A study of the failure behaviour of key-lock joints in glass fibre reinforced plastic pipework

C’onrjx~.v//c~; PCII.~,4 27A (1996) 429.. 135 Copyright (’ 1996 Elsevier Science Limlted Printed in Great Britain. All rights reserved 1359~835X/961$15.00

1359-835X(95)00071-2

A study of the failure behaviour of key-lock joints in glass fibre reinforced plastic pipework

A. Fahrer”

and A. G. Gibson

Centre for Composite Materials Engineering, Newcastle-uponTyne, NE7 7RU UK

University

of Newcastle,

and P. Tolhoek Ameron Fiberglass Pipe Division, (Received 8 August 7995)

4790 CA Geldermalsen,

The Netherlands

load tests on keyylock joints showed a linear correlation between the internal failure pressure and axial tensile or flexural loads. In all tests the joint consistently failed at a location near to the bottom of the groove in the male part, which made it possible to model the failure in terms of load per unit length of crack circumference. A PC-based finite element package was used to model the local stress distribution around the joint. The Tsai-Wu and the maximum stress criteria accurately predicted the location of the initial damage. The pressure corresponding to predicted first damage was approximately the same as that at which the first audible cracks were heard, which was considerably lower than the final failure pressure. The critical stage of joint failure was observed to involve an interesting practical example of mode II fracture. For the particular geometry in question the predicted fracture load is independent of crack length, suggesting that the mode II crack propagated under approximately constant load. Values of Gnc from the pipe tests agreed well with those for glass/epoxy laminates found in the literature. Combined

(Keywords: glass/epoxy

pipes; key-lock

joint; failure bebaviour)

INTRODUCTION The high specific strength and corrosion resistance of filament wound glass reinforced epoxy (GRE) pipework has led to its increasing use in the offshore oil and chemical industries to transport a variety of fluids. One special offshore application is in firewater pipework’, where GRE pipes have been shown to offer a safe and reliable alternative to steel and other metals. Other applications 2-5 include chemical process piping, temporary pipelines and pipe systems for corrosive liquids. GRE pipework offers the choice of two joining methods: one based on adhesive bonding, the other on a mechanical joint, known as the key-lock. The adhesive joint, as shown in Figure 1, has either straight or tapered mating pieces, which are bonded with a thermosetting

*To whom correspondence

should

be addressed

epoxy adhesive. This is the strongest form of connection. Once the adhesive is cured however, the joint cannot be taken apart again, which restricts the range of use of this type of joint. The principle of the key-lock joint is shown in Figure 2. The axial load due to internal pressure is supported by a cylindrical nylon key, which can be inserted into preformed grooves in the male and female pipe ends. The hydraulic seal is achieved by means of a separate O-ring. The joint is assembled by inserting the nylon key into the matching grooves by passing it through a tangential hole in the female pipe end. The joint can be dismantled by removing the key. Although the strength of this mechanical joint is less than that of an adhesive pipe joint, the system is quicker to assemble and can also be taken apart. This type of joint is widely used in (temporary) pipelines and can operate underground under high pressures, up to 60 bar working pressure for a 209 mm (8 in) diameter pipe, and elevated temperatures, up to 95°C.

429

Failure of key-lock

joints in GRE pipework:

Female end

Adhesive bond \ \

I

A. Fahrer et al.

Male end

I

ti

1

R Figure 1 Schematic of an adhesively bonded pipe joint. This type of joint is commonly bonded with a thermosetting adhesive and is the strongest form of connection. However, once the adhesive is cured the joint cannot be taken apart again, which restricts its range of use

Figure 4 Experimental arrangement for combined pressure and axial or flexural load testing. In all tests the external load was maintained, while the internal pressure was increased linearly, until the joint failed. In the flexural tests, the joint was placed in such a way that the entry point of the key was located at the neutral axis of the joint Figure 2 Schematic of the key-lock joint. Axial load is supported by means of a cylindrical nylon key, which is inserted into preformed grooves in the male and female pipe end. The hydraulic seal is achieved by means of a separate rubber O-ring

100

0 Applied

Axial Tensile

Load [kN]

