1. OTC-25034-MS
Influence of Local Effects on Fatigue Assessments during Hold Periods of
Offshore Pipeline Installation by Reeling
Arek Bedrossian and Paul Brett, KW Subsea
Copyright 2014, Offshore Technology Conference
This paper was prepared for presentation at the Offshore Technology Conference Asia held in Kuala Lumpur, Malaysia, 25–28 March 2014.
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 been
reviewed 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, its
officers, 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 to
reproduce 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.
Abstract
There are many situations during offshore installation of pipelines in which the pipe is held clamped for extended periods on
the lay vessel. In the case of installation by reeling this is done by frictional grip at the hang off clamp (HOC) below the
workstation of the reel-lay vessel. The increased fatigue damage in the pipe over the duration of the hold period, especially at
or just below the HOC, then has to be evaluated and included in the overall fatigue assessment.
The behaviour of pipelines during offshore installation under various anticipated sea states have long been routinely predicted
using established and mostly purpose-written numerical methods. In these assessments the time histories of vessel movement
and wave and current action on the immersed pipe are used as input and the dynamic response of the pipeline is evaluated. The
higher local fatigue damage at the girth weld nearest the clamp is then evaluated by effectively using a stress concentration
factor (SCF) applied to the predicted stress ranges at that location to take into account the local fixity. In most simulations,
however, the SCF is implicitly assigned an unnecessarily high value based on the assumption of full fixity at the clamp which
usually results in the predicted fatigue damage being significantly higher than that at the sagbend. The allowable hold period
for a given sea state could then be governed by the SCF.
In reality, many different factors would be at play at the HOC which reduce the magnitude of an SCF based on full fixity. The
more important of these factors are the local frictional slip between pipe outer surface and the pads, the thickness dependent
higher compliance of the outer thermal insulation coating material that may be on the pipe, and the local flexibilities of the
supporting platework of the HOC housing assembly on the vessel. This paper provides a discussion and semi-numerical
evaluation of the effects of some of these factors on the SCF and consequent assessment of fatigue damage. The method of
calculating the local compliance effects and using these in subsequent dynamic simulations is illustrated for a notional case.
Introduction
Almost all offshore pipeline installation involves, at some stage in the operations, an extended period of stoppage over which
pipe is held clamped on the vessel. This is to allow various pipe cutting and welding operations to take place on the vessel
above the point where pipe is held. The holding force varies according to the length of pipe catenary, pipe dimensions and
contents, and installation method. The holding capacity of the vessel is normally one of the factors limiting installation in
deeper waters. On reel lay vessels the clamping is achieved exclusively by frictional grip, with the lateral gripping force
applied through sets of pads distributed around and along the pipe outer surface. The pads are normally arranged into lower
and upper housing modules to provide a degree of operational flexibility and reliability. The assembly of the pads and
supporting framework of lever arm and hinge mechanisms is referred to as the hang of clamp (HOC) module. Part of a typical
HOC assembly is shown in Figure 1. It should be noted that many variants exist, and the materials and surface finish of the
pads and the details of the arrangements for attaching the pads could also vary to suit a particular operation on a given vessel.

