[oc] end fittings for composite risers

8/12/2019 [OC] End Fittings for Composite Risers

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OTC 23977

Development and Qualification of End Fittings for Composite Riser PipeStephen Hatton, Luke Rumsey, Praveen Biragoni, Damon Roberts, Magma Global Limited

Copyright 2013, Offshore Technology Conference

This paper was prepared for presentation at t he Offshore Technology Conference held in Houston, Texas, USA, 69 May 2013.

This paper was selected for presentation by an OTC program committee following review of information contained in an abstract submitted by the author(s). Contents of the paper have not beenreviewed by the Offshore Technology Conference and are subject to correction by the author(s). The material does not necessarily reflect any position of the Offshore Technology Conference, itsofficers, or members. Electronic reproduction, distribution, or storage of any part of this paper without the written consent of the Offshore Technology Conference is prohibited. Permission toreproduce in print is restricted to an abstract of not more than 300 words; illustrations may not be copied. The abstract must contain conspicuous acknowledgment of OTC copyright.

AbstractWhilst much interest is often focussed on the composite pipe body it is often forgotten that a reliable end fitting is a

prerequisite of a successful composite pipe application and further that the design challenge of the end fitting is more

challenging than the pipe itself.

The purpose of this paper is to present the end fitting arrangement for a composite pipe manufactured from carbon fibre and

PEEK polymer. The paper describes the design approach and testing/qualification program employed for the end fitting to

demonstrate reliable application in a critical and structurally demanding application such as for a deep water dynamic riser.

The paper summarises historical design approaches and design alternatives and explains the reason for selecting the proposed

arrangement. It presents the design process used to develop the design and to predict its structural response. It discusses

manufacturing issues and describes the test program conducted to prove the end fitting performance.

Industry focus is often on the pipe rather than the end fitting. However the latter often presents a more difficult design

challenge. Without a reliable design solution and methodology composite pipe application cannot be considered. The

development work presented in the current paper therefore presents an important step towards application of compositetechnology on demanding riser applications.

An end fitting with reliable structural and sealing performance is a prerequisite for the successful application of compositepipe. Historically, this has proven to be a difficult challenge and end fitting performance limitations have been cited for the

slow introduction of composite pipe technology. The current paper presents a new design approach to the problem, made

possible by a unique manufacturing process and a better understanding of the composite metallic interface, which together

allows the end fitting problems to be resolved. The paper describes fundamentals of the design approach, the development

work conducted, manufacturing and qualification testing. The paper discusses the function specification, key design features,FEA approach and results, codes and standards, testing results under combined load conditions and fatigue testing results.

Introduction

Design, fabrication and installation of riser systems for floating production is a complex challenge. Constraints such asmaximum payload capacity and design issues such as internal and external corrosion, fatigue capacity, thermal insulation,

weldability and susceptibility to sour service conditions conspire to make riser design one of the most demanding challenges in

the offshore industry today. The magnitude of this challenge has grown, perhaps non linearly, as water depths have continued

to increase over the last 20 years. Also operating pressures and temperatures have increased, particularly in the last 10 years,

with the focus on presalt and subsalt reservoirs.

Composite pipe technology has been proposed many times in the last 20 years as a potential technology to resolve some of thedesign challenges [1,2]. Composite pipe can bring important performance advantages over steel pipe risers and in some

instances unbonded flexible risers with metallic armor layers. These advantages include:


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Lower weight

Improved fatigue capacity

Corrosion resistance

Higher strain limits

High strength

Ease of deployment

The design advantages offered by composite materials can simplify the problems that riser designers typically face when

working with steel pipe and designing for demanding applications. In such cases it is typical that many of the following design

strategies need to be employed inorder to achieve the levels of performance required:

More sophisticated analysis methodologies

Provision of higher quality design basis assumptions (eg environmental and fluid data)

Increased materials and component testing

Application of higher strength steels

Weld quality improvement (Welding and inspection)

Internal cladding such as inconel or polymer liners

Application of VIV strakes

More compliant riser configurations eg the use of buoyant configurations

Chemical injection for corrosion control and process flow management

Higher capacity installation vessels

Higher capacity host production vessels Lower motion host production vessels

More demanding inspection strategies

More regular replacement strategies

The above design strategies for steel pipe risers have cost and schedule impacts meaning that riser costs can be a significant

percentage of the overall development cost. Unfortunately, the complexity and knock on effect of some of the above issues canbe underestimated or even missed at the early project stages leading to a difference in the as installed riser cost compared to

the predicted cost, leading to project cost and schedule overruns.

Therefore the proposed use of composite pipe brings the potential for some simplification of the riser design process and the

potential for improved performance, reliability and overall cost benefits. Whilst there appears to be a belief across the industry

that composites can deliver these benefits there is also an uncertainty as to whether the technology is sufficiently mature forproject application. One of the key design areas that needs to be better defined is that of the composite pipe end fitting, which

despite previous initiatives remains an area of some uncertainty. This is because the problem of terminating a composite pipewith a metallic end fitting, poses some seriously complex design and manufacturing challenges.