Figure 5 Failure envelope showing a strong linear correlation between failure pressure and applied axial load. A similar linear correlation appeared to apply for the pressure at which the first audible damage was heard

i water SUPPlY

Figure 3 End fitting for combined load tests on 209mm key-lock joints. The fittings are based on an adhesive bond between the tapered GRE pipe ends and aluminium fixtures

EXPERIMENTAL Installed pipe systems are often subject to axial or flexural loads in addition to internal pressure. These loads can be caused by, for example, misalignments in the system, thermal fluctuations or self-weight. Such

430

loads can of course influence the failure pressure of the pipe system and the joints in particular. Pipe joints (209mm inside diameter) were tested under internal pressure in combinations with axial tensile or flexural loads. The end fittings, shown in Figure 3, are based on an adhesive bond between the tapered GRE pipe ends and fixtures. All tests were performed by bonding the fittings to the pipe ends, applying the external tensile or flexural load, and recording the joint failure pressure. The external loads, shown in Figure 4, were kept constant during pressurizing. In the flexural tests the joint was placed in such a way that the entry point of the key end was positioned at the neutral axis of the joint, thereby ensuring that the bending stresses on the discontinuity of the key were minimal.

RESULTS During the tests audible acoustic events were occasionally heard, which indicated material damage. Joint failure was usually accompanied by a loud cracking noise. Results for the combined pressure and tensile load are shown in Figure 5, and for the combined pressure and flexural load in Figure 6. The results showed a good

Failure of key-lock

0

2

4 6 8 Applied Bending Moment [kNm]

10

12

Figure 6 Failure envelope showing a strong linear correlation between failure pressure and applied bending moment. A similar correlation appeared to apply for the pressure at which the first audible damage was heard

+70° covers P-. _

254’

\

Hoop wound covers

Interply

crack

Crack ‘- Cracks in liner

7 Datnagc and failure sequence. Damage was observed to initiate in both grooves, but only propagated axially in the male component. Furthermore. cracks in the linear material were observed

linear correlation between the failure pressure and the externally applied load or moment. When the best-fit line in Figure 5 is extrapolated to zero pressure conditions the axial failure load equals 340 kN, which corresponds with an internal pressure of approximately 80 bar. As this is nearly the final failure pressure of the joint it can be concluded that the hoop stress, induced by the internal pressure, does not greatly influence joint failure. The pressure at which the first audible damage was noted is shown in Figures 5 and 6. These results also showed a linear correlation between pressure and external loading.

AND

DAMAGE

As sketched in Figure 7, failure appeared to be initiated from a location near to the root of the male key-groove and damage progressed in the spreading of a circumferential crack. As the winding angle in this part of the joint was 1t70”, the fibres offered little resistance to this propagation. When the crack reached the interface between the different orientated covers an interply crack was observed to propagate along the pipe in the axial direction. A possible explanation is that the change in winding angle obstructed crack growth in its initial direction and forced it to propagate in the axial direction as an interply crack. Circumferential cracks around the key-groove were also found in the female component. but were not observed to grow in the axial direction as in the male component. Explanations for this are the lower load, caused by the slightly larger diameter. and the wavy pattern of the covers, caused by pressing the groove into the material.

THEORETICAL

Figure

FAILURE

A. Fahrer et al.

covers

Crack

Liner

joints in GRE pipework:

ANALYSIS

between internal Considering the linear correlation failure pressure and externally applied loads, and the observed failure mode, it was felt that failure could be expressed in terms of axial load per unit circumferential length, it’, of the crack in the male pipe end. In most IV will be influenced by the axial circumstances, component of the internal pressure in the pipe. There will, however, be additional components due to any externally applied loads, so the overall load per unit length can be expressed as the summation I(’= IVP+ )I’[ + H’l‘,nar

(‘)

where tcp, M’,and ~~~~~~are I(’due to internal pressure, axial tensile loading and the maximum component from flexural loading, respectively. The circumferential length in question will be 27rR,, crack where R, is the radius at which the interply propagates. The overall axial component of pressure loading will be pr(R + II)‘, in which p is the internal pressure, R is the inner pipe diameter and k is the wall thickness of the male joining part. This will be assumed to act over the crack circumference ~TR,. giving

SEQUENCE

Sectioning the joint after testing showed a consistently repeatable failure mode involving the male component, as shown in Figure 7. The male component is wound on a mandrel using 1t54’ and ~t70” orientated covers (measured from the pipe axis) and the key and sealing grooves are machined into the material. The female component is wound with identical orientations and additional 90 ’covers. However, unlike the male component, the groove is pressed into the material during the winding operation.

