Reinforced Plastics — Part 2: Adhesively Bonded Joints,” Composites, Jan., vol. 13, no. 1,

pp. 29-37.

Renton, W. J. and Vinson, J.R. (1975), “The Efficient Design of Adhesive Bonded Joints,” J.
Adhesion, vol. 7, no. 3, pp. 175-193.

Thamm, F. (1976), “Stress Distribution in Lap Joints with Partially Thinned Adherents,” J.
Adhesion, vol. 7, pp. 301-309.

Volkersen, O. (1938), “Die Nietkrafturteilung in Zugbeanspruchten Nietverbindungen mit
konstanten Loschonquerschnitten,” Luftfahrtforschung, vol. 15, pp. 41-

3-4


for a combination of 0° / + 45° plies. Figure 2.5 shows that, at low levels of + 45° fibers, shear
failure predominates. As the number of + 45° fibers increases, so does shear strength, until
bearing becomes the critical mode of failure. Since + 45° laminates are weak in tension, further
increases in the proportion of + 45° fibers changes the failure mode to tension.

2.4.2 Lateral Constraint

Lateral constraint due to clamping pressure can significantly increase joint strength, although.
over-tightening of bolts may cause surface damage to the laminate. Collings (1977) and Garbo
and Ogonowski (1981) recommended an optimum bolt clamping pressure of 22 MPa for CFRP.

Resin creep causes clamping pressure to decrease in time (Shivakumar and Crews, 1982). This
behavior is more pronounced in high temperature and high moisture conditions (up to 60%
clamping pressure ‘ss at 60 ° C, 1 % moisture content). This loss in clamping pressure does not
necessarily translate in an equal loss in joint strength. Indeed, finger-tight constraints produce
joint strength of 88 % of optimally constrained joints for dry, 0°/ + 45° laminates at room
temperature (Collings, 1982). See Figs. 2.6-2.8.

24.3 Stacking Sequence

Collings (1977) showed that, for bolted CFRP joints consisting of 0° (2/3) and + 45° (1/3) plies,
there was no difference in shear strength, but a difference of 6 % in tensile strength between two
different stacking sequences. Bearing strength, however, shows a significant drop (16 %) for the
more "grouped" laminates (Fig. 2.9, Table 2.1).

For pin-loaded holes (without clamping pressure), Quinn and Mathews (1977) showed that a
90°/445°/0° la-inate produced the highest bearing strength, and a 0°/90°/+45° the lowest, with a
relative reduct..0n of 30 % (Fig. 2.10).

244 Joint Geometry

Width: In contrast to metals, FRP cannot take advantage of plasticity to relieve stress
concentration near bolt holes. As a result, for a given hole diameter, net tensile failure stress
depends strongly on width. The effect of width is most marked for laminates with a high
proportion of 0° fibers and least for + 45° laminates (Collings, 1977, Figs. 2.11 and 2.12).

Hole Size: Hole size has little influence on the net tensile strength and shear strength of 0°/ +
45° CFRP (Collings, 1977) or 0°/ + 45° GFRP (Kretsis and Matthews, 1985). The bearing
strength of CFRP is unaffected by hole diameter d provided sufficient clamping pressure is
provided. However, for GFRP, the low elastic modulus of glass causes out-of-plane cracking for
d/t > 3 regardless of Jateral constraint, in 0°/ + 45° lay-ups. Values of d/t < 1.5 should not be
used due to the risk of bolt shear. Figures 2.13 - 2.15 show the effect of hole size, non-
dimensionalized with respect to plate thickness t or width w. Joint strength is limited by bearing
failure for small d/w ratios, and by tension failure through the hole for large d/w ratios. A
reduction in end distance e, below a certain value of e/w ~ 1 further decreases joint strength.

2-4


WIS S555

Optimized
shape

Figure 5.3 Concepts for sandwich panel connections
(from Hutchinson, 1997)

PTT Ty yds

PUNK

0

TOCCLE CONNECTOR

SWAY CONNECTOR (GROOVE CONNECTOR FLAT TRIMMER, CHANNEL CORNECTOR

Figure 5.4 Principal components of advanced composite
construction system (ACCS) (from Head and Churchman; 1989)

5-4


1.1 References

Clark, J.L., ed. (1996), “Structural Design of Polymer Composites,” EUROCOMP Design Guide
and Handbook, E&FN SPON, 751 pp.

Hutchinson, A.R. (1997), “Joining of Fibre-reinforced Polymer Composite Materials,” CIRIA
Report 46, Nov., 182 pp.

Matthews, F.L., ed. (1987), “Joining Fibre-Reinforced Plastics.” Elsevier, 319 pp.

Strongwell (1998), “EXTREN Design Manual.”


Chapter 3
Bonded Connections

3.1. INTRODUCTION

As discussed in the previous chapter, FRP members can be joined by bolts or rivets. However,
these mechanical fasteners sometimes require the use of load-spreading inserts bonded to the
structure (Fig. 3.1, top). An alternative method of joining is by bonding using adhesives. An
advantage of bonded joints is the possible use of co-curing techniques, in which the composites
are joined without using a separate adhesive. In this process, the composites are prepared in their
uncured (pre-impregnated fiber) form and heat and pressure are applied. As the composite
begins to cure, the pressure squeezes out the excess matrix material, which contacts the adjacent
component. When curing is complete, the components are bonded together. However, this
technique would be difficult to apply to large-scale civil structures, which tend to be field-
assembled. This overview of bonded joints covers stress analysis, failure modes, design
implications and the effects of the operating environment.

3.2. STRESS ANALYSIS

With the assumption that the adherents are rigid and the adhesive deforms elastically and only in
shear, a simple stress analysis of an adhesively bonded, single-lap joint produces a uniform shear
stress distribution in the adhesive. While this simple analysis is approximately true for a short
overlap, Volkersen’s (1938) more refined analysis, which accounts also for the elastic
deformation in tension of the adherents, shows shear stress concentration at the ends of the
connection (Fig. 3.2).