Whilst throughout this paper we use the term Composite, it should be remembered that the term Composite is allencompassing and as such covers technologies from e-glass /polyester structures at one end of the performance spectrum

through to carbon fibre/ PEEK at the other. The final end termination design and reliability are ultimately dictated by the

inherent performance on the Composite choosen for the application in question.

Background and Historical Approach

Composite pipe must be terminated with an end fitting that allows connection of the composite pipe to standard metallic oil

industry interfaces such as API or ANSI flanges or hub connections. The design of these composite end fitting is potentiallymore complex than the pipe itself.

Composite materials are technically more complex to use in design when compared to metallic materials due to their an-

isotropic properties. When such materials are interfaced with metallic components the problem is even more complex. The

difference in structural properties between composite materials and metallic materials, typically steel and titanium, makes thedesign of the interface highly problematic. The main challenge is that steel and composite materials have very different

coefficients of thermal expansion and thermal conductivity and also different stiffnesses and Poissons ratios. Therefore as the

end fitting or pipe is heated or externally loaded the metallic and composite materials respond in a different structural mannerand this can lead to failure of the interface by cracking or disbondment of one material away from the other. As a minimum,

this may cause a failure of the end fittings ability to maintain leak tight integrity. In the extreme, this may reduce the end

fittings structural capacity or fatigue performance. Failure of such interfaces can be progressive, such that on repeated loading

the failure mechanism progresses along the interface until it fails catastrophically. The magnitude and complexity of this

problem is evident from the complexity of the solutions proposed spanning many years.


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TrapLock

The most popular approach has been the Trap-Lock solution patented by Baldwin/Reigle/Drey in 1997 [3]. [1]In this design

the composite material is wound over a metallic mandrel that has been machined with a profile as shown in Figure 1. Each

element of the profile provides the ability to transfer load from the composite into the metallic component via a combination of

tensile and shear mechanisms.

The load path in such a design is dependent on many factors including fiber direction, stiffness, pretension etc. Often a hoop

wound material is laid over the top of the trap lock or an outer steel collar is applied. High levels of preload are required at the

interface to maintain contact between the inner metallic mandrel and outer composite materials under all load and temperatureconditions.

The detail design of the Traplock must prevent the majority of the load from being taken at the first thread. This is achieved

by a combination of preloading of the composite material and optimising the structural stiffness of the mandrel. This preventshigh local stresses and potential for local failure which can progress along the interface on subsequent loading. To avoid this

failure process it is necessary to preload the traplock interface by winding the fibres under tension and/or auto frettage where

the steel mandrel is plastically deformed by internal pressure after winding. This can give good results but remembering thatcomposite materials can creep and initial preloading can be reduced over time, an effect that is accelerated under elevated

temperatures. This effect must be taken into account.

In the Traplock design the profile uses one or more "traplock" grooves in the exterior surface of the mandrel into which thefilament or reinforcements of the composite tube are wound and/or compacted.The axial load is transferred between the

composite tube and the end fitting through bearing on the load-carrying face of the traplock grooves. The surface area of the

load-carrying face is one of the parameters determining the strength of the joint or interface. The bearing area can be increasedby increasing the height of the load-carrying face but the bearing stress which the composite material can support is limited

and therefore the diametrical envelope required by a single traplock groove can become quite large as the height of the load-

carrying face is increased.

The diametrical requirements of the joint can be reduced by the use of multiple traplock grooves, but as discussed above it isimportant to make sure that the grooves carry equal load. This response is optimized by varying the thickness of the root of

the mandrel so that it becomes progressively stiffer from its tip to its other end. This approach further reduces the difference in

stiffness between the two materials and thus its ability to better share the load.

An additional design issue is that of establishing and maintaining a pressure-tight seal between the composite tube and the end

fitting. This is particularly true if the composite tube has an internal pressure sealing liner, which is terminated at the tip of the

traplock mandrel. Designing a reliable elastomeric seal at this location can be a challenge due to practical manufacturing issuesaround the need to lay-up and cure resins, which can easily contact seal surfaces and contaminate them. The reliance on an

adhesive bond between the end fitting and the tube liner is also not that reliable due to the differential movement in the axial

direction that is inherent in traplock system operation. As the fitting moves outboard under load, the liner material andadhesive typically cannot accommodate this differential movement without the possibility of high streses and ultimately

cracking, tearing or disbonding. The solution proposed is to use an elastomeric seal at the interface between the liner andtraplock mandrel in a location where it cannot be easily damaged or contaminated during pipe construction. This is typically

located at the tip of the Traplock mandrel and needs to incorporate profiles that are pressure energised when the pipe is

internally pressurised.