2. 2 OTC-25034-MS
Figure 1: Part of a Typical HOC module
The frictional grip is in essence a mechanism for transferring the vertical load to horizontal force through frictional shear. As
with all such engineering situations where friction plays a key role, the load transfer takes place over a finite length of the
contacting surfaces. For typical values of friction coefficient a small amount of relative slip between the contacting surfaces
therefore exists ‘upstream’ of the load transfer. This in the context of the HOC is at the bottom of the lowest pad in the
assembly where the vertical tensile load in the pipe begins to be transferred by frictional shear to the pads upon application of
the horizontal grip forces to the pads. As the load transfer is fully accomplished over a short length ‘downstream’ along the
pads, no further local slip takes place and pipe is then firmly gripped. This situation would be the steady state case under pure
static tensile load of self-weight of the pipe. Global bending of pipe in addition to the steady tensile load disturbs this static
equilibrium state and produces additional local slip on both the tensile and compressive sides of the bend.
While pipe is held at the HOC, vessel motion and wave action on the pipe catenary induce cyclic global bending of the pipe
involving the dynamics of its entire length. The resulting cyclic stresses in the pipe could cause significant cumulative fatigue
damage especially under extreme installation sea states. Fatigue damage needs to be kept below ‘budgeted’ levels. The
detailed cyclic and dynamic behaviour of a pipeline held on an installation vessel is therefore routinely analysed at design
stage using numerical dynamics methods to predict the cumulative fatigue damage at the girth welds of the pipe. The two
critical locations along a pipeline with relatively high fatigue utilizations are normally the top of the pipe just below the HOC
and the sagbend. However, as in almost all numerical simulations of installation hold periods the connection between pipe and
vessel is effectively assumed to be infinitely stiff in bending, fatigue damage even metres below the HOC may be over-
conservatively predicted. Depending on the location of the nearest girth weld below the HOC, the conservatism could dictate
the allowable hold period. The level of conservatism associated with the assumption of infinite stiffness of the pipe-vessel
connection has not hitherto been evaluated or documented in guidelines.
The aim in this paper is to provide an insight into the influence of local effects at the pipe-vessel connection of a HOC on
predictions of dynamic behaviour of a pipe held at the HOC in typical sea state conditions. The effects of local frictional slip
and other local compliances of pads etc. are presented in a quantitative manner, whilst other effects, such as stiffness of the
HOC supporting platework, are discussed at a qualitative level.
Scope
The study presented in this paper is based entirely on numerical simulation. The local effects are first evaluated using 3-D
solid finite element (FE) models set up in ABAQUS (Ref.1) and the findings are subsequently used in 3-D Orcaflex (Ref.2) for
the calculation of the dynamic behaviour of a typical deepwater pipe installation. The method of transferring ABAQUS results
to Orcaflex is described. A ‘bare’ pipe with minimum corrosion coating and a thermally insulated (wet insulation) variant have
both been considered. The work described in this paper is, however, generic in nature and the postulated geometric and
material properties, friction coefficients, vessel characteristics, water depths and sea states are notional. No systematic
sensitivities with respect of any of these parameters have been undertaken. This is in essence a necessary and pragmatic
compromise as the actual range in pad properties and configurations, effective friction coefficients, vessel types, pipe and wet
insulation properties and dimensions, catenary lengths, and sea states, is quite unwieldy. The intention is to identify salient

3. OTC-25034-MS 3
features and provide insight into the influence of key physical effects and not to present a comprehensive parametric study.
The latter could be carried out for specific cases following the approach presented in this paper.
3-D FE Calculations
The basic components of the FE model of an insulated pipe used in the study are illustrated in Figure 2.
Figure 2: FE Model of Thermally Insulated Pipe
The metal pipe was modelled using the 3-D 4-node S4R shell elements in ABAQUS at pipe mean diameter. The shell elements
are shown in rendered form in Figure 2; hence the overlap of colours. The pads and wet insulation were modelled using the 8-
node 3-D solid continuum C3D8R elements in ABAQUS. In the case of the ‘bare’ pipe the insulation elements were removed
and the pads were brought closer to contact the pipe. Frictional contact interaction was modelled using the surface-to-surface
contact facility within ABAQUS. Geometric and material properties used in the FE calculations are summarized in Table 1.
Parameter Value Units
Metal Pipe Outside Diameter 273.1 mm
Metal Pipe Wall Thickness 15.0 mm
Elastic Modulus of Pipe Steel 207 GPa
Poisson’s Ratio of Pipe Steel 0.3 -
Density of Pipe Steel 7850 kg/m3
Thickness of Thermal Insulation 50.0 mm
Elastic Modulus of Thermal Insulation 2.5 GPa
Poisson’s Ratio of Thermal Insulation 0.45 -
Density of Thermal Insulation 900 kg/m3
Thickness of Pads (4 off around circumference) 20 mm
Effective Elastic Modulus of Pad Material 20 GPa
Poisson’s Ratio of Pad Material 0.45 -
Density of Pad Material 1200 kg/m3
Pad-to-Pipe Outer Surface Friction Coefficient 0.1 -
Table 1: Parameters used in the Pipe FE Models