Goland and Reissner (1944) accounted for bending effects due to load eccentricity in a single
lap joint (Fig. 3.3). In an undeformed lap joint connecting two plates of thickness t, the moment
due to eccentricity of load P is M = P t/2. The deformation of the joint reduces this moment by
a factor K<1. Analysis also shows the presence of transverse normal tensile stresses that cause
peeling near the ends of the joint. Since modern adhesives can be stronger than the polymer in
FRP, this normal stress can cause tensile failure in the adherents if the interface is also strong.

More recent analytical refinements include: zero stress boundary conditions at the ends of the
adhesive, non-uniform shear and normal stress distributions across the thickness of the adhesive,
shear deformation in the adherents, and stiffness of the adhesive (Kutscha and Hoffer, 1969;
Renton and Vinson, 1975; Allman, 1977). A more flexible adhesive produces more uniform
shear and normal stress distributions than a stiffer one (Fig. 3.4), the latter being sometimes
required to limit deformations. Adams and Peppiatt (1973) showed the existence of significant
shear stresses in the adhesive and normal stresses in the adherents in the direction transverse to
the applied loads, across the width of a lap joint. According to Matthews et al. (1982), it is
necessary to consider non-linear adhesive behavior in order to correctly predict joint strength. In
spite of recent advances, numerical models are no substitute for testing and experience, which
are still necessary, especially for a joint that is frequently used.


FURTHER STRENGTH REDUCTION

UNATTAINABLE 7 SX DUE TO NET-SECTION STRESS
0.5 BEARING t '. CONCENTRATION FACTOR
“> F STRENGTH { a
STRENGTH REDUCTION
Ly, DUE TO NET AREA LOSS d
0.4-- i wy BECAUSE OF HOLE 7 r
L
JOINT EFFICIENCY , a
fy
ad i 4 L e/wet P
/ a
O.2-- 4
JOINT STRENGTH 3 BEARING FAILURES
BASE LAMINATE !
STRENGTH i UNATTAINABLE. TENSION. THROUGH-THE-HOLE
Ol TENSION FAILURES
STRENGTH
INCREASING BEARING STRENGTH
0 1 i 4
0 0.2 O4 0.6 0.8

RATIO OF BOLT DIAMETER TO STRIP WIDTH d/w

Figure 2.14 Bolted joint efficiencies for composite laminates
as functions of d/w ratio (from Hart-Smith, 1987)

NO PLATEAU
50
HTS CARBON-EPOXY FIBERGLASS-EPOXY
P
40 / TENSION STRENGTH {
vA
0b AL
4 PLATEAU
JOINT STRENGTH / +
(Ks!) / i
2b 77 BEARING STRENGTH
4
cA
vA TORQUED BOLT
10 1/7 \— pm.coanen HoLe
_,’
0 ! L ! !
) 0.2 04 06 08 1.0
d/w RATIO

Figure 2.15 Relation between bearing and tension failures for
carbon/epoxy and fiberglass/epoxy laminates as functions of
d/w ratio (50 ksi = 350 MPa; from Hart-Smith, 1987)

2-16


CONTENTS

Abstract
Acknowledgment
Disclaimer.

1. Introduction
1.1 References.

2. Bolted Connections.
2.1  Introduction..
2.2 Choice of Fasteners.
2.3. Failure Modes.......
2.4 Important Factors Affecting Joint Strength.
2.5 Effect of Temperature and Moistur
2.6 Fatigue Strength
2.7 Conclusion
2.8 Future Research
2.9 References...

.1 Introduction. .
3.2 Stress Analysis.
3.3 Failure Modes..
3.4 Design Implications...
3.5 Environmental Effects
3.6 Surface Preparation....
3.7 Summary...
3.8 References.

4. Frame Connections.
4.1 Introduction
4.2 Beam-Column Connections.
4.3 Pinned Connections...
44 Semi-Rigid Connections.
4.5 References.

5. Interlocking Connections.... 5-1
5.1 Tubes and panels... 5-1
5.2 Advanced Composite Construction System (ACCS) ve S-1
5.3. “SNAP” Joints...........:::eceeeeeeeeee 5-1

5-2
5-2

5.4 Conclusion.
5.5 References.


ACKNOWLEDGMENT

The support of NIST Advanced Technology Program (Felix Wu, Program Manager, Composite
Materials) and the efforts of Steve Cauffman, Nicholas Carino, Joannie Chin and Geraldine
Cheok in reviewing this document are gratefully acknowledged.

DISCLAIMER

Trade or manufacturer’s names appear herein because they are essential to the objectives of this
document. The United States government does not endorse products or manufacturers.

iv


Interaction between holes: Joint geometry is generally chosen such that potential tensile or shear
failures occur simultaneously at a mean stress as close as possible to the bearing failure stress. If
these requirements are met, and adequate spacing is provided for ease of installation, then
interaction between adjacent bolts and holes is minimal. On the other hand, matching the
stiffness of the joining parts equalizes the load distribution to the bolts and maximizes joint
strength (Fig. 2.16).

2.5 EFFECTS OF TEMPERATURE AND MOISTURE

Exposure of FRP to a wet environment allows the absorption of moisture into the resin
matrix, resulting in plasticization and a lower glass transition temperature. Joint properties most
affected are shear and bearing strengths. Kim and Whitney (1976) showed that temperature had
a more significant effect than moisture on bearing strength. The combination of temperature and
moisture produces a further loss of bearing strength of 10 % compared with the effects of
temperature alone. They measured a maximum bearing strength reduction of 40 % in the case of
combined temperature and moisture.

2.6 FATIGUE STRENGTH

The fatigue performance of bolted CFRP joints can be better than that of open holes. Thus
mechanical joints of composite materials are unlikely to be fatigue critical, in contrast with
aluminum alloys (Clayton and Jones, 1976, Heath-Smith, 1979). Garbo and Ogonowski (1981)
showed that residual strength is generally equal to or greater than the strength of non-fatigued
specimens, because of progressive relief of stress concentration. However, in most cases, the
specimens sustained hole elongations of 0.05 d, a value greater than that normally allowed in
metallic joints. Therefore, elongation, rather than strength, may be the governing fatigue
criterion for CFRP joints.