A commercial consideration of this design approach is that it is a one off shot and in the event that a reliable seal is not

achieved it is not possible to replace it, the joint is effectively irreparable.


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Figure 1 - Traplock Design Example

Swaged End Fitting

An alternative approach shown in Figure 2 is to use a metallic inner and outer sleeve that effectively sandwiches the composite

material and its internal/external liners.

This is a relatively simple design approach that has the benefit of assembly onto plain ended composite pipe. The end fitting

can be assembled in a number of ways but typically the end fitting is fitted over the pipe end as an assembled unit. The inner

spigot has an interference fit with the bore whilst the outer sleeve is a close sliding fit. After assembly of the end fitting overthe plain pipe end the outer sleeve is swaged to reduce its internal diameter and generate a high interference with the

composite pipe. The high interference generates a combination of friction and mechanical interference that is capable of

transferring structural loads.

The inner mandrel is machined with a serrated or saw tooth outer profile that is designed to engage with the composite pipeliner. Additionally, it also incorporates a nose seal, typically an elastomeric O ring that seals against the bore of the liner pipe.

The issues associated with this arrangement are:

Unlike with the trap-lock design, where the axial stiffness of the mandrel can be readily optimised, load transmissionfrom the steel to the composite elements is focussed on the first thread and this can lead to higher local stresses

The inner mandrel design results in a bore restriction

Sealing mechanisms are typically O ring technology and it is not easy to incorporate preloaded seal solutions

There is a high stress concentration factor at the transition from the stiff steel outer sleeve to the nominal compositepipe through both the high difference in bending stiffness and fact that the composite pipe within the fitting ismaintained in hoop compression whilst at the exit from the end fitting the pipe transitions to hoop tension causinghigh local bending.

The steel outer sleeve and inner mandrel interface with the weaker liner and external coating materials and not withthe structural core composite material.

There is potential for damage of the inner liner surface resulting in uncertain mechanical interface as the innermandrel is inserted

Thermal and structural loads can result in differential movements between the steel and composite elements that needto be considered in the design process

High local stresses at the interface between the composite elements and steel elements result in high potential forcreep and relaxation affecting the long term response.

To overcome some of these issues it is typically necessary to thicken the wall of the composite pipe locally to achieve an

acceptable strength consistent with the pipe strength performance. The construction of this end thickening may need to bedifferent to that of the pipe construction to accommodate the nature of the loads at this location.

Such a design is well proven in smaller diameter hose products but is not considered as attractive for larger diameter, highpressure and dynamic applications. This is because the loads required for the swaging process get prohibitively large, the

swaging process is complex metalurgically, especially where there may be H2S and the end fitting must accommodate

dynamic loading


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Figure 2 - Swaged End Fitting Design Example

Metallic Liner End Fitting

Recent composite riser pipe development has considered a Hybrid composite tube [4,5], where a thin walled steel or titanium

liner is used along the entire pipe length and which is structurally continuous, via a weld, with the end fittings. The externallyapplied composite material is then used to enhance the steel pipe performance and is wound over the top of the steel liner. This

reinforcement is achieved by a combination of hoop and axially placed fibers.

This arrangement significantly reduces the complexity of the end fitting design and has three important benefits:

The metallic liner provides an impermeable membrane that precludes the problem of gas migration across the liner,which is a common problem for many polymer materials

It simplifies the end fitting design by avoiding the need for a seal assembly between the pipe and end fitting as thepipe and end fitting are welded.

It allows axial and hoop loads to be accommodated separately ie it allows axial loads to be primarily accommodatedby the steel pipe and the hoop (burst) loads accommodated by the composite material

In the extreme design case, the composite fibers can be used to provide only hoop or burst reinforcement and all axial loads are

taken by the steel liner. In this case there is no need for a trap-lock interface as all loads are hoop. Where the compositematerial is configured to take both hoop and axial loads the trap-lock style interface may still be required.

This configuration does not offer such a large weight saving benefit as an all composite pipe structure, due to the higherweight of the liner, its minimum bend radius and fatigue performance are dictated by the steel pipe and by selected weld

details rather than the composite material and there remains a design challenge regarding differential material strains when

subjected to temperature gradients and/or external loads.

In the hybrid tube approach corrosion is a remaining issue and often the steel liner needs consideration to both internal and

external corrosion allowances. An external corrosion allowance may be required if there is potential for water to ingress

between the steel pipe and composite material where such an interface can present a challenging corrosion environment. Theneed to provide such corrosion allowances can further degrade the weight benefits of such a structure and when this is coupled

with fatigue and corrosion issues of the steel pipe it is not evident that such an approach offers a significant advantage over an

all steel approach.

Where this approach does offer an important benefit is for ultra-high pressure applications where without such an approach

extreme metallic wall thicknesses may be required, potentially beyond the practical limits of manufacturability or weldability.In such an application, the hybrid tube approach may be considered an important enabling technology.