4. 4 OTC-25034-MS
It should be emphasized that the parameters shown in Table 1 are entirely notional and do not represent any specific real case.
The chosen values of pipe and insulation properties were considered to be typical of deep water reeled pipe. The pad properties
are gross simplifications of the complex local compliances and diversity of materials that could be used for pads and backing
blocks. The elastic modulus assigned to the pads should be regarded as being a composite of all these compliances. The
friction coefficient of 0.1 between the pads and pipe outer surfaces was assumed not to vary with level of pad compression and
was also used for the bare pipe. In reality, all parameters used will vary and differ from the values used, but it is believed that
the basic principles highlighted in the paper and the method of analysis presented will still be valid.
The first three steps in the FE analyses were used to establish the frictional force balance between the steady tensile vertical
load of 64 Te applied to the bottom of the pipe (position ‘L’ in Figure 2(b)), representing the assumed weight of pipe catenary
below the HOC, and the lateral compressive forces applied to the pads. At the end of the 3rd
step the weight of pipe catenary
was taken up fully by the gripping forces applied to the pads and the friction between pad and pipe outer surfaces.
In the subsequent quasi-static steps alternating rotations of ± 0.05 radians were prescribed in small increments about the global
y-axis at the location ‘L’ of Figure 2(b) whilst maintaining the vertical tensile load constant. The pad outer surfaces were held
rigidly fixed in space in all three global directions. The calculated local rotation response of the pipe at the point ‘R’ of Figure
2(c) was then recorded, and plotted against the moment reaction obtained at location ‘L’. It should be noted that pipe rotations
about global y-axis at the pipe section containing position ‘R’ vary somewhat around the pipe circumference due to ovalisation
of pipe section and pad compliance. The rotations at ‘R’ were found to be representative of the mean values of rotation of the
whole local pipe section. The short length of modelled pipe and small prescribed rotations ensured that additional moment
induced by the tensile load at the section of point ‘R’ was negligibly small. The tensile load was applied as a ‘follower’ load
rotating in direction with the rotation of the pipe. The non-linear geometry, or large deformation, option was invoked, and the
calculations were performed using quasi-static force equilibrium principles.
The calculated moment-rotation responses of the bare and insulated pipes are shown in Figure 3.
Figure 3: Variation of Rotation of Pipe at Bottom of HOC with Applied Bending Moment
It can be seen that the frictional hysteresis effect in the bare pipe is more pronounced in comparison with that of the insulated
pipe. This is believed to be the result of the additional compliance of the thermal insulation, which through its higher
compressibility allows higher reduction in contact forces on the tensile side of the bend, thus producing less frictional
dissipation of energy. The effect could also be due to the method used in ABAQUS to apply pad compression but has not been
investigated further.

5. OTC-25034-MS 5
The differing values of frictional damping were calculated by releasing the last prescribed rotation and allowing pipe to vibrate
freely in a subsequent dynamic step in the ABAQUS model. The large deformation direct integration algorithm of the code
was invoked. The resulting calculated decay in the free vibration amplitudes of the two pipe systems is shown in Figure 4.
Figure 4: Predicted Free Vibrations of the Pipes
It can be seen that the frictional damping in the bare pipe is indeed higher than that in the insulated pipe. By considering the
logarithmic decrement, the critical damping ratios have been estimated to be 5.3% and 1.6%, in the bare and insulated pipes,
respectively. It is therefore abundantly clear that local frictional slip (see Figure 2(c)) not only results in increased bending
compliance, it also produces local damping, which may be equally beneficial in reducing local pipe dynamic amplitudes.
The contribution of local frictional slip to the overall bending stiffness of the local pipe section at the bottom of the HOC was
estimated by re-running the bare pipe case with a very high friction coefficient of 100. The resulting comparison of the
predicted moment-rotation responses is shown in Figure 5.
Figure 5: Effect of Local Frictional Slip

