Composite Structures 93 (2011) 1538–1548

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Composite Structures

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Compressive strain limits of composite repaired pipelines under combinedloading states

Ahmed Shouman, Farid Taheri ⇑Department of Civil & Resource Engineering, Dalhousie University, 13360 Barrington Street, Halifax, Nova Scotia, Canada B3J 1X1

a r t i c l e i n f o

Article history:Available online 22 December 2010

Keywords:Composite wrapPipesRepairCombined loadingBendingStrain limit

0263-8223/$ - see front matter � 2010 Elsevier Ltd. Adoi:10.1016/j.compstruct.2010.12.001

⇑ Corresponding author. Address: 1360 BarringtonCanada B3J 1Z1. Tel.: +1 902 494 3935; fax: +1 902 4

E-mail address: farid.taheri@dal.ca (F. Taheri).

a b s t r a c t

Recent literature and the industry have widely accepted the benefits of strain-based design in regards tosteel pipelines. In accordance with the recent increase in the use of strain-based design for pipelines, inthis paper, the authors investigate composite repaired pipelines and its applicability in reference to thestrain-based design. In an earlier work, the authors found that under combined loading conditions, arepaired pipe would tend to buckle locally in a location adjacent to the composite repaired wrap. In anattempt to limit the initiation of local buckling response in composite repaired pipelines, the compressivestrain limit was investigated. For that, a finite element study was conducted. The results clearly showedthat the maximum strains would not occur in the post-repaired defect region. The local buckling wouldhowever occur in the unrepaired (undamaged) section of the steel pipelines. An experimental study wasalso conducted on Fibre Reinforced Polymers (FRP) repaired pipelines to verify the finite element results.A parametric computational investigation was also conducted to understand the limits of compositerepair wraps.

� 2010 Elsevier Ltd. All rights reserved.

1. Introduction

Composite repaired pipelines have been proven to be a viablepipeline repair alternative over the past two decades. It has beenshown that it can perform adequately under different environ-ments and industrial projects (Duell et al. [7]). The limiting crite-rion on which the composite repair system has been evaluated isthe bursting strength of the steel pipeline. Moreover, the compos-ite repair systems have been configured of unidirectional fibreswrapped in the circumferential direction, with the thickness ofthe repair wrap established based on simple mechanics equation[2] as shown below:

Pburst ¼rulttpipe þ rult-wraptwrap

rið1Þ

where Pburst is the burst pressure of the pipe, rult is the ultimate ten-sile strength of the steel pipe, tpipe is the thickness of the steel pipe,rult-wrap is the ultimate tensile strength of the composite in the hoopdirection for composite wrap, twrap is the thickness of FRP repairwrap and ri is the internal radius of the pipe.

This repair philosophy have been successfully implemented inmost cases until now on many different projects, but that doesnot deny the fact that pipes may often be subjected to different

ll rights reserved.

Street, Halifax, Nova Scotia,84 6635.

loading states. The study on the influence of various loading condi-tions and their combinations on such repair pipes is quite scares.Lee et al. [12], developed a new resin transfer molding (RTM)process for repairing/reinforcing damaged underground pipesusing fibre reinforced composite materials. They used glass fibrereinforcement covered by tarpaulin films that worked as a flexiblemold and protection skin, and a porous breathing tube was used toremove air entrapped in the reinforcement. To assess the increasein the capacity of the pipe rehabilitated by their reinforcement,they subjected the repaired pipe only to compressive line loads,applied on the top (12 o’clock) and bottom (6 o’clock) and observedapproximately 15% increase in the load capacity of the repairedpipe.

As can be seen, the work in this area is very scares, and there-fore, it is of utter importance to consider the influence of loadsother than the internal pressure on the response of such repairedpipes. The response of repaired pipelines under combined loadingsuch as the internal pressure, axial force and bending was investi-gated by the authors [13]. It was found that in an unrepaired pipe,wrinkling would occur at the defective area. For the repaired pipe,this local buckling was found to have the tendency of moving to anarea outside but adjacent to the composite repair section. This isdue to the added stiffness provided by the composite to the steelsection, thus making it act like a steel collar. The local bucklingphenomenon is associated with large strains. Thus, the strain-based design of steel pipelines would be more appropriate toaccount for the plastic response of Fibre Reinforced Polymer

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(FRP) repaired pipelines. As a result, the strain-based design isinvestigated in this study in reference to composite repairedpipelines.