2.7 CONCLUSION

Mechanical fasteners are an effective method of joining FRP members. The best overall
performance for the three main failure modes (tension, shear and bearing) is obtained for 0° / +
45° laminates. As a general rule, there should be between 1/8 and 3/8 of the fibers in any one of
the basic laminate directions, 0°, + 45°, - 45° and 90° (Hart-Smith, 1987). Maximum joint
strength is developed when the joint geometry is designed to suppress tension and shear failures,
and sufficient clamping pressure across the thickness of the laminate is provided.

2.8 FUTURE RESEARCH

The EUROCOMP Design Guide and Handbook (1996) provides general design procedures for
mechanical connections loaded in shear or tension, as well as a more rigorous finite-element
method. The Guide recommends designing other mechanical joints by testing. Clearly, further
research is needed to define design procedures for all mechanical joints.

The use of thicker plies, or the stacking together of thinner plies significantly reduce laminate
strengths, a fact not accounted for in most laminate theories. Since civil engineering applications
generally use thicker members than aerospace engineering does, more research in the joining of
thicker elements is necessary. The development of stronger and tougher resins is also considered
the key to better joints (Hart-Smith, 1978).

2-5


Chapter 1
Introduction

Fiber-reinforced polymers (FRP) represent a new class of construction materials. In the past two
decades, their use has spread from the aerospace industry to civil infrastructure, which presents a
new set of challenges. For seismic rehabilitation, such as wrapping of bridge columns, and
applications where corrosion is an important consideration, FRP are competitive with more
traditional construction materials. Compared to these, FRP offer significant advantages, such as
resistance to corrosion, high strength, light weight and ease of installation. Their disadvantages
include high initial cost, possible durability problems in the presence of moisture, ultra-violet
radiation, freeze-thaw cycles, or sustained stresses, and the absence of design standards and
experience.

FRP can be made into structural shapes by pultrusion, a process that combines extrusion and
pulling of molten or curable resin and continuous fibers usually arranged in unidirectional layers,
or plies, through a die of a desired structural shape of constant cross-section. The assembly of
structural members, made of FRP or other materials, requires connections, which are therefore
integral parts of a structure and can be classified as:

e Primary structural connections, whose integrity is crucial to the function and safety of a
structure. They are designed to provide major strength and stiffness to the structural
assembly for the life of the structure.

e Secondary structural connections, which provide some strength and stiffness to the assembly,
but their failure would not cause damage beyond the structural components they join. An
example is the field connection of two modular units.

e Non-structural connections, e.g., joints of decorative panels.

Although FRP structural members can be made in a variety of shapes (Fig. 1.1), large volume
construction requires standard shapes, which must be limited to certain sizes for ease of
transportation and assembly. In addition, individual building details vary greatly in geometry,
making assembly of standardized components much more practical than custom fabricated
members. For all these reasons, connections are necessary components of a structure. Figures
1.2, 1.3 and 1.4 show examples of FRP connections in civil structures. It is apparent that they
are similar to steel bolted connections, an expected development in this early stage of use of a
relatively new material such as FRP. It will be seen in this review that there are, in fact,
important differences between the behavior and design of steel and FRP connections.

This review of the state of the art of connections of FRP members covers primary structural
connections, in which at least one of the members is made of FRP. The connections consist of
mechanical joints (Chapter 2), bonded joints (Chapter 3), interlocking joints (Chapter 5), or a
combination of those, such as frame connections (Chapter 4). For more details, the reader is
referred to excellent reviews of the subject, such as Hutchinson’s (1997) “Joining of Fibre-
Reinforced Polymer Composite Structures”, and Matthews’s (1987) “Joining Fibre-Reinforced
Plastics”. Also of interest is the EUROCOMP Design Guide and Handbook (1996), “Structural
Design of Polymer Composites”, which is a comprehensive design guide for FRP structures
following a Load and Resistance Factor Design (LRFD) format.


200
BEARING STRENGTH
(KSI)
400- HTS CARBON-EPOXY
QUASHISOTROPIC LAMINATE PATTERN
Q PINLOADED HOLE 0.25 INCH DIAMETER BOLTS
NORMAL TORQUES NORMAL METAL
FINGER-TIGHT FOR COMPOSITES BOLT TORQUES
pt L tL
0
BOLT TORQUE

Figure 2.7 Effect of bolt torque on bearing strength of fibrous composite
laminates (100 ksi = 690 MPa, 1 in = 25 mm) (from Hart-Smith, 1987)

140

H
120}- 8
8
100 |-
80;-
cuales ENGTH 60 2g 7300/N5208 CARBON-EPOXY LAMINATES
8 QUASHISOTROPIC LAMINATE PATTERN
40 BOLT DIAMETER = 0.25 INCH
TESTS CONDUCTED
20 AT ROOM TEMPERATURE
° PIN CONNECTION PROTRUDING-HEAD BOLT

Fig 2.8 Comparison between bearing strengths of connections with
and without clamping pressure (100 ksi = 690 MPa, 1 in = 25 mm)
(from Hart-Smith, 1987)

2-11



NISTIR 6532

of tnetrdeere

Connections of Fiber-Reinforced
Polymer (FRP) Structural Members:
A Review of the State of the Art

Si RT

Dat Duthinh

August 2000

Structures Division
Building and Fire Research Laboratory
Gaithersburg, Maryland 20899

NIST

United states Department of Commerce
Technology Administration
National Institute of Standards and Technology

@ 1007-1061





iJ
e
%

mm
Ly Bi Stele
PR AL. SR UDD

£

2
z.

Fa

£

Le = JOR §

or
ft
£
x

on

tt

a4
By

{a) Beam — column

a
Loy
ig
we

.
3
3 4 Be
reo
ES
o. liv . _ 2 zz
* a & e
8 §| & Zo
ate af tT = at al Tat 3 ef
BS +t 3
Bs & &.