Similarly. give

an axial tensile load F, acting on the joint will

ll’,

=

__

4

2TR,

The engineering theory of bending (ETB) was used to derive an expression for lrrmax. The flexural loading on a key-lock joint, as shown in Figure 8, was assumed to act as a line load around the circumference of the interply crack at radius R, and vary linearly with distance ~1from the neutral axis. Hence, lt’f can be expressed as

431

Failure of key-lock

joints

A. Fahrer et al.

in GRE pipework:

e

Figure 8 Bending stresses applied on the key-lock joint. The flexural load is assumed to act as a line load around the circumference of the interply crack at radius R, and vary linearly with J

IQ0

Case I

F___+J

Elastic condition (I), partial load redistribution (II) and total load redistribution (III) conditions. The actual stress distribution is thought to be partially redistributed over the crack circumference, as in case II

Figure 10

STRESS REDISTRIBUTION 0

Ii’.,

.I..’

50

Ii’,

250

300

Failure envelope comparing axial load test results [equation (3)] with elastic [equation (7)] and total load redistribution [equation (9)] conditions. The total load redistribution approach shows a better fit with the experimental results

Figure 9

where wfmax is the maximum load per unit circumferential length (which occurs at y = R,). Although the load due to flexure varies around the circumference of the crack, failure of the joint will be assumed to correspond to the maximum tensile value, wfmax. Making use of y = R, sin B and dc = R, do, the bending moment on a cross-section at the distance y from the neutral axis, dM, can be written as dA4 = wfy dc = wfmaxsin 0 R, sin 0 R, d0 = wfmaxR%sin* 8 de

(5)

The total bending moment M over the crack circumference can be found by integrating equation (5) over the circumference, as M = 2wfmaxR;

s *I*

Later in this paper it will be argued that failure occurs, driven by axial load, through the propagation of an axial crack in the male joining part. This crack spreads initially in the circumferential direction and then axially. It will be argued that, under the condition of mode II crack propagation (which applies here), the crack, between the *54” and f70” covers, will propagate under approximately constant load. The effect of this will be to redistribute the local loading around the joint over a much larger distance than would be the case for purely elastic bending. In some respects, this load redistribution can be considered to be analogous to the elastic-plastic behaviour that would occur in a metallic component under flexure. Case III in Figure IO shows a condition that would be analogous to the fully plastic bending situation. Although this case is not thought to occur in the GRE pipes (case II being nearer to the real situation), it represents one extreme of behaviour and is simpler to describe than case II, so it will be used here. If wr were assumed to be constant over the crack circumference, as in case III, this would give wf =

Twfmax

Wfmax= 4R,2 (6)

Rearranging equation (6), wrmaxcan be found as M Wfmax = -

TR,2

(7)

where M is the bending moment on the key-lock joint. Substituting equations (2), (3) and (7) in equation (1) the failure envelopes for initial pressure versus axial load and internal pressure versus flexural load can be combined on one plot of internal pressure versus w, as

432

(8)

M

sin* 0 do

R2c

M’fmax

Using the same approach as previously, this leads to

-*/2

=

APPROACH

“‘-1

~i~~~~i

150 200 100 Unit axial load [kN/m]

Case III

shown in Figure 9. In this figure it can be seen that, for an identical w, the samples tested under flexure fail at slightly higher pressures than those under tensile load. A possible explanation for this is that crack propagation around the circumference of the pipe allows a significant redistribution of stress, in a way that no longer corresponds with the ETB. Notwithstanding this, equation (7) could be used for design purposes, as it would give a positive safety margin.

I

Ots..