One of the few notable and very comprehensive studies is thework of Alexander and Ochoa [3], who developed an integratedanalytical and experimental methods to assess the integrity of anonshore composite repair technique that was developed for appli-cation to offshore risers. Their repair technique constituted design-ing a carbon/epoxy based composite system, fitted on a steel pipe.The development of the repair system included the incorporationof computational simulation, prototype fabrication and experi-mental verification. The design philosophy included embodimentof the limit state analysis and strain-based design methods forthe steel/reinforced composite system. The authors further vali-dated the fabrication, installation and the proposed design meth-odology by conducting tensile, burst and bending tests.

1.1. Stage 1: Numerical modelling

1.1.1. Outline of the procedureThe ABAQUS finite element (FE) code was used to conduct the

analysis. ABAQUS’ three dimensional reduced integration, eight-

Table 1Materials’ properties.

Parameter Value

Glass epoxy composite repair wrap material’s properties (from Hyer [10])El (GPa) 55E2 (GPa) 15.2E3 (GPa) 15.2m12 0.254m13 0.254m23 0.428G12 (GPa) 4.7G13 (GPa) 4.7G23 (GPa) 3.28

Steel pipe material’s properties (from Walker and Williams [14])Young’s modulus (GPa) 205Poisson’s ratio 0.3Minimum yield stress (MPa) 413Yield strain 0.5%Ramberg–Osgood’s model yield offset 1.48%n parameter used in Ramberg–Osgood’s model 18.99

Filler material’s properties (from Clock Spring [6])E (GPa) 3.5

Table 2A typical series of considered pipes with their geometric, dimensionless parameters and theas a sample of pipes with D/t = 60; similar parameters were considered for pipes with the

Pipe’s designation Geometric parameter Dimensionless parameters

D (mm) t (mm) D/t Pt/Py rh/ry

D50C45P00 508 10.2 50 0.26 0D50C45P20 508 10.2 50 0.26 0.2D50C45P40 508 10.2 50 0.26 0.4D50C45P60 508 10.2 50 0.26 0.6D50C45P80 508 10.2 50 0.26 0.8

D50C10P00 508 10.2 50 0.06 0D50C10P20 508 10.2 50 0.06 0.2D50C10P40 508 10.2 50 0.06 0.4D50C10P60 508 10.2 50 0.06 0.6D50C10P80 508 10.2 50 0.06 0.8

D50T30P00 508 10.2 50 �0.17 0D50T30P20 508 10.2 50 �0.17 0.2D50T30P40 508 10.2 50 �0.17 0.4D50T30P60 508 10.2 50 �0.17 0.6D50T30P80 508 10.2 50 �0.17 0.8

Note: Pt is the thermal axial load and Py is the axial load causing yielding.rh is the hoop circumferential stress and ry is the yield stress.

node linear solid element (C3D8R) was used to construct the mesh[1]. According to a study done by Fatemi et al. [8], C3D8R elementprovides adequate accuracy in capturing the moment–curvatureresponse in the buckling and post-buckling regime. A mesh conver-gence study was conducted to establish the integrity of the meshand its numerical convergence. A finer mesh was generated inthe repair region and an even denser mesh was generated at thedefect region. A schematic of the FE model is shown in Fig. 4. Notethat for the sake of clarity, only a quarter of the pipe’s FE mesh isillustrated. A uniform and moderately fine mesh was used for theremainder of the steel pipe. A total of 16,320 C3D8R elements,connected by 20,962 nodes, were used to construct the model.Moreover, due to geometric and loading conditions, only a quar-ter-symmetry of the geometry was modeled. A mesh convergencestudy was also carried out to ensure the effectiveness of the mesh.

To apply the loading and constraints, a reference node was lo-cated on the axisymmetric center-line of the pipe, at each of itsends (see Fig. 4). The axial load and bending moment, as well asthe constraints were applied to these reference nodes. Appropriatemulti-point constraints (MPCs) were used to connect the referencenodes to all the nodes on the mesh that were located at pipe’s ends.Some representative MPCs are schematically shown in Fig. 4(again, for the sake of clarity, not all the MPCs are illustrated). Asfor the translational degrees of freedom, the translational degreesof freedom of the reference nodes were restrained, except for theone at one of the ends, which was allowed to contract or expandfreely in the axial direction. Both ends were allowed to rotatearound the x-axis, but restrained from rotating about the othertwo axes, thus mimicking the actual rotational degree of freedomaccommodated by the experimental setup as shown in Fig. 3.

