1. An integrated approach to the design of
high performance carbon fibre reinforced
risers - from micro to macro - scale
Angelos Mintzas1
, Steve Hatton1
, Sarinova Simandjuntak2
,
Andrew Little2
, Zhongyi Zhang2
Magma Global Ltd1
University of Portsmouth2
ABSTRACT
The increasing demand for ultra-deep water oil and gas extraction has led to the consideration
of high performance materials, such as carbon fibre reinforced composites, in the
manufacturing of risers. Whilst the unique properties of composites offer great capability in
the design, the lack of field experience and of relevant design codes and standards raise
concerns regarding appropriate design methodologies within the oil and gas community.
In this study, an integrated design approach is proposed, whereby information from the
micro-scale (i.e. at the fibre and matrix level) is used for the prediction of the structural
response at the macro-scale (i.e. of the whole pipe). In order to do so, mathematical models
describing the pipe behaviour at the micro, meso and macro-scales are developed and linked
together. Experiments in all three length scales are performed in order to reveal the failure
mechanisms associated with different loading conditions and to validate the proposed
method. Predicted and experimental results from pipes subjected to internal pressure, tensile,
compressive and bending loads as well as their combinations are presented and discussed.
INTRODUCTION
Composites are nowadays extensively used in automobile, marine, civil and aerospace
industries with the latter being the main conductor of industrial level composite technology.
Weight saving in aerospace has been a critical selection driver and the excellent strength to
weight ratio of fibrous composites renders them ideal for lightweight applications. The
significantly reduced weight of composite structures, when compared to ‘conventional’ steel,
aluminium and titanium ones, is not, however, the only advantage. Due to their
microstructure, composites are inherently fatigue resistant and can be tailored to suit the
particular application. For example, a structure with high corrosion resistance to seawater,
CO2 and H2S which has also the capability of operating at high strain levels (up to 1%
operational strain) can be manufactured with the appropriate choice of matrix, fibre type and

2. orientation. It is therefore apparent that the utilization of high performance composites such
as PEEK / carbon fibre, can potentially improve riser safety, reliability and field viability
while at the same time reduce deployment and operational costs. Moreover, such materials
will ultimately enable the industry to unlock the potential of ultra-deep water applications
where conventional steel and non-bonded flexible risers have reached their operational limits.
The oil and gas industry has identified the benefits of using composites as a potential offshore
riser material for over 20 years. However, their utilization has been mostly limited to non-
bonded flexible risers where the heavy, costly and fatigue sensitive steel armour wires are
replaced by carbon fibre reinforced ones. Some of the reasons for the lack of composite
applications in this industry include the lack of familiarity of composite technology, field
experience, applicable test data and established design codes. A thorough understanding of
the composite behaviour is therefore important in order to help address the aforementioned
deficiencies. The development of a multi-scale design approach, whereby information at the
micro-level (i.e. μm scale) is used for the prediction of the structural response at the macro-
level (i.e. m scale), is fundamental for the understanding of the complex failure mechanisms
that develop within composite materials and thus for the prediction of their behaviour.
MODELLING STRATEGY
A graphical illustration of the pipe and its break-down into the three different length scales is
shown in Figure 1. It should be noted here that unlike non-bonded flexible pipes, the
composite pipes in this study are monolithic structures. The structure break-down into
smaller components is therefore purely done for analysis purposes and does not correspond to
any physical discontinuity. The pipe consists of a number of repeatable fibre groupings that
are laid along predefined angles (θ) with the fibres being held together by a PEEK matrix
fused in via a laser assisted thermoplastic process1
. Each of these groupings can be treated as
the building block (meso-scale model) of the pipe. Consequently, each fibre grouping
consists of thousands of individual fibres surrounded by PEEK matrix which can be now
treated as the building block (micro-scale model) of the different fibre groupings. The
response of the final pipe structure can be therefore predicted by merely using the behaviour
of the constituent materials (i.e. carbon fibre and PEEK matrix) as an input. This provides a
great flexibility to the design as pipes with any matrix / fibre combination and fibre
orientation can be analysed. The predicting capability of this approach clearly depends on the
proximity of the assumptions made at each length scale which can be, however, quantified by
performing experiments on all three length scales.
1
This process is completely different to the usual thermoset processes as typically employed by
aerospace industry and it is more controllable, repeatable and recordable.