Case II

It can be seen that equation (9) is similar to equation (7) as the dominator merely increases from rr to 4. Figure 9 shows the failure pressure versus w under stress redistribution conditions. This approach appears to fit the experimental data. FINITE ELEMENT BEHAVIOUR

(FE) MODELLING

OF JOINT

The NISA PC-based finite element package was used to

Failure of key-lock

joints in GRE pipework:

A. Fahrer et al.

Table 2 Failure properties for unidirectional laminated ply of E-glass fibre reinforcement in an epoxy matrix with a fibre volume fraction of 0.60 Tensile strength in the I-direction, (or,, (MPa) 1000 Compressive strength in the l-direction, CT,,~(MPa) 600 Tensile strength in the 2-direction, (T~,~(MPa) 30 Compressive strength in the 2-direction, g? u (MPa) 110 Tensile strength in the 3-direction. 03 t (MPa) 30 Compressive strength in the 3-direction, qc (MPa) 110 Shear strength in the 12-plane, ~t?,~ (MPa) 30 Shear strength in the 13-plane, ~,s,~ (MPa) 30 Shear strength in the 23-plane. IT??,, (MPa) 30

Matrix Figure 11 Unidirectional ply coordinates. For modelling purposes. long fibre composites are often thought to be built up of unidirectional plies. In these plies all fibres are assumed to be aligned in the l-direction and perfectly bonded with a plastic matrix. Usually, the I-. 2- and 3directions are called the fibre, the transverse fibre and the throughthickness direction respectively

Table 1 Unidirectional Halpin Tsai equations

elastic

ply constants,

E, = 39.06GPa G‘,z = 2.00 GPa I‘,’ = 0.31

& = 5.52 GPa G,, = 2.00 GPa 1’13= 0.31

estimated

with

the

Ei = 5.52GPa C&s = 2.12GPa l’?j = 0.30

model the stresses in a key-lock joint under pure pressure conditions. Although such an analysis will not be able to predict total joint failure, being a case of crack propagation, it is possible to predict damage initiation in the key-groove. In the discussion of this analysis, all stresses are expressed in the unidirectional laminated ply coordinate system shown in Figure Il. The key-lock joint geometry was modelled using design drawings supplied by Ameron, and the ply arrangements and fibre orientations according to optical microscopy results. Micromechanics were used to calculate the unidirectional elastic ply constants from the elastic properties of the matrix and fibres and their volume fractions in the laminate. Application of the HalpinTsai equation&’ resulted in the ply constants the shown in Table 1, which were used to simulate mechanical behaviour of the filament wound composite material. Failure was modelled with the TsaiiWu failure criterion8.9, which is an empirical polynomial criterion that is often used in finite element analysis because of its computational convenience. It defines a failure surface of a fractional form, .f’, and assumes some interaction of stress components. First ply failure under combined loading is assumed whenf‘ reaches unity. By combining all the stress components in a functional form no prediction of a failure mode can be made. The coupling coefficients for the Tsai-Wu failure criterion, accounting for the interaction of stress components, were estimated as suggested elsewhere’. As failure properties for composite materials have been extensively studied in the past, it was felt that these properties could be obtained from the literature”. The suggested unidirectional laminated ply failure properties for E-glass fibre reinforcement in an epoxy matrix with a fibre volume fraction of 0.60 are presented in Table 2.

The PC-based version of the finite element package was limited to 5000 degrees of freedom, which made it necessary to model the male and female component separately and ignore the contact problem between the key and key-groove. Therefore, a sensitivity analysis was carried out using empirical load distributions between key and groove. simulating an operating pressure of 20 bar. The problem was treated as an axisymmetric one and analysed under static conditions. As shown in Figure 12, the effect of the different empirical load distributions on the predicted ply stresses appeared to be small and the model was observed to be insensitive to the way in which the stresses were applied to the key-groove. The highest stresses were found at 0.7 to 0.8 times the depth of the key-groove. At these locations the TsaiiWu criterion predicted conditions close to first ply damage, although the modelled stresses were clearly higher in the male component. Comparing the ply stresses with maximum allowable stresses, as in Table 2, it was found that the first damage conditions were initiated by high (interlaminar) shear and throughthickness stresses. The predicted location of first damage corresponded with the failure initiation point in the male key-groove and the predicted pressure value agreed approximately with the observed acoustic events in the load tests.