During manufacture the composite is cured at an elevated temperature but when the pipe cools down the steel pipe contracts

more than the composite material and a gap occurs between the two. To prevent the two materials adhering to each other and

causing unwanted stresses this process must be managed by using a release agent applied to the pipe prior to coating.Additionally, a rubber layer is applied as the first layer before the composite material is wound this is also applied to the

traplock area so that the composite materials can be completely encased with an outer rubber layer.


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Winding the composite material into the 3D traplock profile must be conducted accurately and care must be taken to manage

winding turnarounds to make sure that fiber angles are correct and material placement achieved to avoid potential for gaps and

defects during the turnaround process.

After the composite has been fully laid and the traplocks filled with 90degree fiber the whole joint undergoes a curing and then

autofrettage process that produces an interference fit between the steel pipe and the composite material. During the

autofrettage process it is difficult to control the interference between the traplock and the composite material due to its muchgreater hoop stiffness than the pipe body itself. Consequently, there is a possibility that at the critical end locations optimum

preloading of the traplock profile is difficult to achieve.

Figure 3 - Metallic Liner Design Example

Magma End Fitting Design

The end fitting design, developed by Magma and discussed in the following sections builds on the experience and knowledge

developed from previous end fitting arrangements. The end fitting is designed to allow termination of fully composite pipes

that have no metallic liner. Fundamentally, the design separates structural and sealing functions and thereby simplifies the

problem of having materials with dissimilar structural and thermal properties. The end fitting manufacture is also somewhat

simplified by virtue of the composite material selected, which is a thermoplastic rather than a thermoset as typically used in theabove designs. This is an important difference as it allows materials to be easily added at multiple manufacturing stages which

cannot be easily achieved with the thermoset process.

A schematic of the arrangement is shown in Figure 4. The composite material under consideration is a high performance

composite comprising high strength carbon fibre and Victrex PEEK.

PEEK is a thermoplastic semi-crystalline polymer material with a glass transition temperature of 143degC but with goodstructural properties well above this temperature up to 200deg C.

The carbon fibre used is supplied by Toray, a high performance fibre with a tensile strength of 4.9GPa.

Both materials offer good general performance characteristics and together produce a composite material with some of the

highest possible performance characteristics with respect to structural strength but also corrosion, fatigue, chemical resistance,

aging and permeation. The combination of PEEK and carbon fiber is well proven and regularly used in critical oil & gas welldrilling and completion and aerospace applications.

An important point, with respect to the proposed end fitting design, is the ability of the manufacturing process to allow thewall thickness to be built-up locally at the pipe end whilst maintaining the structural performance through the full section,

consistent with the main pipe body and with no strength knock-down.

This build-up can be conducted during the initial pipe manufacture or as a secondary thickening process onto the end of plainended pipe as discussed above. The properties of this build-up can also be optimised in terms of fiber orientations to best suitthe structural needs of the end fitting. Also, importantly from an operational point of view, an end fitting can be cut off a pipe

and a new build-up and end fitting applied. This could be conducted many years after the initial pipe manufacture.

Once the pipe end has been thickened in this manner the material can be machined to accurate dimensions to suit the end

fitting design.


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In the proposed design the pipe end is gradually thickened in two stages. The first thickening provides a transition from the

nominal pipe dimensions to the much stiffer end fitting and therefore allows the stress concentrations that occur at this point to

be appropriately managed.

The second thickening is more significant than the first and allows provision of tapered section to interface with a steel outer

collar. The benefit of increasing the wall thickness in this manner is that it significantly increases the local strength and

stiffness of the pipe at this section. This allows higher local loads to be applied with small strains.

The taper angle selected is a compromise between generating radial preload, minimising axial movement, allowableinterlaminar shear properties and spreading peak loads. The angle was selected only after completing a number of tests to

understand the effects of different angles and also surface finishes and friction factors.

The outer steel collar has a similar mating tapered profile to the composite pipe and is preloaded onto the taper using a

hydraulic tool. The preload is then locked into the end fitting using a locking nut that is wound in to take up the gap that occursduring preloading. This approach ensures that the preload is accurately applied and takes account of the deformations that

occur in both the steel and composite materials. The preload is selected to be greater than the highest service load, ensuring

positive load is maintained.

Sealing is provided by the use of a conventional pressure energised bore seal with an AX profile. This arrangement provides asmooth internal bore and as a percentage of the end fitting preload is taken by the seal it is considered to be a high integrity

sealing arrangement. The bore seal material can be either PEEK or stainless steel. On the composite pipe side the seal sits in aPEEK lined pocket and on the steel side a conventional inconel inlaid pocket is used.