The composite repair wrap consisted of eight layers of glassepoxy with the plies orientations assumed to be unidirectional,in the hoop direction (along the circumference) of the pipe. Threeintegration points were requested at each ply’s thickness direction.The composite material properties are shown in Table 1.

The pipe was made of X60 grade steel per API 5L [4] specifica-tions (see Table 1). The stress–strain relationship was definedusing the Ramberg–Osgood material model expression. TheRamberg–Osgood model was found to represent the relationshipaccurately, with its parameters extracted from Walker and Wil-liams [14]. The model is described mathematically by:

e ¼ rEþ a

r0

Err0

� �ð2Þ

combined loading conditions considered (all pipes with D/t = 50 are reported, as wellother D/t ratios).

Values of the individual loading component

Pressure (MPa) Dt (�C) Pt (kN) Py (kN) Pburst (MPa)

0 45 1722 6620 18.773.33 45 1722 6620 18.776.67 45 1722 6620 18.7710 45 1722 6620 18.7713.33 45 1722 6620 18.77

0 10 383 6620 18.773.33 10 383 6620 18.776.67 10 383 6620 18.7710 10 383 6620 18.7713.33 10 383 6620 18.77

0 �30 �1148 6620 18.773.33 �30 �1148 6620 18.776.67 �30 �1148 6620 18.7710 �30 �1148 6620 18.7713.33 �30 �1148 6620 18.77

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where � is the strain, r is the stress, E is modulus of elasticity, a isthe yield offset, r0 is the yield strength of the material and n is thehardening material constant.

The defect cavity is usually filled with epoxy putty. The fillermaterial properties were extracted from the Clock Spring literature[6], see Table 1). The filler material has a relatively low modulus ofelasticity, and is just there to fill the gap between the compositeand steel surfaces, acting as a stress-transferring medium for therepair to function.

1.1.2. Parametric analysisPipes with D/t ratios of 30–100 with increments of 10 were con-

sidered. A typical set of parameters considered for one of the D/tratios (i.e., D/t = 50) is tabulated in Table 2; moreover, the specifi-cations for a group of pipes with D/t = 60 are also noted in thetable. It should be noted that similar parametric combinationswere used for all other D/t ratios considered in this research. Thethicknesses of the pipes were selected to be close to the availablepipe sizes used by industry. The length of the full pipe(3000 mm) was kept constant in all models. This was to allowthe local buckling to occur without the interactions with thestrains due to the boundary conditions. In total, 15 combinationsof various parameters, for a total of 45 series of D�C�P� combina-tions were analyzed by finite element. The identifier D refers tothe D/t; C or T refers to the sense of the applied axial load (com-pression or tensile, respectively); and P refers to the internal pres-

Fig. 1. Epoxy-putty filled defect

Fig. 2. (a) A typical repaired pipe used in the investig

sure ratio, and finally the � appearing after each letter reflects thevalue of the aforementioned parameters. Thus, D30C45P20 refersto a pipe with D/t of 30, subjected to a thermally induced compres-sive axial load corresponding to 45 �C thermal gradient, and aninternal pressure corresponding to 20% of the operating pressure.

The loading sequence investigated was assumed to model field-like conditions. It was assumed that the pipe would be pressurizedon a daily basis, and at some point of time during a given year, theambient temperature would change, thus creating a thermal gradi-ent. This thermal gradient would in turn develop an axial force.Lastly, on a less frequent basis, ground movement would occur,producing increasing curvatures in the pipe. Thus, the loading se-quence investigated was as follows:

1. An internal pressure was applied, and kept constant.2. The equivalent axial force induced by the thermal gradient was

applied and kept constant.3. Finally, the curvature was increased incrementally within the

pipe, to simulate the bending resulting from the groundmovement.

1.2. Stage 2: Experimental design

1.2.1. Specimen preparationFor the experimental test pipes to be exemplary of the research

completed by previous authors [11,9,7], the size was chosen in the

in a sandblasted steel pipe.

ation schematic. (b) Close up of the wrap region.

Fig. 3. Experimental setup.

A. Shouman, F. Taheri / Composite Structures 93 (2011) 1538–1548 1541

range of pipes used commonly in practice. This size of the pipe rep-resented typical line pipes used in the oil and gas transportationindustry.