FRACTURE

MECHANICS

A common mode of fracture in composite laminates is delamination, which may be regarded as the propagation of an interply crack. A measure of the resistance of the material to delamination is the interlaminar fracture energy, which can be measured using linear elastic fracture mechanics, thus enabling the critical energy release rate Cc to be determined. Unlike homogeneous metals or polymeric materials. the fracture of continuous fibre composites cannot be modelled by a single linear fracture parameter Gc. Three modes of fracture are identified: mode I (tensile opening), mode II (in-plane shear) and mode III (scissor blade). Furthermore, combinations of these fracture modes can occur and each mode has its own critical energy release rate. The most frequently observed modes of fracture arc mode I and mode II, as shown in Figure 13. Mode I is the

433

Failure of key-lock

joints in GRE pipework:

Stress Assumed

A. Fahrer et al.

distribution

in

the

male

key-groove

loading

CJ23 TSAI-WU

CT

Y--O

.2 8 y=R

Case

y=R

1

023

y=R

TSAI-WU

y=R

Case

2

All

stresses

in

MPa

Figure 12 Stress distribution

and Tsai-Wu failure parameter in the male key-groove for two types of assumed load distribution: case I parabolic and case 2 linear. The following seven diagrams show the stresses in the principal material directions and the Tsai-Wu failure parameter, for a modelled pressure of 20 bar. The stress distributions around the groove were found to be similar for both cases. Failure of the male component is predicted to occur at 0.7 to 0.8 times the depth of the groove

Mode I

Mode II

Figure 13 Mode

I (tensile opening) and mode II (in-plane shear) failure. Characteristic of each mode is the direction in which the crack c propagates with respect to the applied forces

lowest fracture energy for isotropic materials and, thus, a crack will always propagate along a path normal to the direction of the maximum principal tensile stress. Hence, a crack in an isotropic plate will propagate in mode I fracture, regardless of the orientation of the initial flaw with respect to the applied stress. However, this is not necessarily the case for crack growth in fibre composites, which are inhomogeneous and highly anisotropic materials. In these materials the initial interlaminar defect is constrained and often continues to propagate in the same plane between the laminae regardless of the orientation of the crack to the applied loads. Thus in these materials mode II failure is possible and, obviously, mixed mode I/II may also be observed. In the combined load tests final failure consistently occurred in the male part of the joint. Although only the final failure event was observed to propagate in mode II, it was felt that the overall joint failure could be modelled in terms of mode II fracture. Considering a mode II

434

-L--i F

___l?_/_/F

Figure 14 Schematic

representation of mode II fracture in the male pipe end in a cross-section of the key-lock joint. The crack length is represented with c, the inner pipe radius with R and the wall thickness of the male pipe end with h

crack propagating in a body, the energy release rate Gnc can be derived from

(10) in which U is the total elastic energy in the body, and B and c are the width and the crack length respectively. Figure 14 shows a sketch of the mode II failure in the male component. In this figure h is the wall thickness of the male part, R is the inner pipe radius and Yis the depth of the key-groove. By applying an axial load F to the joint, the crack length c increases with a length dc. The change in total body energy dU due to crack growth from c to (c + dc) can be expressed as (11) in which E is the modulus of the joint in the direction of

Failure of key-lock

crack propagation,

and A,, Ab, ga and gb are given by

A, = nR,2 - rR2

F 0, = A,

Ab = T(R + h)’ - nR2

F a,, = Ab

(12)

This results” in 1

An important feature to note in equation (13) is that the driving force for crack propagation, F, is independent of crack length. This implies that the mode II crack will propagate under approximately constant load. By substituting equations (2), (3) and (9) in equation (I), F can be rewritten as F=F,+plr(R,+h)‘+g

C

(14)

Numerical values for Gnc were calculated from the combined load tests as 1.79 f 0.21 kJme2, taking R, R, and h equal to 104.5mm, 109Smm and 12Smm, respectively. E was taken as the axial pipe modulus, which was 11.OGPa. In the liferature’2p’4, energy release rates for unidirectional fibre-reinforced laminates were found to vary from 0.65 to 2.05 kJmp2. Although the numerical values calculated here are for filament-wound specimens with *54” and *70” oriented covers, they are consistent with the literature. On the assumption that Gnc is constant, an expression for the axial failure load of key-lock joints can be found by rewriting equation (13) as

in which Q is a constant geometry.

depending

(15)

ACKNOWLEDGEMENTS This work was carried out as part fulfilment of the Degree of Master of Science at the University of Twente, The Netherlands. The authors wish to thank Professor P. Powell, Group of Engineering Design in Plastics, Department of Mechanical Engineering, for his help and guidance.