3. MICRO SCALE MODELLING
The random array of fibres within the PEEK matrix can be idealized as a periodic array of
repetitive unit cells as shown in Figure 2. The response of a fibre grouping can be therefore
defined by simulating a unit cell with periodic boundary conditions [1]. Due to the shape of
the fibre, it becomes apparent that the response of the unit cell will be directionally dependant
i.e. the overall material behaviour will be anisotropic. Five elastic constants are required for
the characterization of a fibre grouping with fibres lying along the same direction [2]. These
are
 the elastic modulus along the fibre directions (E1)
 the elastic modulus transverse to the fibre direction (E2)
 Poisson’s ratio in 1-2 plane (v12)
 the shear modulus in 1-2 plane (G12)
 the shear modulus in 2-3 plane (G23)
E1 and v12 can be derived from the unit cell by applying a tensile load along the 1-direction
and then obtaining the resulting displacements along the 1- and 2-directions, E2 by applying a
tensile load along 2-direction and obtaining the resulting displacement, G12 by applying a
shear load and obtaining the change in 1 - 2 angle and G23 by applying a shear load and
obtaining the change in 2 - 3 angle. The only input required for the simulations above is the
volume fraction of either the matrix or the fibre and their constitutive behaviours.
MESO AND MACRO SCALE MODELLING
At the meso-scale level, the material is treated as a continuum having the anisotropic
properties calculated from the micro-scale analysis. However, these properties correspond to
the principal coordinate system of each fibre grouping (i.e. 1-2-3 coordinate system in Figure
1) and can be used as calculated only when the latter coincides with the global coordinate
system (i.e. only if fibres are laid along θ = 90o
). The material properties of the fibre
groupings laid at angles θ ≠ 90o
, can be mathematically derived through a stiffness matrix
transformation [3]. The pipe can be therefore treated as a layered structure with each layer
corresponding to a group of fibres with the same direction.
The layered version of element 186 in the ANSYS finite element software has the capability
of modelling up to 1000 different such layers within the same element. This allows for the
calculation of the stresses and strains at the meso-scale level while performing a single global
analysis (i.e. on the whole pipe geometry) within reasonable computational time. Obtaining
the stresses and strains at the meso-scale level is very important as most of the existing failure
criteria for composites [4, 5, 6] are only applicable at this level. In this study, a maximum
stress allowable along the fibre direction2
was employed for the prediction of the pipe failure
2
Due to their microstructure, composites have less strength when subjected to compressive loads. The
allowable tensile stress was obtained from testing unidirectional coupons with fibres along the loading
direction. The allowable compressive stress was taken to be within 0.4 to 0.6 of the tensile one.

4. load with its value obtained from the unidirectional coupon tests described in the next
section.
EXPERIMENTAL PROGRAMME
PEEK AND CARBON FIBRE TESTS
The material properties of PEEK and carbon fibre have been provided by the manufacturers
i.e. Victrex and Toray respectively and are given in Table 1. It should be noted here that
carbon fibres exhibit transversely isotropic behaviour hence the 5 elastic constants.
UNI-DIRECTIONAL COUPON TESTS
For the experimental validation of the material properties obtained from the micro-scale
model, two different coupon types have been tested. The first type is a cylindrical coupon
with fibres laid along +45o
and -45o
angles as shown in Figure 3(a). Flat ±45o
coupon testing
[7] is the most widely used method for obtaining the shear modulus G12 and the shear
strength S12 of fibre reinforced composites. A cylindrical coupon is, however, preferred in
this study because (a) it is loyal to the manufacturing procedure, (b) it can be easily tested in
compression3
and (c) there are no edge effects to contaminate the test results. The second
type is a flat coupon with fibres laid along the X-direction as shown in Figure 3(b). By
applying a tensile load and monitoring the longitudinal and transverse strains, the elastic
properties E1, v12 and the tensile strength along the fibre direction X11T can be obtained.
Form the two tests above, three (i.e. E1, v12 and G12) out of the five elastic constants needed
for the meso-scale analysis have been obtained. Tests for the determination of E2 are yet to be
performed. However, E2 corresponds to the modulus transverse to the fibre and is a matrix
dominated property. Composite structures are designed to carry loads along the fibre
direction and therefore small variations in the transverse properties do not significantly affect
the load distribution within the structure. Tests for the determination of G23 are difficult to
perform and give high scatter. For this reason, DNV OS C501 code [8] suggests that G23 is
taken equal to G12.
COMPONENT TESTS
Two 2” 10 ksi rated pipe configurations have been tested, one designed to be flexible and one
to have higher axial capacity. The flexible pipe consists of fibre groupings laid along two
directions θ = a and θ = b where a, b ≤ 45o
. The same a, b angles are used in the axially
3
Shear behaviour of ± 45 composites is similar in tension and compression and thus shear properties
can be defined via either tension or compression tests as indicated in DNV OS C501 [8]. The
cylindrical shape of the designed coupon prevents it from buckling thus the latter can be easily loaded
up to failure.