7
F

é
3

L4x3xte L4x4xte

ial

(nP) x
RB tp x Lp rs

SECTORS ~SecTOn_(By

Figure 1.2 Examples of connection details for standard pultrusions.
(dimensions in inches, 1 in = 25.4 mm, from EXTREN Design Manual)

£ tp x lp

1-4


{a) Standard sections (bd) Non-standard sections

Figure 1.1 Standard and non-standard pultruded sections
(dimensions in mm; from EXTREN Design Manual)

1-3


) :

Figure 5.1 Examples of bonded tubular joints
(from Hutchinson, 1997)

a
Cox

Figure 5.2 Examples of bonded interlocking system components
(from Hutchinson, 1997)



_—-——
GRAPHITE-EPOXY GRAPHITE-EPOXY }
TITANIUM i TITANIUM eS

CONSTANT. THICKNESS (STIFFNESS-UNBALANCED} JOINT CONSTANT-STIFFNESS (VARIABLE -THICKNESS) JOINT

pw) fhe uu

ADHESIVE SHEAR STRESS DISTRIBUTION ADHESIVE SHEAR STRESS DISTRIBUTION

Figure 3.14 Stepped lap bonded joints (strength improvement by matching stiffnesses)
(from Hart-Smith, 1987)

STEPPEO-LAP JOINT
FAILURES SHOWN
REPRESENT THE LIMIT ON

EFFICIENT DESIGN FOR
EACH GEOMETRY

BONDED JOINT STRENGTH

2g
BY |

& PEEL FAILURES

i
&
SF BENDING OF
Roy SINGLE-LAP JOINT ADHERENDS DUE TO.
_ ECCENTRIC LOAD PATH

ADHEREND THICKNESS

Figure 3.15 Relative strength of different bonded
joint types (from Hart-Smith, 1987)

3-12




Chapter 4
Frame Connections

41 INTRODUCTION

The previous chapters covered connections designed primarily to transmit in-line forces, with
secondary bending moments caused by the geometry of the connectors. These types of
connections have benefited from research in aerospace engineering. In contrast, frame
connections (between beams and columns) are primarily of concern to civil engineers and are
designed to transfer moments as well as shear and in-line forces between thick members, such as
I-sections.

4.2, BEAM-COLUMN CONNECTIONS

Beams transfer floor and roof loads to columns to which they are joined by-rigid (moment-
resisting), flexible (free to rotate), or semi-rigid connections. For uniformly loaded beams,
flexible connections (or simple supports) cause the maximum moment to occur at midspan,
whereas rigid connections (or fixed supports) shift the maximum moments to the ends. Semi-
rigid connections provide an opportunity to equalize the maximum negative moments (upward
bend) at the supports with the maximum positive moment (downward bend) at midspan, an
efficient design for a symmetrical cross-section, such as an I-shape (Fig. 4.1).

In steel buildings, beam-column connections usually consist of angles or cleats bolted or welded
to the flanges of the column and the beam, or the flanges of the column and the web of the beam.
In addition, gusset plates or stiffeners are sometimes used where necessary to tie the column
flanges together, thus transferring the beam load to both column flanges and mitigating web
crippling. The same designs are essentially carried over to FRP beam-column connections,
which include bolts only or a combination of adhesive bonding and bolts. Whereas steel frame
connections are designed primarily for (ultimate) strength limit state (failure under overload
conditions), concern with FRP connections is more with limiting deflections (serviceability limit
state) and preventing instability.

4.33, PINNED CONNECTIONS

Figure 4.2 shows a FRP connection consisting of an angle bolted to the web of an I-section and
designed to act essentially as a pinned connection. It must be able to resist shear force and
accommodate rotation. The use of adhesive alone is not recommended because it may lead to
catastrophic failure at small rotation or deflection due to inter-ply splitting of the web angle.
Bolting alone can satisfy both ultimate and serviceability limit states, e.g., the maximum
deflection must be less than 1/250 of the beam span [EUROCOMP (Clark, 1996)]. A
combination of bonding and bolting enables the connection to incur no slip even at the
serviceability limit, but localized cleat failure may occur.

The EUROCOMP Design Code (Clark, 1996), which is a practical design code limited to glass

FRP components, connections and assemblies for the construction industry, recommends bolt
hole clearance of 2 mm (0.08 in) for constructibility, and a gap of 10 mm (0.4 in) between beam

41



BOLT LOAD DISTRIBUTIONS
4-ROW BOLTED JOINT

14,
12--

10-

BOLT LOAD
(1000 LB)

fymmemnd’ CONFIGURATION 1
Oo. Q CONFIGURATION 2
O==+=0 CONFIGURATION 3
m= 8 CONFIGURATION 4

it i 1 1

CONFIGURATION 4
Eros * 0.006 INJIN.

CONFIGURATION 3
Eyou * 0.0046 NIN,

CON FIGURATION 2
Coy * 0.0045 INJIN,

9.60 IN. 1234 9.25 IN,

1 2 3 4
BOLT NUMBER

—o
t

=

|
CONFIGURATION 1
Eyrou = 0.0088 ININ.

Figure 2.16 Effect of joint configuration on bolt load distribution
(1000 Ib = 4 448 N, 1 in = 25 mm; from Hart-Smith, 1987)

123 4 oa6in,

—
—




SHEAROUT
FAILURE

TENSION FAILURE BOLT PULLING THROUGH LAMINATE

\\

“—

CLEAVAGE -TENSION FAILURE BEARING FAILURE BOLT FAMLURE

Figure 2.1 Modes of failure for bolted joints in
FRP composites (from Hart-Smith, 1987)

) DELAMINATION

NO CHANGE IN

FIBER STRESS

<= THROUGHOUT,
HIGH PEAK —f DELAM|
TENSION STRESSES REGION

IN FIBER
ADJACENT | HOLE

REDUCED PEAK FIBER
STRESS BECAUSE
STRAIN SPIKE IS
AVERAGED OVER
FINITE LENGTH

pos IN FIBERS ALONGSIDE HOLE

Figure 2.2 Stress concentration relief in fibrous
composites by delamination. (from Hart-Smith, 1977)

2-8


Chapter 5
Interlocking Connections

A major advantage of FRP construction is the ease of transportation and construction that comes
with light-weight members. Interlocking joints, sometimes bolted or bonded as well, are
particularly well suited for FRP construction because they allow connections to be made quickly.
Geometric interference and surface friction may provide sufficient integrity, without any
adhesive or mechanical fasteners, for some large, interlocking structures to carry light loads.
The main disadvantage of such connections is that members must be fabricated to close
dimensional tolerances.