REFERENCES 1

3

on the key-lock 4

5

CONCLUSIONS 6

1) Combined

load tests on key-lock joints with a 209mm inner pipe diameter showed that there is a linear correlation between failure pressure and applied axial load or bending moment, and failure consistently occurred in the male component. 2) Damage in the male joint was observed to initiate at a location near to the root of the key-groove. This damage appeared to propagate through the top f70” oriented covers and changed to an interply crack when it reached the f54” covers, causing the joint to fail. 3) Simple mechanical theory, based on the engineering theory of bending, underestimated the internal pressure in combined pressure and bending moment conditions. A different approach, allowing for stress redistribution, showed better agreement with the

A. Fahrer et al.

experimental results. This indicated that the failure pressure of a key-lock joint under combined loading conditions is a linear function of the internal pressure, the axial load and the bending moment. 4) Linear elastic fracture mechanics for mode II resulted in an expression for Gnc that was independent of crack length. This indicates that the interply crack, the final failure event, propagates under approximately constant load. The obtained numerical values for Gnc were consistent with values for unidirectional fibrereinforced laminates found in the literature. 5) With a PC-based finite element analysis it was possible to predict the location of onset of damage in the keygrooves, which also agreed with the first audible events. Additionally, damage was predicted to occur first in the key-groove in the male part, which coincided with experimental observations.

2

F=cY@&

joints in GRE pipework:

7 8 9 10 11 12

13

14

Ciaraldi, S., Alkire, J. D. and Huntoon, G. Fibreglass firewater systems for offshore platforms. Paper OTC 6926, ‘Proc. 23rd Annual Offshore Technology Conf.‘, Houston, TX, May 199 1, pp. 477-484 ‘Ameron Bonstrand Reference Guide, FP-123A’, Ameron Fibreglass Pipe Division, Geldermalsen, The Netherlands, 1986 ‘Ameron Bonstrand Reference Guide, FP-311 A’, Ameron Fibreglass Pipe Division, Geldermalsen, The Netherlands, 1986 Williams, J. G. Developments in composite structures for the offshore oil industry. Paper OTC 6579, ‘Proc. 23rd Annual Offshore Technology Conf.‘, Houston, TX, May 1991 Marks, P. J. The fire endurance off glass-reinforced epoxy pipes. Presented at ‘Polymers in Marine Environment Conf.‘, October 1987, Institute of Marine Engineers Hull, D. ‘An Introduction to Composite Materials’, Solid State Science Series, Cambridge University Press, 1986, p. 86,88 and 108 Chou, T. W. ‘Micro Structural Design of Fiber Composites‘. 2nd Edn, Cambridge University Press, 1992 Sih, G. C. and Skudra, A. M. ‘Failure Mechanics of Composites’, Elsevier Science Publishers, Amsterdam, 1985, pp. 96-97 ‘User Manuals NISA’, Engineering Mechanics Research Corporation, Wild & Partners Ltd, Stockport, UK, September 1992 Correspondence with College of Aeronautics Laminate Analysis, Cranfield Institute of Technology, Cranfield, UK Williams, J. G. Personal comments Hashemi, S., Kinloch, A. J. and Williams, J. G. Mechanics and mechanisms of delamination in a poly(ether sulphone)-fibre composite. Compos. Sri. Technol. (1990) 31, 429 Hashemi, S., Kinloch, A. J. and Williams, J. G. The effect of geometry, rate and temperature on mode I, mode II and mixed-mode I/II interlaminar fracture of carbon-fibre! poly(ether-ether ketone) composites. J. Compos. Muter. 1990, 24,918&956 Hashemi, S., Kinloch, A. J. and Williams, J. G. The analysis of interlaminar fracture in uni-axial fibre-polymer composites Proc. Roy. Sot. Lond. 1990, A427, 173

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