The steel pipe specimens were API 5L, X56 seamless carbonsteel pipes, schedule 10 (thickness t = 0.134 in.) with an outerdiameter of 6.625 in., with the length of 5 ft (1520 m). Our earlierinvestigations have demonstrated that this length could simulatethe response of the actual longer pipe and removes the boundarycondition effect on the repaired section. The desired defects weremachined into the pipe wall using a computer numerical con-trolled (CNC) machine, producing a defect with a longitudinallength of 6 in. (152.4 mm) and a depth of 80% of the wall thickness(0.1072 in. or 2.72 mm). In the hoop direction, defect widthspanned over a 2-in. arc. The defect size was kept constantthroughout the tests.

Once the defects were machined, 10 � 10 � 0.5 in. end plateswere welded appropriately to the pipe ends. The steel pipes werethen sandblasted, to allow proper adhesion of the FRP repair tothe steel surface. The steel pipes were then mounted on a supportstructure, to allow its rotation while being wrapped and repairedby composite fibres. The defects were filled with the epoxy putty,such that the defect profile would return to its original unalteredstate. This is shown in Fig. 1. Afterwards, the FRP repair, which con-sisted of an eight layer unidirectional E-glass fibres and epoxyadhesive, was applied. The unidirectional E-glass fabric waswrapped around the pipe using hand tension, while simulta-neously applying well-mixed epoxy onto the fabrics. It wasensured that the center-line of the FRP repair and the steel pipe-lines coincided at all times. Once the FRP repair was applied, thecomposite was left to air-cure at ambient temperature for 24 h.Fig. 2 shows a typical FRP repaired pipe specimen.

1.2.2. Experimental setupA detailed schematic of the experimental frame setup is shown

in Fig. 3. This experimental setup is capable of applying a combina-tion of axial and bending forces. The pipes are placed vertically un-der the loading head of a Tinus Olsen loading frame. The TinusOlsen machine is a mechanical loading frame, with the capacityof 252,000 lbs. In this test setup, the machine was used to applythe concentric axial compressive loads. Each end plate, originallywelded during the FRP repair wrap process, was bolted to other10 � 10 � 1 in. plates that were welded to the top of the hollowsteel section (HSS) used as the loading arm. Five ASTM A4900.5 in. bolts were applied to the tension side of the pipe, and twoon the compression side. The loading arms consisted of a6 � 6 � 5/8 in. HSS section. The arms had a length of 80 in. andwere stiffened with an internal beam with a cross-section dimen-sion of 500 � 200 � 7000, as shown in Fig. 3. On the opposite side of theHSS sections, right under the 1000 � 1000 � 100 end plates, a hingesystem was constructed to allow for the end rotations of thebeams, hence the pipe specimens. This in turn would create thedesired bending during the test. This hinge system included twodouble L-angles as the base of the hinge system, which was weldedto another 10 � 10 � 0.5 in. plate, which was in turn bolted to theTinus Olsen base. Four half-inch bolts were used (one on each cor-ner of the plate), to secure each plate. Lastly, the double back-to-back L-angles were stiffened in the bending plane.

On the other end of the beam, 72.5 in. away from the center-lineof the pipe, a hydraulic jack system was placed to apply an eccen-tric axial force thus generating the bending forces to the pipe ends.

1.2.3. Instrumentation and measurementsVarious types of instrumentation devices were used to record

the data. Six strain gauges were placed along the compressive sideof the pipe, along the bending plane of the pipe. The positions andmeasured strain component by each strain gauge was as follows:

1. Strain gauge #1 was placed 300 below the bottom end of the FRPrepair (this is where the local buckling was anticipated to occuraccording to the FE analyses);its purpose was to record thehoop strain in the steel pipe.

2. Strain gauge #2 was placed 300 away from the bottom end of theFRP repair; its purpose was to record the axial compressivestrain on the pipe.

3. Strain gauge #3 was placed on the FRP repair section, just abovethe midpoint of the defect area; its purpose was to record thehoop strain on the composite repair section.

4. Strain gauge #4 was placed at the midpoint of the steel defectsection, under the filler material, beneath the FRP repair; itspurpose was to record the hoop strains on the steel pipe inthe defect region.