5. stiffer pipe with extra fibres laid along angle θ = c where c > 45o
. The characteristics of the
two pipes are given in Table 2.
For the qualification of the two pipe configurations, an extensive experimental testing
programme has been completed which includes axial tension, axial compression, burst,
collapse, bending and combined bending - burst tests. For the axial tension and compression a
Tinius Olsen machine was used along with a video extensometer to monitor the strains. For
the burst tests special end fittings were designed to simulate open-ended conditions and a
video extensometer was used to monitor the strains at low pressure levels. For the bending
and the combined bending - burst tests the purpose built rig shown in Figure 4 was used
along with optical strain gages. Accurate strain data are not available for collapse therefore
this test could not be used for the validation of the modelling method. A summary of the type
and number of tests used for the model validation is given in Table 3.
RESULTS AND DISCUSSION
Tables with the predicted and experimental properties along with plots of the stress-strain
response as obtained from analysis, coupon and pipe testing are given below.
MICRO SCALE MODEL VALIDATION
The number of coupon tests, the coefficient of variation and the mean experimental E1, v12
and X11T values normalised by the predicted ones are given in Table 4. Predicted and
experimental elastic properties are shown to be in very good agreement (within 4%).
However, the predicted tensile strength is shown to be 26% higher than the experimental one.
This is not surprising since this property is sensitive to the coupon manufacturing process and
the test procedure. Having a benchmark value for the maximum achievable strength is very
useful in assessing the quality of (a) the supplied materials, (b) the manufacturing process and
(c) the experimental procedure.
The mean experimental shear properties normalized by the predicted ones are given in Table
5. Predicted and experimental shear strengths and moduli are shown to be in extremely good
agreement. This is because shear in 1-2 plane (see Figure 2) is a matrix dominated property.
Therefore, provided the shear behaviour of PEEK is known, the shear response of the
unidirectional carbon-PEEK composite can be accurately predicted. Parameters such as the
coupon manufacturing process and the experimental procedure do not significantly affect
shear strength. This is most probably due to the highly non-linear behaviour of PEEK which
was also evident in the nearly elastic - perfectly plastic response of the ± 45o
coupons.
MACRO SCALE MODEL VALIDATION
Typical axial strain and hoop strain - stress curves as obtained from pipe compression tests
and FE analyses are shown in Figure 5 and Figure 6 respectively. In both plots, the stresses
are normalized against the predicted compressive failure stress of each pipe configuration.
From Figure 5, it can be seen that FE overestimates the axial stiffness of both pipes by