Figure 4 - Magma End Fitting Design Example

Magma End Fitting Design Issues

The design features that were identified as being important at the start of the development process are summarised as follows:

Be structurally as strong as the rated pipe

Have a fatigue performance as good as the pipe

Have smooth unobstructed bore

All seals can be back pressure tested and are replaceable

The end fitting can be disassembled, inspected and reassembled with replacement parts as necessary

The design must use a preloaded seal to ensure high integrity

Metallic sealing surfaces are stainless steel inconel inlaid to prevent seal surface degradation

The design is preloaded to ensure fatigue performance of steel components and ensure seal integrity at high load

The overall dimensions are minimised to reduce weight and cost

Critical interface stresses are minimised to avoid the problem of creep which in all composite design problems is akey design issue encountered in the development of the end fitting are discussed below.

Creep is time, temperature and load dependent and generally it can degrade the long term strength of a composite structure. Inthe case of the end fitting, creep has the potential to reduce initially applied preloads such that over time the sealing integrity

or load capacity is compromised. The consequence of creep in the current end fitting design is to allow preload at the end face

of the build up section to be relieved such that under extreme loading a gap may appear and this can influence sealperformance and ultimately fatigue response of the laminate and particularly the steel components.


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In this respect the combination of PEEK and carbon fiber is a key advantage. PEEK has a high creep resistance, even at

elevated temperatures, in comparison with other polymers. Also carbon fiber, which has a negligible creep response, further

stabilises the polymer.

In the current design the interface between the steel and composite materials is formed by relatively large accurately machined

and well toleranced surfaces allowing high local stresses to be avoided. This combined with the high section stiffnesses helps

to reduce the potential for creep and end fitting relaxation.

However, the design is also attractive from the point of view of thermal and external loading as it does not rely on bondedsurfaces and can accommodate some relative movement between the steel collar and the composite pipe. The highly preloaded

connection ensures that even when the end fitting is heated the resulting differential movement can be accommodated.

FEA APPROACH and RESULTS

The end fitting design has been subjected to a full FEA evaluation using ANSYS. Rather than using a complex composite

modelling process a simplified approach is used that quickly and efficiently provides results for both the composite andmetallic components. The analysis process uses material data generated by testing, primarily on full scale pipes but also some

coupon materials. This test data is also supplemented where necessary by specialist material design software to produce

characteristic orthotropic material properties.

The starting point for the analysis work was a series of small scale pull tests to investigate the response of taperedcomposite/metallic interface and understand the issues related to slip, friction factors, surface finish, creep and necking. This

also allowed benchmarking of ANSYS results against test data.

This preliminary testing was conducted using 2in 10,000psi rated bore pipe specimens, shown in Figure 5 which were axially

loaded and accurately monitored to understand their response. This provided confidence in the proposed interface and

provided confidence in selected friction factors, which are surface finish dependent.

Figure 5 End Fitting Pull Test Configuration

Subsequently, a full FE model of the end fitting was prepared that correctly models the interfaces between all the steel andcomposite components including the gasket. The end fitting was modelled axisymetrically using PLANE182 elements and

using a range of orthotropic properties to model the different sections and layers of the pipe and build-up.


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Figure 6 End Fitting Assembly

The geometry used for the analysis is a 3.25in 5,000psi end fitting. A 2D axisymmetric model is used, created by importing a

3D Parasolid file, generated in Solidworks CAD software, and refined within the ANSYS Workbench software ensuring thatthe geometry accurately reflects the detailed design and enables better mesh and analysis. The model is quarter sectioned and

small chamfers and bolt holes are removed.

The entire model is meshed using 4-noded plane182 elements, with linear elastic behaviour. The pipe and collar sections of the

model are divided into smaller areas such that a regular pattern of mesh is achieved. Mapped mesh is applied on these sections.

Other steel parts such as the Locking Ring, Steel Hub, and threaded part of the collar are meshed using the free meshingoption. The collar is also divided into smaller areas and free meshed such that the contact surfaces on the inner layer and the

steel have the same mesh pattern.

In the contact regions, the mesh is refined to achieve virtually matching mesh. The contact regions are modelled, using

TARGET171 and CONTA169 elements, as asymmetric frictional contacts.

The model contains 5 contact regions to simulate the locking taper friction, the pipe end abutment with the hub, the bore seal

behaviour and influence of elastomer (rubber) which is used as a compliant filler between the internal diameter of the steel

collar end transition to the composite pipe. This manages local contact stresses. The five areas are as follows:

1. Outer surface of the locking taper of the pipe and the corresponding inner surface of the collar, the friction factor isassumed 0.3.

2. Forward surface of the pipe and the aft face of the hub, friction factor 0.2.3. Top left surface of the bore seal and forward and bottom surfaces of the inner pocket layer, friction factor 0.2.4. Top right surface of the bore seal and bottom taper face of the hub, friction factor 0.4.5. Elastomer, between the thin end of the collar and composite build up, friction factor 0.6.

The FEA model is loaded using a range of load conditions including internal pressure, bending and axial load. The internalpressures used in the current study are relatively low but other studies have considered operating pressures of 15,000psi and

test pressures in excess of 45,000psi.