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5. Strain gauge #5 was placed 300above the top end of the FRPrepair section. It measured the hoop strain in the steel pipe,similar to strain gauge #1, so the comparison of the strains atboth ends of the FRP repair section could be facilitated.

6. Finally, strain gauge #6 was placed 300 above the top end of theFRP repair section. It measured the axial compressive strain; itspurpose was to produce data that could be compared with thatobtained from strain gauge #2.

Moreover, two Linear Variable Differential Transducers (LVDTs)were used; the first LVDT was placed at the mid-span of the pipe,on the tensile side of the FRP repair section. It was used to measurethe mid-span deflection of the pipe, from which the curvaturecould be obtained. The second LVDT was placed at the bottomHSS section. It was to measure the vertical end displacement ofthe HSS section as the load by the hydraulic jack was being applied.In addition to the LVDTs, two digital inclinometers were used tomeasure the inclination angles of the HSS sections, hence the pipeends rotations. Each inclinometer was placed at the end of the HSSbeam section close to the hydraulic jack (72 in. away from the neu-tral axis of the pipe). They estimated the end rotations at both endsof the pipe.

1.3. Stage 3: Limit design of composite wraps

ASME has incorporated composite repairs for damaged pipes, byrecently developing the ASME Post Construction Repair Standard-PCC2 [5]. This standard has set a minimum FRP repair thicknessto be applied to damaged pipelines that can be computed by thefollowing equations:

trepair ¼1

Ecec� PD

2� ðSMYSÞts

� �ð3Þ

where Ec is the tensile modulus of the composite in the hoopdirection, ec is the allowable hoop strain in the composite, P isthe applied internal pressure, D is the outer diameter of the

Fig. 4. Finite elem

steel pipe, SMYS is the Specified Minimum Yield Strength of thesteel pipe and ts is the minimum remaining wall thickness inthe pipe.

Moreover, Alexander [2] determined through a combination ofexperimental and analytical investigation that the minimum re-quired thickness of composite repair for a damaged pipe could beobtained from the following equation:

trepair ¼1sc� Pf D

2� sats

� �ð4Þ

where sc is the composite’s tensile strength in the hoop direction, Pf

is the failure pressure of the undamaged pipe, D is the outer diam-eter of the pipe, sa is the yield strength of the pipe and ts is the min-imum remaining wall thickness in the pipe. The difference from Eq.(3) being that the allowable hoop strain is not a defined criterion incomparison to the composite tensile strength, which is well definedin codes and textbooks.

Once the composite repair thickness is determined, it can bechecked against the yield strength or ultimate tensile strength ofthe pipe. The hoop stress in the pipe can be calculated by [2],

rhoop ¼PR

tp 1þ Ectrepair

Eptp

h i ð5Þ

where rhoop is the hoop stress in the steel pipe, P is the applied pres-sure, R is the mean radius and tp is the thickness of the steel pipe.

In addition to reinforcing the pipe against burst pressure, thepipe might also be under an axial force, in which case the repairhas to be of adequate length to maintain the bond between thesteel and composite interfaces [2]. The required minimum addi-tional length on either side of the defect region is determined as[2],

L PR

pDsadhessive� SF ð6Þ

where L is the minimum repair length on each side of the defect, F isthe longitudinal axial force (tensile), D is the outside diameter of the

ent model.

A. Shouman, F. Taheri / Composite Structures 93 (2011) 1538–1548 1543

pipe, sadhessive is the adhesive lap shear strength and SF is a safetyfactor equal to 1.5.

In an effort to improve the economic feasibility of the compositerepair, a short parametric finite element study was conducted. Theprincipal aim of this exercise was to develop limits on the FRP de-sign, incorporating the design requirements and loading conditionsapplied on the FRP repaired pipelines.

The following parameters were examined:

� Geometry of the composite repair.� Material properties of composite repair.

From the parametric study, the following results were extractedand investigated:

� Stresses and strains in the composite repair system.� Strains in the steel.� The minimum repair thickness of FRP repair to withstand burst

pressure.

The geometry of the models in this stage was based on stage 1.The steel and filler material properties were not varied. The FRP re-pair was to be varied in thickness and length.

Fig. 6. Plot of moment versus longitudinal compressive strain response for D70C45.