5.1. TUBES AND PANELS

Tubular FRP members benefit from a long tradition of pipe fitting. Figures 5.1 and 5.2 show
some examples of interlocking and bonded joints for circular and square tubing. Figure 5.3
shows a system of construction of sandwich panels that can be assembled quickly by inserting
them into ready-made corner and other joints, followed by bonding or bolting. The panels carry
flexural loads and consist of glass FRP skins bonded to a low-density core.

The interlocking connectors should maximize flexibility of orientation of standard components.
A truly universal connector requires a large number of fibers to run in as many directions as
possible, which is difficult to achieve with pultrusion. Rather, a molding process is used for
special or low volume connectors.

5.2 ADVANCED COMPOSITE CONSTRUCTION SYSTEM (ACCS)

Interlocking joints are mostly ingenious, proprietary inventions protected by patents. The ACCS
consists of pultruded panels joined together by special pultruded dog-bone connectors that can
also be adhesively bonded, if dismantling of the cellular, monocoque structure is not planned
(Fig. 5.4). (A monocoque structure possesses sufficient strength and stiffness by itself, here
through mechanical interlocking alone, and does not require additional framing). The connectors
are constructed with ridges that control the thickness of the bond line (Fig. 5.5). The system is
designed for stresses up to 5.0 kPa (0.7 psi) and a serviceability limit strain of 0.3 %
(Hutchinson, 1997). It is suited for one or two story field offices or emergency buildings (Fig.
5.6) and bridge enclosures (designed to protect structural members permanently from the
environment).

5.3. “SNAP” JOINTS

The “SNAP” joint is an interesting type of connection that locks in place and cannot be
dismantled. Figures 5.7 and 5.8 show pultruded members whose fiber architecture is designed so
that the load-bearing faces of the ridged connectors have very high interlaminar shear strength.
These connections were an integral part of a 28 m transmission tower (Fig. 5.9) that was
assembled on its side, then uprighted by a crew of three in a few hours (Goldsworthy and Hiel,
1998). Light weight was crucial to this construction and allowed significant cost savings. The


Figure 5.11 “Snap” joint design for composite high rise
(Designed by WBG&AAI for Commercial Spaceport .S.A.)
(from Goldsworthy and Hiel, 1998)

5-8


Adherend

c _ Adhesive

=

High strain
concentration
High strain
concentration

t{shear)

Shift adhesive (high E)

Ty

& a(shear)

2 ale <7
2 av
3 Ne, _

Ss 0 ===

3

2 (normal)

3S

Ss

a

Figure 3.4 Typical elastic bondline stress distributions in a single lap joint
under load, made with adhesives of low and high elastic modulus.
(from Hutchinson, 1997)

‘Tensile stress

Intertaminar shear stress

‘interface shear stress
Peel stress.

Figure 3.5 Failure modes of single-lap bonded joints


Reseave

SHEAR STRENGTH
RESERVE woe
PEEL STRENGTH whe

STRENGTH. P
ADHESIVE PEEL FAILURES

ADHESIVE SHEAR FAILURES

ADHEREND FAILURES.

ADHEREND THICKNESS ¢

Figure 3.6 Relative severity of adhesive shear and peel
stresses in single lap-joints (from Hart-Smith, 1987).

AT oom

Figure 3.7 Peel stress failure of thick composite
joints. (A), (B) and (C) indicate failure sequence.
(from Hart-Smith, 1987)


M.

H

1 inch @ composite rod Pre-preg cleat
with square nuts — Pultruded stiffener
' Outer fillet radius
= of 40 mm
na
ae
Hy L6x6x1/2 inch
H
2 an.
5 —

“a 1 M16 bolts grade 4.6
Dimensions: mm

ya Ls |

Figure 4.6 FRP connection TLmj
(from Mottram and Zheng, 1996)

@ First failure

rn L 2 es
0 10 20 30
mrad

Figure 4.7 Moment-rotation curves of isolated semi-rigid connections
DTLmj = Double thickness top cleat, single thickness lower cleat, bending about major axis;
TLmj = top and lower cleats, major axis; Stmj = Steel connection of steel member, major axis.
(from Mottram and Zheng, 1996)

4-7


Net tensile strength (MN/m? )

10005 XAS/914 vt 0-6
d=6-35mm

Torque 3-4Nm

4 4245°

°

2% 0% *45°

3% 0°'2 t4s°
4% 0°'% 45°

4
4

9 T T T |
oO 40

Wi ¢mm)>

Figure 2.12 Variation of net tensile strength with width for 0/+45° CFRP
composites with 6.35 mm hole (from Collings, 1987)


Figure 4.12 | Schematic of connection test setup.
(from Banks et al., 1994)






Shear strength (MN m2)

&

Dry
€
30 Dry = K-4) Wet
2
3
Wet 3S
5 H 15 20°C
0 4 1 ti)
-50 i) 50 100
Temperature (°C) Shear strain (%)

Figure 3.19 Typical effects of temperature and moisture on the mechanical
behavior of an unmodified cold-curing epoxy adhesive in shear.

(from Hutchinson, 1997)

30
< Dy
=
z
= 20
3 Wet
2
a
g 1 20°¢
co
&
0 L 1 |
0 1 2 3
Tensile strain (%)

Figure 3.20 Typical effects of temperature and moisture on the
mechanical behavior of an unmodified cold-curing epoxy
adhesive in tension (from Hutchinson, 1997)

3-15


|

>
Single fap
x SSS i
Single lap
SF Se ee
Toggte lap
_——_—
>
~
Double lap

= SS

Beveled double lap

>

~< = —_
Double lap

<— 1

_—— Wn
Double fap

Multiple step jap

Ss -

Tongue and groove

~ Cee

Simple scarf

eae

Simple scarf

<
<—_
Double scarf

<1

Double scarf

eo single lap

Corner single lap

Corner double lap

4

intersecting wall lap

Figure 3.16 Bonded shear joint concepts
(from Hutchinson, 1997).