Model pressures are as follows:

Operational Pressure 3,103 psi (214 bar)

Test Pressure 5,555 psi (383 bar)

A PEEK bore seal is used. This has a small interference fit with the pocket and results in an initial contact pressure when the

end fitting is assembled. The seal is then pressure actuated as the bore pressure increases. In accordance with the findings ofprevious analyses and tests, the end fitting taper needs to be pre-loaded to minimise movement at the taper face when high

COLLAR

LOCKING

TAPER

HUB

LOCKING

RING

m-pipe

BORE SEAL


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external loads are applied to the end fitting. An axial pre-load of at least 1.3 times the maximum axial load (pressure end cap

plus applied load) is applied. Practically, this is achieved on the end fitting by using a hydraulic puller on the collar and then

holding this with a threaded locking ring. Once the locking ring has taken up the slack and has been tightened up against the

hub, the hydraulic puller is released. This process, using a hydraulic tool rather than a applying torque, ensures that an accuratepreload is applied without the need to make assumptions for friction factors or component deformations. In the FEA the

preload is achieved by thermally shrinking a ring of elements until the appropriate preload is achieved.

The material properties used, the contact definitions and loading applied are explained in the following sections.

Structural Loading and Boundary Conditions

The structural loading applied in the analysis is summarized in Table 1. The analysis is carried out in 2 load steps as below:

Load steps Load type Application type Quantity

1 Pre loadShrink a row of collar elements to achieve stresses of ~66

MPa in the collar by applying -100C

(34 tonnes)

2

Axial load

Operational Pressure Load23.9 MPa

(10.87 tonnes)

Test Pressure Load42.75 MPa

(19. 5 tonnes)

Internal

pressurePressure Load 38.3 MPa

Table 1 Design Load Cases

Figure 7 - Loading and Constraint

The results from the analyses of the model under loading, as described in the previous section, are summarised and illustrated

in the following tables and images.

The positions of the nodes 2356 & 2394 are within the pipe laminate build-up and are at the thick and thin ends of the taper,

Figure 8 Nodes 148 & 1015 are within the steel (collar) and at the thin and thick ends of the taper, local to the peak hoop

stresses indicated.


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Figure 8 Key Node Positions

Contact pressures at the locking taper are shown in Figure 9 for preload and preload plus pressure load. It can be seen that thenominal contact pressure load and distribution is relatively low along the majority of the contact face and that it does not

increase significantly as the internal pressure is applied. This is a result of the the preloading process.

Stresses in the steel collar and locking rings are relatively low even at the critical threaded collar sections.

a) b)

c)

Figure 9 - Contact pressures (MPa) at the taper. a)Preload b) Preload+Operational Pressure c) Preload+Test

Pressure

Node 2394Node 2356

Node 1015

Node 148

Node 277


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a)

b)

Figure 10 - Equivalent (Von Mises) stresses (MPa) after a) Preload+Operational Pressure b) Preload+Test Pressure

NODE Location Sz Sy Sxy Von Mises

2356 Pipe Thick End of taper 24% 29% 5% N/A

2394 Pipe Thin End of taper 25% 17% 0.2% N/A

148 Collar Thin End of taper 22%

1015 Collar Thick End of taper 19%

277 Collar in region of the locking taper 22%

Collar to Lock Ring thread roots. Secondary stress. 50%

Collar adjacent to Lock Ring take-up. Primary stress. 27%

Bore Seal Nib Root 17% 27% N/A

Bore Seal ID centreline 45% 30% N/A

Table 2 - Stresses after Load Step 1 (Preload) % Utilisation of Ultimate StrengthSz=radial, Sy=hoop, and Sx=axial.


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NODE LocationSz Sy Sxy VM

2356 Pipe Thick End of taper 24% 30% 6% N/A

2394 Pipe Thin End of taper 27% 17% 1% N/A

Pipe Laminate ID at start of build-up 28% N/A

Pipe laminate ID 18% 21% N/A

Pipe laminate OD 10% 21% N/A

148 Collar Thin End of taper 39%

1015 Collar Thick End of taper 26%

277 Collar in region of the locking taper 33%

Collar to Lock Ring thread roots. Secondary stress. 17%

Collar adjacent to Lock Ring take-up. Primary stress. 38%

Bore Seal Taper Root 9% 26% N/A

Bore Seal ID centreline 56% 68% N/ATable 3- Stresses after Load step 2 (Preload + Internal Pressure) % Utilisation of Ultimate Strength

Sz=radial, Sy=hoop, and Sx=axial.