2. Results and discussion

2.1. Stage 1: parametric modelling results

The conducted parametric study investigated the effect of D/tratio, the applied and axial load and the internal pressure effectson the response of FRP repaired pipelines. The logarithmic strainswere extracted and analyzed. The effect of pressure is illustratedin Figs. 5–7. The pressures considered corresponded to the ratioof hoop stresses to yield stress. Pressures amounting to 0%, 20%,40%, 60% and 80% of the yield stress in the hoop direction wereinvestigated. It is evident that as the pressure increased, themoment capacity of the pipeline decreased. Therefore, an increaseof pressure would have an adverse effect on the ultimate momentcapacity of the pipe. This is due to the pressure increasing the cir-cumferential stresses within a pipe wall, which in turn would al-low the pipe to reach the onset of buckling at lower moments,thus reducing the ultimate moment capacity.

Moreover, it can be seen that at lower pressures, higher ulti-mate moment capacity is reached at lower strains. Furthermore,

Fig. 5. Plot of moment versus longitudinal compressive strain response for D50C45.

pipelines subjected to higher pressures were able to undergo high-er strains, such that the severity of the local buckling is increased.Moreover, it was seen that pressurized pipes buckled locally in theform of an outward bulge, in contrast to the unpressurized pipesthat locally buckled in the form of inward bulging, or diamondshaped buckles. Both modes of buckling initiated as wrinkling, inthe form of minor sinusoidal waves. However, only one of the min-or wrinkle waves increased in magnitude, and formed into eitheran outward bulge or diamond shaped buckle.

Moreover, the effects of the axial force and D/t ratios on thestrain-based response of composite repaired pipelines were alsoinvestigated. The moment-compressive strain response varied sig-nificantly with the change of D/t ratio. It can be seen from Fig. 8that as D/t ratio increased (i.e., as pipe’s wall becomes thinner),the ultimate moment capacity decreased. It must also be notedthat the lower the D/t ratio (thicker pipe), the faster the ultimatemoment capacity was reached at lower strains. In addition, thestrain capacity undertaken by the pipe increased as the D/t ratio in-creased. This is because pipes with D/t ratio of 70–100 behave in amore ductile manner during the onset of buckling phase, leading tothe formation of local buckling and attainment of a lower limit mo-ment capacity.

Fig. 7. Plot of moment versus longitudinal compressive strain response forD100C45.

Fig. 8. Plot of moment versus longitudinal compressive strain response for differentD/t ratios at rh/ry = 0.8.

Fig. 9. Plot of moment versus longitudinal compressive strain response for differentapplied axial forces at rh/ry = 0.8 for D/t of 50.

Fig. 10. Plot of moment versus longitudinal compressive strain response fordifferent applied axial forces at rh/ry = 0.8 for D/t of 70.

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The third dimensionless parameter considered was the ratioof the axial force due to the thermal gradients, to the axial forcethat would cause the pipe to yield. The effect of the variation ofapplied axial force on D/t ratios of 50, 70 and 100 are shown inFigs. 9–11, respectively. It can be seen in all three figures thatthe maximum ultimate moment capacity was attained whenthe pipe was subjected to a tensile force. This is due to the pipegaining additional stiffness due to the tension within the pipewall. This tensile force urges the pipe to respond in a ductilemanner, thus inhibiting any form of local buckling. The failureof pipes subjected to tension would probably occur as a tensilefracture before any compressive local buckling failure would oc-cur. On the other hand, as the compressive axial force is in-creased, the moment capacity of the pipe is decreasedsignificantly.

Fig. 11. Plot of moment versus longitudinal compressive strain response fordifferent applied axial forces at rh/ry = 0.8 for D/t of 100.

Fig. 12. The GDS pressure machine (on table) and the Tinus Olsen Controller (inbackground).

Table 3Strain readings from experimental specimens and the corresponding values obtained by the finite element models.

Longitudinal compressive strain Hoop strain (in the pipe) Hoop strain (in the defect region) Hoop strain (in the composite wrap)

Expert’l FEA Expert’l FEA Expert’l FEA Expert’l FEA

Pipe 1Strain �1.36% �1.57% 1.17% 1.04% 0.13% 0.12% 0.24% 0.26%% error 13.5% 12.3% 12.0% 8.1%

Pipe 2Strain �4.42% �4.68% 1.29% 1.15% 0.019% 0.11% 0.09% 0.12%% error 5.5% 12.4% 89.4% 25.5%

Pipe 3Strain �0.99% �1.16% 0.78% 0.66% N/A 0.14% 0.25% 0.27%% error 14.3% 18.2% N/A 7.9%

A. Shouman, F. Taheri / Composite Structures 93 (2011) 1538–1548 1545

2.2. Stage 2: experimental results

The experimental study was conducted to verify the results ofthe finite element modelling. The strains on the experimental pipeswere compared with those of the finite element modelling. More-over, the location and shape of the buckle were also compared.