6. approximately 15% to 20%. At normalized stresses greater than 0.5, this difference is partly
due to the non-linear axial deformation of the pipe which is not taken into account in the
current analysis. This does not, however, explain the differences in the linear region. FE gives
much better predictions in the hoop direction, where, predicted and experimental stress -
strain curves are shown to be in good agreement (see Figure 6). The failure stress of the (a /
b) pipe is over-predicted by 21%, whereas, that of the (a / b / c) pipe is under-predicted by
5%.
Typical axial strain and hoop strain - stress curves as obtained from pipe tension tests and FE
analyses are shown in Figure 7 and Figure 8 respectively. The plotted stresses are normalized
against the predicted tensile failure stress of each pipe configuration. These tests were run up
to approximately 0.1 of the ultimate failure load to ensure loading within the linear region
(could not be loaded to failure due to clamping restrictions). Similarly to the compressive
tests, the axial stiffness is over-predicted in this case by approximately 10% to 15%, whereas,
the hoop stiffness is once again accurately predicted.
The variations of hoop strain with normalised internal pressure as obtained from the burst
tests are plotted in Figure 9. Strains were monitored for pressures up to 7 % and 16% of the
pressure capacity of the (a / b) and the (a / b / c) pipe respectively. In both cases, the
predicted and experimental hoop strains are shown to be in very good agreement. After the
low pressure tests, the video extensometer was removed and both pipes were taken to failure.
The ratio of the experimental to the predicted burst pressure is 1.01 for the (a / b) pipe and
1.16 for the (a / b / c) pipe.
It should be noted here that all the input properties used for the FE analyses of the pipe come
directly from the micro-scale analysis without using any fit factors. This is because the latter
were found to be in very good agreement with the ones obtained from unidirectional coupon
testing. The only property taken from experimental data is the tensile strength along the fibre
direction which was determined from tests on unidirectional coupons that have fibres along
the loading direction. This is because the latter property was overestimated from the micro-
scale analysis by 26% (see Table 4).
The axial tension and compression tests revealed that the pipe axial stiffness is over-predicted
by 10% to 20%. An axial stiffness knock down factor of 20% was therefore applied on the FE
analyses for the prediction of the (a / b) pipe subjected to bending and combined internal
pressure - bending loads. The axial strain versus the normalized bending moment for the pure
bending test is shown in Figure 10. The structure was bent in the linear region and the FE
results are shown to give good predictions. The experimental axial strains were measured in
both tension and compression sides of the pipe to ensure that pure bending load was applied,
whereas, the hoop strains were very close to zero and are thus not plotted. Figure 11 shows
the axial and hoop strain versus the normalized bending moment for the (a / b) pipe subjected
to bending and 690 bar internal pressure. It should be noted that the bending moment is
normalised with the bending moment at failure as predicted for a pipe with zero internal
pressure. Internal pressure relieves the axial stresses (due to the contraction in the axial
direction) hence the higher bending capacity of the pipe (i.e. bending moment / bending

7. moment at failure = 1.5 in Figure 11). Experimental and FE predictions are shown to be in
very good agreement.
The tests show that the hoop strain - stress response is more accurately predicted than the
corresponding axial one for both (a / b) and (a / b / c) pipe configurations (both pipes have
more fibres along the hoop direction). Moreover, the strength predictions are more accurate
when the pipes are loaded along the fibre direction (i.e. better strength predictions for (a / b /
c) pipe when axially loaded and better predictions for the (a / b) pipe when tested in burst).
The above indicate that the current modelling approach is more efficient in predicting the
pipe response along the direction that is more heavily reinforced by fibres. This is not
surprising given that the failure criterion employed is based on a maximum allowable tensile
and compressive strength along the fibre direction. Coupon experiments for the determination
of the transverse to the fibre properties are expected to help in the better understanding of the
axial response of the pipe configurations studied.
CONCLUSIONS
A multi-scale modelling approach has been used in this study for the prediction of the
response of composite pipes subjected to various loading conditions. This approach allows
for the calculation of the stress / strain fields at different length scales within the structure and
thus for the application of physically based failure criteria. The latter can be linked to the
failure modes found in composites and can thus yield more accurate strength predictions. The
maximum allowable tensile and compressive stress along the fibre direction was the failure
criterion employed for the strength predictions in this study. Despite its simplicity, this
criterion was shown to provide acceptable predictions especially for the cases where the pipe
was loaded along the more highly reinforced directions. More sophisticated constitutive
models (e.g. ones that can account for material non-linearity) and failure criteria must be
employed in order to obtain more accurate predictions. The latter can be determined from
micro-scale analyses combined with coupon testing.
ACKNOWLEDGEMETS
A portion of this work was carried out within the framework of the UK Knowledge Transfer
Partnership scheme and the support of University of Portsmouth is gratefully acknowledged.