Chapter 2
Bolted Connections

2.1 INTRODUCTION

The use of mechanical fasteners to join FRP members is a logical extension of their use in
joining metallic members. Early on, there was an excessive reliance on design methods
originally developed for isotropic materials, as well as a reluctance to punch holes in laminates
for fear of cutting the fibers. Researchers have since developed design methods more
appropriate to FRP and showed that, although mechanically fastened FRP joints share the same
basic failure modes with metals, the mechanisms by which damage initiates and propagates can
be fundamentally different and so classical metal failure criteria are not always applicable.

FRP components can indeed be considerably weakened by holes and cut-outs; this can be
attributed to the large stress concentration at such discontinuities and to a lack of plasticity. In an
infinite, uni-directionally reinforced sheet, the stress concentration at a circular hole can be as
large as 8, compared with a value of 3 for isotropic materials (Collings, 1987). Most isotropic
materials exhibit some plasticity, which relieves the stress concentration and causes it to have
only a small effect on the net failure stress. This is not the case for uni-directional FRP, which is
elastic to failure, and where stress concentration causes low net failure stress. The various
modes of failure, the design parameters affecting joint performance, and the effect of
environmental conditions will be discussed in this brief overview, which mostly covers bolted
joints.

2.2. CHOICE OF FASTENERS

Self-tapping screws are perhaps the simplest form of fasteners. They are used where there is no
access to the reverse side of a joint, but are inefficient due to the possibility of stripping of the
threads cut in the FRP laminates.

Rivets are suitable for joining laminates up to 3 mm (0.1 in) thick. The closing pressure is not
always readily controllable, resulting in a wide variation in clamping pressure. In addition, the
riveting operation can potentially damage the laminates.

Collings (1977) showed that bolted joints are the most efficient form of all mechanical fastening
for carbon FRP (CFRP). On a specific strength basis (strength/mass), they are even superior to
conventional structural materials.

2.3. FATLURE MODES

In addition to failing in modes similar to those found in metallic connections, i.e., by shear,
tension or bearing, FRP connections may also fail by cleavage of the laminate or the connector

pulling through the laminate (Fig. 2.1).
2-1


Threaded
Rod

|
Threaded 4 “Universal Connector”
Rod | i

Beam Nl

Column

Figure 4.10 Use of threaded rod stiffeners

(from Mosallam, 1994)

ie

=
co
Yo

Figure 4.11

Thickening and stiffening of flanges of column

and beam using leg angles (from Mosallam, 1994)

4-9


Pips to control
bondline thickness.

Toggle connector

4
Adhesive

Figure 5.5 Pips or ridges molded into surface of composite to
control bondline thickness (from Head and Churchman, 1989)

Figure 5.6 Advanced composite site office building, Second Severn Crossing,
Avon, UK. (from Head and Churchman, 1989)

5-5


2.3.3. Bearing Failure

Bearing causes compression on the loaded half of the bolt hole. Thus the compressive strength
of the 0° fibers, as well as the clamping pressure are important parameters in determining bearing
strength (Collins, 1977, 1982; Kretsis and Matthews, 1985; Stockdale and Matthews, 1976;
Matthews and Kalkanis, 1985). An average bearing stress assumed to act uniformly on the
cross-sectional area of the hole can be calculated,
P
o=- 7 (2-3)

Working with CFRP, Collings (1982) showed that failure initiated in shear at the hole edge,
through both fibers and matrix. Laminates containing some 90° and / or + 45° plies perform well
under bolt bearing conditions. This is contrary to what might be expected from the compressive
behavior of 90° and + 45° laminates, thus demonstrating that bearing failure mechanisms are
different from those of compressive failure. The difference is due to the lateral pressure exerted
by the clamping action of the bolt. Constraint against Poisson expansion normal to the plane of
the laminates causes the fibers transverse to the load to fail in constrained transverse
compression.

Kretsis and Matthews (1985) suggested that the low elastic modulus of glass fibers favors an
instability mechanism that is not fully prevented by lateral constraint. Matthews and Kalkanis
(1985) showed that the fibrillar nature of Kevlar fibers causes weak planes of failure. Bearing
strength thus depends more on this weak bonding between "fibrils" than on the interface between
fibers and matrix.

2.3.4 Cleavage Failure

Cleavage failure only occurs in 0 ° / + @ lay-ups with a high proportion of 0 ° fibers (@ is the
orientation of the plies that are not at 0 °). Failure initiates in a single shear mode followed by
failure of the net section on one side of the laminate. If such lay-ups are needed, the region
around the hole should be reinforced.

2.3.5  Pull-out Failure

Pull-out failure is associated mostly with rivets. Since rivets are used mostly in single shear,
axial in-plane loading imposes out-of-plane bending and peeling (pulling apart) of the joint.
Under extreme loading, rivets may fail by bending or head shearing.

Of the above modes of failure, the most desirable are bearing and tension (Fig. 2.4).

2.4 IMPORTANT FACTORS AFFECTING JOINT STRENGTH

2.4.1 Fiber Orientation

As mentioned in the discussion on modes of failure, fiber orientation has a strong effect on joint
strength. Collings (1977) showed that the best overall joint performance for CFRP is obtained

2-3


NISTIR 6532

Connections of Fiber-Reinforced
Polymer (FRP) Structural Members:
A Review of the State of the Art

Dat Duthinh

August 2000

Structures Division

Building and Fire Research Laboratory
National Institute of Standards and Technology
Gaithersburg, MD 20899

U.S. Department of Commerce

Norman Y. Mineta, Secretary

Technology Administration

Dr. Cheryl L. Shavers, Under Secretary of Commerce for Technology
National Institute of Standards and Technology

Raymond G. Kammer, Director

a


(a)

{b)

Figure 3.2 Exaggerated deformations in loaded single lap joint:
(a) with rigid adherends; (b) with elastic adherends.
(from Adams, 1987)

Pe r — _—— Bending moment Mo =Pt [2
= P

{a) Undetormed joint

Bending moment My =KPt

K<1

(b} Detormed joint

Figure 3.3 A geometrical representation of the Goland and
Reissner (1944) bending moment factor (¢ = adherent thickness).