Fig. 14. Comparison of the actual buckling shapes an

Fig. 13. Comparison of various the strain components recorded on the pipespecimens.

Three pipes were tested. Table 2 also shows the typical loadingconditions imposed on the pipes (in this case for the D/t ratio of50). Four strain gauges were placed on each pipe at identical loca-tions. All the strain gauges measured the circumferential strain ex-cept for strain gauge 1 which measured the longitudinalcompressive strain.

The loading pattern applied was as follows, with the magnitudeof individual loads varying for each specimen:

� Internal pressure was applied and kept constant using the pres-sure machine shown in Fig. 12.� Concentric axial load was applied using the Tinus Olsen

machine, and kept constant throughout the bending phase.The Tinus Olsen Controller is shown in Fig. 12.� Eccentric axial load was applied using a hydraulic jack. This

served as the curvature to be applied at the ends of the pipe.The axial load was applied as a displacement controlled load.

The top and bottom end rotations of the pipe were recorded peri-odically using the digital inclinometers as the eccentric axial loadwas applied (see Fig. 3). The finite element models were constructedsimilarly to the experimental setup, with identical applied loadingconditions. However, the unequal end rotations occurred in theexperiments were ignored; instead, the average of the top and bot-

d those predicted by the finite element method.

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tom rotations was applied to the models. Moreover, the materialproperties of both the steel pipe and composite wrap were extractedthrough ancillary tensile tests. However, it should be noted thatsince the composite test specimens were extracted from plates thatwere hand-laid, better consolidated and vacuum bagged, there couldbe slight variation in the properties between the test coupons andthe actual repair wraps.

Table 3 shows the comparisons between the strain gauge read-ings extracted at the limit moment from the experimental dataand the corresponding values generated by the finite element mod-els. It should be noted that in the cases of pipe #2 and #3, straingauge #3, which was placed at the defect region, under the filler,was damaged during the fabrication, and its readings remained un-changed during the tests.

It can be seen that there is reasonably good agreement betweenthe experimental and finite element generated values. As seen fromthe results reported in Table 3, the maximum errors varied frompipe to pipe, with the maximum error occurring in pipe 2, equalling25.45%. Regardless of the loading conditions applied, the finite ele-ment was able to yield reasonably close predictions. Moreover, thecomparison between the strain gauge readings on different pipes isshown in Fig. 13. It can be seen that pipe 2, which was subjected tothe harshest loading combination, endured the highest strains (re-corded by strain gauges 1 and 2, in the axial and hoop directions).This agrees with the finite element results, in which the severityof the local buckling increased, as the pressure increased.

The location and type of buckle in each experimental pipe spec-imen was also compared to those predicted by the finite elementmodels. The observed buckling modes are in concert with those re-ported by Alexander and Ochoa [3]; they also observed that withthe increase of the bending moment, the pipeline would tend tobuckle away from the corroded region, and thus would most likelyfail outside the repaired region. A comparison of the pipe specimencontaining outward bulging is shown in Fig. 14. It can be seen thatthe finite element was able to predict the mode of local bucklingand its location quite accurately. Moreover, the strains occurringclose to those local buckles were also closely similar to those ob-served during the experiment. The slight errors in the strains andamplitude of the buckle might be due to one of the following:

� In the models, it was assumed that the end rotations wereequal, so that the models could be reduced to a quarter-symme-try. However, it was found this was not the case in the actualexperiments, in which unequal rotations was commonlyexperienced.

Fig. 15. Effect of thickness of composite wrap on moment–curvature response ofrepaired pipelines.

� Possible variations in composite material properties in theancillary tensile tests and the composite repair wrap.� The combined loading state made it rather complicated to

maintain the loading rates constant throughout the test (i.e.,the bending loading rate vs. concentric axial loading rate).� Pressurized pipe specimens leaked during the plastic deforma-

tion phase in the tests, due to the high amount of stresses inthe pipe wall.� The residual stresses created during the welding of the end

plates to pipes were not incorporated in the FE models.