LocationMax Contact Stress

(MPa)

Pipe seal pocket 5

Hub seal pocket 6

Table 4 Bore Seal ContactStresses after Load Step 1 (Preload)

LocationMax Contact Stress

(MPa)

Pipe seal pocket 50

Hub seal pocket 60 (Mean) 80 (Peak)

Table 5 Bore Seal Contact Stresses after Load Step 2 (Preload + Pressure)

Location Contact Stress (MPa)

Average over the contact face 30

Peak value at contact face edges 142

Table 6 Locking Taper ContactStresses after Load Step 1 (Preload)

Location Contact Stress (MPa)

Average over the contact face 40

Peak value at contact face edges 186

Table 7 Locking Taper Contact Stresses after Load Step 2 (Preload + Pressure)


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Testing Approach

A series of mechanical tests have been performed on the proposed end fitting. The objective of each test is to prove the

structural performance and seal integrity for a range of applied loads.

Pressure

The internal pressure test comprises a closed end pipe filled with water. The end of the pipe is terminated with a Magma endfitting and assembled with a PEEK bore seal. The pump allows accurate control of the pressure so that any form of test profilecan be conducted.

There are four typical test profiles:

1. Step test: pressure increased in steps to a target maximum and then released in steps2. Dwell test: valve locked at a target pressure and left for a period of days3. Cyclic test: pressure repeatedly increased and decreased in quick succession4. Failure: pressure increased in steps until failure

After each test the fittings are inspected for any evidence of leaks. The pressure data is analysed to check for changes in the

pressure profile. A concrete bunker ensures the safety of personel during all tests that operate near predicted failure pressures.

Individual loadsIn addition to the tensile tests performed to evaluate the metal-m-pipe interface (see FEA Approach and Results), two

individual load cases have been assessed: axial tensile load and cantilever bending load. The test approach includes multiplefibre optic strain gauges on the surface of the pipe in the axial and hoop direction. In addition, a video extensometer is used for

the tensile test to monitor the strains at any location along the pipe or end fitting. Calibrated load cells are linked to the

hydraulic cylinders that apply the loads.

Combined loads

The combined load test has been assembled to enable any combination of internal pressure, axial tension, and cantilever

bending loads to be applied. The following combinations were chosen:

1. Internal pressure and tensile load2. Internal pressure and bending load

3. Internal pressure, tensile and bending loads4. Failure: pressure and bending to failure

Figure 11 - Combined Load Rig GA

Fatigue

Fatigue has been evaluated using a resonant cantilever beam arrangement. A vibration motor with offset weights was secured

to one end of an m-pipe test sample with the other end terminated in a Magma end fitting bolted to a vertical bracket. Thepipe was instrumented with fibre optic strain gauges so that the oscillation frequency and strain response could be monitored at

all times. The pipe was filled with water and pressurised to a nominal level throughout the resonant fatigue tests.

m-pipe

End Fitting


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OTC 23977 15

Figure 12 - Resonant Fatigue Rig GA

Test Results

Pressure

A number of pipes were prepared over a range of two sizes: 2in and 3.25in. A summary of the key results is presented in Table

8.

m-pipe Test Target Pressure Test Description Result

2in 5000psi Step 11,000psi 2900psi steps, hold for 2 minutes

at each step

No sign of leak

Step 14,500psi 2900psi steps, hold for 2 minutes

at each step

No sign of leak

Dwell 14,500psi 2900psi steps, hold for 2 minutes

at each step. Dwell for 150 hours(~6 days)

No sign of leak

2in 10,000psi Failure - 2900psi steps, hold for 2 minutes

at each step continue untilleak/failue

Data from step dwells

indicated no sign of leaks.Pipe burst at 26,527psi

3.25in 3100psi Step 5555psi 1450psi steps, hold for 5 minutes

at each step

No sign of leak

Step 6600psi 1450psi steps, hold for 5 minutes

at each step

No sign of leak

Dwell 6600psi 1450psi steps, hold for 5 minutes

at each step. Dwell for 65 hours(~2.5 days)

No sign of leak

Cycle 5555psi Increase to 5555psi in 1 minute,

dwell for 1 minute, releasepressure in 1 minute. Repeat 10

times

No sign of leak

Table 8 - Pressure Test Results Summary

Cantilever End Fitting

Vibratory Motor


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16 OTC 23977

Each test sample was assembled and dismantled several times during the sequence of pressure tests showing that the seal and

the seal pocket remained undamaged.

Figure 13 - Photo of a 2in 5000psi m-pipe in the pressure chamber

Individual loads

The test results for the individual loads are shown in Table 9.

m-pipe Test Load Test Description Result

2in 5000psi Tensile 120kN 20kN steps, dwell for 1 minute at eachstep

No sign of axialslippage or damage

Cantilever Bend 0.9kNm 0.15kNm steps, dwell for 1 minute at

each step

No sign of damage

Table 9 - Individual Load Test Results

The tensile only test could have presented difficulties due to the hoop shrinkage as a result of the Poissons effect. The benefits

of the selected manufacturing method include the ability to vary the fibre angles in the design of the additional thickeningaround the fitting and so the Poissons effect did not cause any issues during this test.