3-6


Table 2.1
Lay-Up Number and Stacking Sequence
from Garbo and Ogonowski, 1981

Ply number Ply orientation (degrees)?
to centreline? Lay-up number
1 2 3 4 5
1 + + + + +
2 0 - - - -
3 - (0) + 1) i}
4 QO 0 - ce) 0
5 90 + 90 90 0
6 0 90 1) oO i)
7 + - 0 + oO
8 0 0 0 - +
9 - QO i) 0 -
10 0 i) i) oO 90
Number of plies 20 20 20 20 20
Percentage of
0/+45/90° 50/40/10 50/40/10 50/40/10 50/40/10 50/40/10

*4 and - refer to +45° ply orientation.
>Lay-ups are all symmetric about the centreline.

13805 —

3(see table 2)

Bearing strength (MN/m?)

— T T 7
100
Percent of 0° plies grouped together

Figure 2.9 Effect of grouped 0° plies on bearing strength.
See Table 2.1 for lay-up (from Garbo and Ogonowski, 1981)

2-12





end and column faces to allow rotation. The EXTREN Design Manual (1989), published by
Strongwell Corporation for the design of glass FRP structural shapes, assumes the use of FRP
bolts at least 10 mm (0.4 in) in diameter, in holes with clearance of 1 mm to 2 mm (0.04 in to
0.08 in), and tightened up to a torque of 24 N-m (210 Ibf-in).

4.4 SEMI-RIGID CONNECTIONS

Figure 4.3 shows a typical failure of a bolted FRP beam-column connection, in which angle
sections are used to join the flanges of a beam and a column. The FRP angle fails by formation
of hinges due to opening (top) or closing (bottom) moments. Other likely locations of failure are
under the bolts going through the column (crushing) and at the junction of web and flange of
FRP I-beams, an area relatively rich in resin and poor in fibers (Fig. 4.4). The simplest solution
is to replace the FRP angles with stainless steel ones, if weight is not a concern.

Mottram and Zheng (1997) improved on the basic design of semi-rigid connections by:

e Increasing the stiffness of the connecting angles by doubling their thickness or adding more
angles (Fig. 4.5).

e  Stiffening the flanges of beam and column with bolts (Fig. 4.6).

These FRP semi-rigid connections are stiffer, stronger and rotate further before failure than steel

connections (Fig. 4.7). The combination of bolting and bonding improves performance over

bolting or bonding alone (Fig. 4.8).

Mosallam (1994) increased the stiffness of the connecting angle by adding triangular gusset
plates. This so-called “universal connector” is especially made with molded glass / vinyl ester
but is not yet a readily available item (Fig. 4.9). The use of over-stiff connectors transfers failure
to the FRP beam and column, so they must be stiffened as well (Figs. 4.10 and 4.11).
Alternatively threaded rods may also be used in conjunction with the “universal connector”. One
criticism of the tests performed by Banks, Mosallam and McCoy (1994) is that their test set-up
produces both bending and compression in the members (Fig. 4.12). This compression is usually
not present in the beam of actual frames and may explain some of the failure modes observed.

42


3.3. FAILURE MODES

Adhesively bonded joints can fail in three distinct modes (Hart-Smith, 1987):

© The strongest joints do not fail in the adhesive at all. Rather, they fail outside the joint area
at 100 % of the adherent tensile strength. The strength of the joint is thus proportional to the
laminate thickness t. However, delamination of fibrous composite adherents is a
considerably weaker mode of failure.

e The next highest strength is limited by the shear strength of the adhesive. The strength of the
joint is here proportional to the square root of the laminate thickness.

e The weakest failure mode is associated with failure of the adhesive under peel stresses. The
peel strength is proportional to the quarter power of the laminate thickness.

e In addition, failure of the interface between adhesive and adherents is a possibility.

Adhesives are most efficient only in the thickness range of 0.125 mm to 0.25 mm (0.005 in to

0.010 in). Figures 3.5 and 3.6 show the failure modes and strength envelopes of adhesively

bonded single-lap joints. Depending on the coefficient of proportionality between the various

strengths and the powers of t, the strength envelope can be more or less complicated. Fig. 3.7

shows the peel stress failure of thick composite joints.

3.4 DESIGN IMPLICATIONS

Due to the non-uniform shear stress distribution in the adhesive, extending a lap joint beyond a
certain length does not increase its strength, but does provide safety against creep rupture,
provided the stress in the central part of the joint is sufficiently low (Figs. 3.8 and 3.9). The
stress trough in the middle of a lap joint should thus not be regarded as an inefficiency to be
designed away. In fact, Hart-Smith (1982) considers that limiting the minimum strain (to 10 % of
the maximum at ultimate load) in a lap joint is more important than limiting the maximum strain.

A symmetrical design, such as a double lap joint, alleviates bending effects to some extent, but
load eccentricity and bending are still present, thus leading to tensile peeling stresses (Fig. 3.10).
The addition of a fillet at the end of the lap is very effective in reducing stress concentration and
increasing joint strength (Fig. 3.11). Other designs, such as tapering the end of the adherents,
accompanied by local thickening of the adhesive, reduce stress concentration for both shear and
peel stresses, but do not make as significant an improvement in joint strength as a well designed
fillet does (Fig. 3.12).

A stepped lap joint is an attempt to render the shear stress distribution more uniform and allows
an efficient use of a long lap (Fig. 3.13). Each step, in effect, functions as an individual lap joint,
with the butt faces at the end of each step transferring a significant portion of the load. An
important design consideration is the need to balance adherent stiffness at both ends of the joint.
The joint on the right in Fig. 3.14 is almost twice as strong as the one on the left, yet uses less
material!