2.3. Stage 3: limit design of composite wraps

A parametric finite element study was conducted to investigatethe limits of the design parameters of the composite wrap. For thispurpose, the thickness and length of the composite wrap were var-ied. The thickness was varied as the ratio of the original thicknesscalculated from Eq. (4), as follows: 0.25, 0.5, 0.75, 1.5, 1.75 and 2.The wrap length was calculated based on Eq. (6), and was variedas no-length (i.e., no wrap), the length obtained by using Eq. (6),and double that length.

Fig. 15 shows that the thickness of the FRP would have noapparent effect on the moment curvature response. This is becausethe FRP was assumed to be unidirectional in the circumferentialdirection of the pipes, thus it does not add much stiffness to thepipe against bending. Furthermore, Fig. 16 supports this hypothe-sis, as it is seen there is no visible effect on the longitudinal strainsin the bulging region adjacent to the composite repair section.Again, this is due to the lack of axial reinforcement in the compos-ite wrap system. Moreover, Fig. 16 shows that as the thickness ofthe wrap changed, the severity of the outward bulge remainedthe same. Fig. 17 shows the relation between the moment andthe hoop strain measured at the mid-span of the pipe at the defectregion. It is seen that as the thickness of the composite wrap is in-creased, the initial hoop strain at the beginning of the bendingphase would decrease when the thickness becomes greater thanthe thickness calculated by Eq. (4). Moreover, under such a circum-stance, regardless of the value of the moment applied, the defectarea would not yield, (assuming steel would yield at 0.2% strain);however, the defect area would yield and cause the pipe fail underthe internal pressure, if the wrap thickness is taken as theminimum.

Figs. 18 and 19 show the effect of wrap’s length variation on thebending response of the repaired pipelines. The results indicate

Fig. 16. Effect of thickness of composite wrap on moment-longitudinal strainresponse of repaired pipelines.

Fig. 17. Effect of thickness of composite wrap on moment-defect hoop strainresponse of repaired pipelines.

Fig. 18. Effect of length of composite wrap on moment–curvature response ofrepaired pipelines.

Fig. 19. Effect of length of composite wrap on moment-defect hoop strain responseof repaired pipelines.

A. Shouman, F. Taheri / Composite Structures 93 (2011) 1538–1548 1547

that as the length of the composite wrap increases, the ultimatemoment capacity also increases, and the curvature decreases. Thisis due to the added stiffness provided by the wrap to the pipe.Fig. 19 also shows that increasing the length would in turn increasethe hoop strain carrying capacity within the defect region.

3. Conclusion

Composite repaired pipelines have been traditionally investi-gated under internal pressure. However, combined loading statesare often experienced by pipeline in service. A numerical paramet-ric study was conducted to further our understanding of the behav-iour of composite repaired pipelines under different loadingconditions, focussing on its applicability in reference to strain-based design. The results of the study indicate that the longitudinalstrains bypass any strain limits because of its direct relation to thelocal buckling phenomenon occurring adjacent to the compositerepaired section.

The second part of the paper focused on the experimental studyof composite repaired pipelines under internal pressure, axial load-ing and bending force. Strains were extracted at various locationsand compared to finite element models based on the experimentalspecimens. Results were found to compare reasonably well, verify-ing that finite element modelling is capable of capturing the re-sponse of composite repaired pipelines under various loadingconditions.

The third and last section of this paper focused on investigatingthe limiting parameters that can be used in composite repair wrapdesign. It was shown that increasing the thickness of the wrapcould not enhance the strength of the pipe in the axial direction;however, the added thickness could prevent the yielding of thepipe at the defect region. Nonetheless, it is highly recommendedthat a safety factor be applied to the minimum thicknesses thatis computed by the available equations in the recent codes, thusensuring that the defect strains remain within an acceptable limit.Moreover, the length of the composite repair was found to affectthe bending response of the repaired pipelines. Wrap length mighttherefore be a possible parameter that could enhance the axial per-formance of such repaired pipes.

In conclusion, composite repair wraps offer a viable solution forrehabilitating damaged pipelines subject to combined loading con-ditions. However, a more optimized design with varying fibre ori-entations should be considered to enhance the safe load carryingcapacity of such pipes.

Acknowledgement

The financial support of the Natural Sciences and EngineeringCouncil of Canada (NSERC) in support of this work is gratefullyacknowledged.

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