6. 6 OTC-25034-MS
It can be clearly seen that local frictional slip induces both a reduction in bending stiffness and also local damping dissipation.
It is interesting to note that the bending stiffness of the sliding case is initially coincident with that of the non-sliding case at
very small bending moments, and it is only when frictional sliding begins to take place, it progressively becomes less stiff. The
fully bonded non-sliding case naturally produces no local damping.
The finite global bending stiffness observed in Figure 5 for the fully bonded case is the result of combination of pad stiffness
and possibly local pipe flexibility. An insight into the effect of pad stiffness alone was gained by doubling the elastic modulus
of the pads to 40 GPa. The resulting comparison for the fully bonded bare pipe case is shown in Figure 6.
Figure 6: Effect of Pad Stiffness
It is seen that pad stiffness has some influence on the bending stiffness of the pipe-HOC connection even from a base case of
already fairly stiff pads. However, for the specific parameters used in the study, the effect appears to be less significant than
that caused by local frictional sliding. It should be emphasized that whilst the numerical values presented above are entirely
notional, being applicable only to the particular combination of parameters used, the trends will nevertheless be expected to be
true for all real cases encountered in practice.
Cyclic local frictional slip under a constant tensile vertical load may raise concerns regarding longitudinal ratcheting, which is
known not to be an issue in practice. For the particular combination of parameters used in the ABAQUS simulations, the local
axial ratcheting was observed to be minimal, at 10 microns per cycle, even under the rather large cyclic bending excursions
applied in the model. The top of the pipe above the HOC was confirmed to be completely stationary in the analyses.
The effects of the various stiffness values presented above on the predicted fatigue damage response in the pipe in the vicinity
of the HOC was evaluated by transferring the ABAQUS results into Orcaflex and running several simulations, as described
below.
3-D Orcaflex Installation Calculations
The method used to transfer the results from the ABAQUS calculations into Orcaflex simulations was by adding a short length
of a special element in Orcaflex having the calculated hysteretic bending stiffness properties. This special element was initially
developed for representing flexible pipes (Ref.3), but in this case is inserted at the top of the normal length of the pipe catenary
held by the HOC. A schematic illustration of the location of the short element is shown in Figure 7. The length ‘l’ of the
element was chosen to be 50 mm, which is considered to be a representative length over which full fixity conditions are
established. However, the results are not expected to be dependent upon the choice of this length.
Fully fixed models in this analysis incorporated a 50mm stiff element, fully fixed to the vessel, whereas the hysteretic bending
model incorporated the properties of the ABAQUS model interface over the 50mm active element length. Fully fixed in this
context implies a linear (non-hysteretic) perfectly vertical line in the moment-rotation diagrams of Figures 3, 5 and 6. A check

7. OTC-25034-MS 7
was conducted of the hysteretic active element in Orcaflex by forcing a set pipe length to deform, and then releasing it, causing
free damped vibrations. The predicted damping matched the ABAQUS model damping shown in Figure 4.
Figure 7: Location of Additional Special Element in Orcaflex
The parameters used in Orcaflex to investigate the extent to which the properties of the HOC interface affects allowable hang-
off periods are given in Table 2. The same pipe properties as in Table 1 were used. The chosen metocean criteria represent a
fairly harsh wave environment, but the effect of current and wind were not considered, since they are not expected to
significantly affect first order wave fatigue calculations. Fatigue due to vortex induced vibrations was also not considered.
Parameter Value Units
ENVIRONMENTAL
Flat Water Depth 1200 m
Significant Wave Height ‘HS’ 3.0 m
Mean Up-Crossing Period ‘TZ’ (Head Seas) 8.75 sec
Mean Up-Crossing Period ‘TZ’ (Beam Seas) 13.50 sec
Spectral wave model ISSC -
Current 0 m/s
Wind 0 m/s
INSTALLATION VESSEL PARTICULARS
Vessel ILLC Perpendiculars Length ‘LBP’ 197 m
Vessel Maximum Moulded Breadth 49 m
Moulded Depth 22.4 m
Operational Draft 8.4 m
Operational Displacement 65,708 mTe
Centre of Gravity (COG) Coordinates (x,y,z) 96.2, 0.4, 20.5 m
HOC Coordinates (x,y,z) 0.0, -15, 22.4 m
Reel-Lay Tower Angle (from horizontal) 80 deg
Table 2 – Environmental and Vessel Parameters for Installation Analysis