8. NOMENCLATURE
COV Coefficient of Variation
FE Finite Element
MBR Minimum Bending Radius
REFERENCES
[1] Sun C. T. & Vaidya R.S., Prediction of Composite Properties from a Representative Volume
Element. Comp. Sci. & Tech., 56 (1996) 171-179.
[2] Hull D. & Clyne T.W., An Introduction to Composite Materials. Cambridge University Press,
(1996) 2nd
Edition.
[3] S. G. Lekhnitski, Theory of Elasticity of an Anisotropic Elastic Body, San Francisco, Holden-
Day, 1963.
[4] Tsai, S. W. & Wu, E. M., A General Theory of Strength for Anisotropic Materials. Journal of
Composite Materials., 5 (1971) 58–80.
[5] Puck, A. & Schurmann, H., Failure Analysis of FRP Laminates by Means of Physically Based
Phenomenological Models, Composites Science and Technology, 62 (2002) 1633–1662.
[6] Hashin, Z., Failure Criteria for Unidirectional Fiber Composites. J. Appl. Mech., 47 (1980) 329–
334.
[7] ASTM D 3518 / D3518M - 94(2007) Standard Test Method for In-Plane Shear Response of
Polymer Matrix Composite Materials by Tensile Test of a ±45° Laminate.
[8] Det Norske Veritas –“Composite Components”, Offshore Standard DNV-OS-C501, October
2010.
TABLES AND FIGURES
Table 1 - Material properties of PEEK and carbon fibres as provided by Victrex and
Toray respectively
Material E11
(GPa)
E22
(GPa)
G12
(GPa)
G23
(GPa)
v12 X11T
(MPa)
PEEK 3.7 3.7 1.36 1.36 0.36 110
Carbon
fibre
230 20.9 27.6 27.6* 0.2 4900
*G23 is not provided by the manufacturer and is therefore assumed to be equal to G12

9. Table 2 – Characteristics of the 2” 10 ksi rated pipe configurations tested
Pipe
Configuration
Pipe characteristics
Inner
Diameter
(mm)
Outer
Diameter
(mm)
Wall
Thickness
MBR
(mm)
10 ksi composite
pipe weight / 10 ksi
steel pipe weight
(a / b) pipe 50.8 63.6 6.4 4600 0.14
(a / b / c) pipe 50.8 62.2 5.7 7000 0.13
Table 3 – Type and number of tests performed for model validation
Pipe
Configuration
Number of tests for model validation
Tension Compression Burst Bending Burst-bending
(a / b) pipe 5 5 5 1 1
(a / b / c) pipe 5 5 5 - -
Table 4 - Experimental and predicted properties along the fibre direction
No. of
coupons
tested
Property COV Mean
Experimental
/ Predicted
23
E1 6.3% 0.96
v12 13.8% 0.98
X11T 7.2% 0.74
Table 5 – Experimental and predicted shear properties in 1-2 plane
No. of
coupons
tested
Property COV Mean
Experimental
/ Predicted
5
G12 5.1% 1.01
S12 1.3% 1.02

10. Figure 1 – Pipe break-down to smaller components for modelling purposes

11. Figure 2 – Idealisation of a random array of fibres as a periodic array consisting of
repetitive unit cells
Figure 3 – Unidirectional coupons for the determination of (a) the shear properties (i.e.
G12 and S12) and (b) the properties along the fibre direction (i.e. E11, v12, X11T)

12. Figure 4 – Combined burst and bend rig assembly model
Figure 5 – Axial strain versus normalised stress as obtained from pipe compressive tests
and FE predictions. Stresses are normalised against the predicted failure stress of each
pipe configuration

13. Figure 6 – Hoop strain versus normalised stress as obtained from pipe compressive tests
and FE predictions. Stresses are normalised against the predicted failure stress of each
pipe configuration
Figure 7– Axial strain versus normalised stress as obtained from pipe tensile tests and
FE predictions. Stresses are normalised against the predicted failure stress of each pipe
configuration

14. Figure 8– Hoop strain versus normalised stress as obtained from pipe tensile tests and
FE predictions. Stresses are normalised against the predicted failure stress of each pipe
configuration
Figure 9 – Hoop strain versus normalised stress as obtained from internal pressure tests
and FE predictions. Stresses are normalised against the predicted failure pressure of
each pipe configuration

15. Figure 10 – Axial strain versus normalised bending moment as obtained from the
bending test on (a / b) pipe configuration and the FE predictions. Bending moment is
normalised against the predicted bending moment at failure.

16. Figure 11 – Axial and hoop strains versus normalised bending moment as obtained
from the combined internal pressure - bending test on the (a / b) pipe configuration and
the FE predictions. Bending moment is normalised against the predicted bending
moment at failure of a pipe with zero internal pressure.