Combined Loads

The combined load tests followed on from the individual load tests using the same equipment. The test plan consisted of a

matrix of target loads and pressures and in all cases the pipe axial and hoop strains were monitored. The key results are

summarised in Table 10.

TestTarget

Load/PressureTest Description Result

Tensile &Pressure

112kN,10,000psi

20kN steps at 4 different levels of internalpressure, dwell for 1 minute at each step

No sign of axial slippageor leaks

Cantilever Bend& Pressure

0.9kNm,10,000psi

0.2kNm steps at 4 different levels of internalpressure, dwell for 1 minute at each step

No sign of damage orleaks

Tensile,Cantilever Bend

& Pressure

40kN,0.6kNm,

10,000psi

2 tensile load steps, 2 levels of internalpressure, 4 steps of increasing bending

moment

No sign of damage orleaks

Cantilever Bend

& Pressure to

failure

10,000psi Increase bending moment until failure 150mm stroke limit

reached on the hydraulic

cylinder no failure

Table 10 - Combined Load Test Results (2in 5000psi Rated)


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These tests present themselves as a fairly rigorous series of events with axial pipe strains in excess of 1.8%. The combined

bending and pressure test highlighted a beneficial aspect of m-pipe in that the axial strain due to pressure had to be over-

come by the compressive strain due to bending before failure in the pipe could occur. In this case the limit of the vertical

hydraulic cylinder was reached and so the test was repeated without internal pressure so that a failure could be achieved. Thefailure occurred on the top surface of the pipe (compressive load) near the exit of the end fitting.

Figure 14 - Combined Load Test Rig

Figure 15 - Bend to Failure

Fatigue

A series of ring-down tests were performed to establish the resonant frequency of the system. This was quite simply achieved

by knocking the end of the pipe with a rubber mallet and recording the axial strains in the fibre optic sensors. The time-strain

plot gives the damping characteristics and an indication of the natural frequency.

The motor weights were initially balanced to give 75% of their maximum vibrational force. An internal pressure of 1000psi

was applied using a water filled pump. The motor control was increased until a frequency close to the natural frequency wasachieved (~10Hz). The motor speed was adjusted to give a target peak-peak strain value in the fibre optic strain sensors that

were positioned in the middle of the pipe. Assuming the bending moment increases linearly towards the fixed end of the pipe

the strain was calculated at the point the pipe exits the end fitting and this extrapolated value is shown in Table 11.

For this particular test the end fitting setup represented an early design of the bore seal and as such does not contain some of

the design improvements from subsequent pressure testing.

The results are summarised in Table 11.


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Strain level at End Fitting/Pipe Interface Number of cycles Result

~0.8% strain peak-peak 2.4millionNo leak,

No structural damage

1.1% strain peak-peak 10.0million

Small leak observed

from seal at 7.8million

cycles,


~1.45% strain peak-peak 0.5million


unable to reach higher

strains due to motorlimit.

Table 11 - Resonant Fatigue Test Results

Further resonant fatigue testing is planned with 2in 15,000psi m-pipe samples and 3in 10,000psi samples to increase the

sample data.

Conclusion

It is believed that carbon fibre/PEEK pipe has the potential to be an enabler for future oil and gas developments in applications

where metallic based pipe technologies struggle to meet project requirements. However, composite pipe end fitting design

issues have often been highlighted as a concern, of sufficient magnitude, to preclude serious consideration of the technology.These concerns have related to structural performance, sealability and general long term reliability of composite materials and

specifically composite/metallic interfaces.

The work reported evidences carbon fibre/PEEK materials and manufacturing methods that together allow a different endfitting approach to be adopted. This approach resolves previous issues facilitating a solution with increased performance and

reliability. Importantly, this removes a serious obstacle to the future broarder application of composite pipe technology.

The key design features of the end fitting design are enabled by the manufacturing method and the selection of carbon fibre

and PEEK as the composite materials of choice. This process allows the end of the pipe to be readily thickened as a secondary

manufacturing operation to provide a reliable structural interface with the steel end fitting. Additionally, the end fitting design

specifically manages the issues related to differential thermal expansion and Poisson ratio effects between metallic and

composite elements. The combination of these features is that the end fitting can be designed to stronger than the base pipe inall respects.

References

[1] Melve, B. First Offshore Composite Riser Joint Proven on Heidrun Offshore Magazine

[2] Blanc, L. Composites Cut the Riser Weight by 30-40%, Mass 20-30% Offshore Magazine

[3] Baldwin, D.Interface System Bewteen Composite Tubing and End Fittings United States Patent 6,042,152

[4] Guesnon, J. Caillard C. Hybrid Tubes for Choke and Kill Lines, Offshore Technology Conference OTC 14021

[5] Cederberg, C.Design and Verification Testing Compsoite-Reinforced Steel Drilling Riser Final Report RPSEA

07121-1401

[oc] end fittings for composite risers

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