Another particularly efficient design is a scarf joint, which causes stresses to be much higher at
the compression end (where the center adherent ends) than at the tension end of the joint (where
the outer adherents end). Tapering of the adherents provides little advantage unless the scarf
joint ends in a knife-edge, which is difficult to achieve in practice. A tip thickness of only 0.025

3-2


Failure load
25mm wee. Joint)

Simpl lt
(a) Simple lap joint Fracture

if composite

_ 3
==. _—. 223
(b) Scarfing of steel
- 23-5
+ = —.
(c) Reverse tapering of steel
3O* angie of
fillet
- 76-3
i. ad —e

Adhesive failure
(d) Reverse tapering of steel and adhesive fillet

Figure 3.12 Bonding of steel to CFRP (length in mm) (from Adams and Wake,
1984). Fracture of composite caused by through-thickness tensile stresses.

:

CFRP (unidirectional) ° _
f—a—
a es onal fe

= $23 Be" | ceRP-crRP acnerends
TET a: 051rm, aluminum-CFRP aanerenas

Shear stress
‘Average applied shear stress

°
ma

6
Distance alona overiap (mm)

Figure 3.13 Adhesive shear stress distributions in
step joints (from Adams and Wake, 1984).

3-11


B0or~ © 635mm ¢
0 953mm ¢
4 1270mm 6
700}—
600}-
Z 500--
2
5 Glass (Ref, §) 0° £45*
2 400--
g
2 Carbon (Ref. 1) O*45*
5 or
8
200- Keviar (Ref.11) 0*/90*
100/—
L 1 ie L A A.
° 1 2 4: 5 6

3
dit

Figure 2.13 Variation of bearing strength with d/t for a laminate without
lateral constraint (from Collings, 1987)

2-15


Figure 1.3 20 m (66 ft) high glass FRP (GFRP)
stair tower in Virginia, USA (from Hutchinson, 1997).

Figure 1.4 Structural framework and bolted connection detail in a
chlorine plant (from Hutchinson, 1997).


Figure 5.9 Composite transmission tower with SNAP joints
(Designers: W.B. Goldsworthy & Associates, Inc. 1994 for Ebert Composites)
(from Goldsworthy and Hiel, 1998)

Figure 5.10 Test setup for truss joint
(at Coordinated Equipment Index Testing Facility, Wilmington, CA)
(from Goldsworthy and Hiel; 1998)

5-7


45 REFERENCES
Banks, L.C., Mosallam, A. and McCoy, G.T. (1994), “Design and Performance of Connections

for Pultruded Frame Structures,” J. Reinforced Plastics and Composites, vol. 13, pp. 199-212,
Mar.

Bass, A.J. and Mottram, J.T. (1994), “Behavior of Connections in Frame Reinforced Polymer
Comopsite Section,” Structural Engineer, Sep., vol. 72, no. 17, pp. 280-285.

Clarke, J.L. (1996), “Structural Design of Polymer Composites,” EUROCOMP Design Code and
Handbook, E&FN SPON.

Hutchinson, A.R. (1997), “Joining of Fibre-reinforced Polymer Composite Materials,” CIRIA
Report 46, Nov.

Mosallam, A.J. (1994), “Connections for Pultruded Composites: a Review and Evaluation,”
Proc. 3" Materials Engg.. Conf., ASCE, San Diego, CA, Nov., pp. 1001-1017.

Mottram, J.T. and Zheng, Y. (1996), “State of the Art Review on the Design of Beam to Column
Connections for Pultruded Frames,” Composite Structures, vol. 35, pp. 387-401.

Strongwell Corp. (1998), “EXTREN Design Manual,” Bristol, Va.

4-3


Composite failure
n in L-shaped cleat

Figure 4.3 Failure of composite
cleat. (from Hutchinson, 1997)

Figure 4.4 Resin-rich zone at flange/web junction of
pultrusion (from Hutchinson, 1997)


2.3.1 Tension

The average net stress 6, across a section is an obvious consideration in the design of a joint:
P
o.=———— 2-1
n* Goond)t @)
where P is the tension carried by the joint of width w and thickness ¢ with n bolt holes of
diameter d within the section.

However, stress concentration often causes failure to initiate at the hole. Unlike metals, FRP are
linearly elastic up to failure, and consequently no yielding alleviates stress concentrations.
However, Hart-Smith (1980) noted that slippage between resin and fibers provides some relief of
stress concentration induced around bolt holes and cut-outs (Figs. 2.2 and 2.3). The most highly
loaded fibers pull out of the resin over a short length to alleviate high fiber strains, i.e., local
debonding or delamination of fibers from resin allows local load sharing between fibers. Thus
slippage between resin and fibers averages out local stress concentrations and results in more of
the fibers being loaded before final failure. This leads to the surprising conclusion that, in the
vicinity of holes, improving adhesion between resin and fibers can in fact lower joint strength.
However, reinforcing the vicinity of holes with fibers in different directions reduces the degree
of anisotropy, and introduces some pseudo-plastic behavior or softening (Hart-Smith, 1980).

Tensile strength depe*s highly on fiber orientation. The fibers parallel to the load (0°) carry
most of it, and for CFRP, failure is initiated at the stress concentration at the edge of the hole at
90° to the loading axis (Potter, 1978). For glass FRP (GFRP), failure is a more complex
combination of shear and tension failures. Failure initiates with the 0° fibers shearing over the
projected area of the hole and propagates by in-plane shearing between the 0° and 45° plies
across the entire width of the laminate. As a consequence, only the 45° plies remain to carry the
axial load, and tension failure occurs immediately after in-plane shearing (Godwin, Matthews
and Kilty, 1982).

Aramid fibers (Kevlar) exhibit low lateral adhesion (compared with its high tensile strength) and
tend to split along the fiber axis (fibrillation). This relieves the stress concentration at the hole
edge and contributes to the tensile strength of the connection (Collings, 1987).

2.3.2 Shear Failure

The shear strength t of an FRP joint of thickness ¢ carrying a force P is calculated in the same

manner as for isotropic materials,
P
T= te (2-2)

where e is the distance from the center of the bolt hole to the ends of the connected plates. As
mentioned above for tension failure, fiber orientation is also critical in determining shear
strength. For 0° fibers, the shear strength of a joint is low compared with the in-plane shear
strength of the laminate, thus indicating high shear stress concentration around the hole. In
contrast, the shear strength of 0°/ 45° joints is high and insensitive to end distance, suggesting

low shear stress concentration.
2-2