8. 8 OTC-25034-MS
A large installation vessel motion set is used for the analysis; with the HOC located 96.2m aft of the vessel COG. The mean
up-crossing period was chosen to place the spectral peak period of the wave spectrum at the maximum vessel response for
pitch (head seas), and roll (beam seas). A typical standard installation catenary was selected by conducting static analysis
using Orcalay software (Ref.4) to determine a favorable installation configuration; this was replicated in Orcaflex by matching
the catenary characteristics and line data.
Figure 8 – Finite Element Model of Installation Catenary (3m HS Head Seas)
A simulation period of 1hr of a single seed was used to provide a representative fatigue state for the purposes of comparing
relative fatigue between models. The rainflow counting method was employed to sum the fatigue damage. All lines were run
in a single model for head and beam seas, so that results could be easily compared – see Table 3 for load cases and results.
Catenary
Wave Incident
Direction
Coating HOC Model
Allowable Hang-Off
Period(1)
Hang-Off Period
Benefit
1 Head Bare Hysteretic Bending 42.6 hrs
10.6%
2 Head Bare Fully Fixed 38.5 hrs
3 Head Insulated Hysteretic Bending 31.7 hrs
7.9%
4 Head Insulated Fully Fixed 29.4 hrs
5 Beam Bare Hysteretic Bending 11.8 hrs
13.5%
6 Beam Bare Fully Fixed 10.4 hrs
7 Beam Insulated Hysteretic Bending 9.4 hrs
8.9%
8 Beam Insulated Fully Fixed 8.6 hrs
(1) Figure obtained by dividing the allowable damage (0.02) by the damage per hour obtained from analysis. The damage per
hour value was extracted at a distance of 2.15m from the HOC, which may be a typical minimum allowable clamping
distance from any weld. DNV-RP-C203 E-Class weld was assumed.
Table 3 – Load Cases and Results
The allowable fatigue damage budget for installation is normally considered to be 0.02 of total SN damage. This originates
from an overall reduction factor of 5 for safety class normal, 10% of which is the allocated installation fatigue budget. This
value governs the allowable hang-off period for an E-Class weld located the minimum assumed distance of 2.15m from the
HOC. It should be noted that the predicted pipe rotations at or near the HOC were less than the range considered in the
ABAQUS evaluations. The extent of local pipe rotations would be dependent on sea state, top tension and vessel
characteristics. The results in Table 3 do, however, indicate the certain benefit of including local compliance and damping into
the installation simulations. For all cases investigated, the lowest fatigue life occurred at the HOC, a typical distribution of
predicted fatigue life, from HOC to seabed, is shown in Figure 9.

9. OTC-25034-MS 9
Figure 9 - Distribution of Fatigue Life over Arc Length
Discussion
The results presented in this paper are indicative being based on the notional properties used. However, the overall method of
deriving local hysteretic moment-rotation characteristics using ABAQUS and transferring these into Orcaflex was shown to
yield reasonable results. It should be noted that the possible additional contribution from the compliance of the HOC
supporting platework was not included in the evaluation. This again is vessel specific but could be included in more detailed
work. The influence of the local effects on allowable hold period is also highly dependent on all the characteristics used,
especially with respect to the dynamic simulations of the installation conditions. Thus although the assumed sea state my be
harsh, the vessel characteristics used were for a very large and relatively stable vessel, not very representative of a more
typical fully laden reel-lay vessel. So, the numerical results should be taken as being indicative only and clearly applicable to
the particular array of parameters used.
Conclusions
The results show that local conditions at the pipe-HOC interface do have a beneficial influence on predicted fatigue damage
during hold periods of an offshore installation campaign. Of the two effects investigated, local frictional slip between pipe and
pads appears to be more significant than local pad and pipe stiffness. The actual extent of the benefit is highly case specific
and should be evaluated using parameters appropriate to each case. The practical implications of such beneficial effects could
be longer allowable hold periods, lower predicted fatigue damage for a given hold period, or perhaps even installability in
harsher sea states. Current dynamic simulations procedures of installation do not take any benefit of such local effects.
This paper presents a technique for incorporating the local effects of clamp-pipe interaction in installation analyses, using
traditional, commercially available, FE and installation software. An illustrative example is provided in the paper, but due to
the gross assumptions made in certain areas, only the method can be regarded as being the key finding since the numerical
results are clearly dependent upon many of the assumptions. However, the overall approach of deriving local properties using
ABAQUS and transferring these to Orcaflex through the hysteretic bending element facility in Orcaflex appears to give
meaningful predictions of influence on predicted fatigue life. More case specific work could to be carried out based on the
method suggested in this paper to yield more specific predictions. Or, further studies, based on more wide ranging parametric
sensitivities, may be undertaken to perhaps reveal greater benefit from this approach.
HOC
Seabed
Sag-Bend

10. 10 OTC-25034-MS
References
1. ABAQUS V6.12 User’s Manual, SIMULIA 2012.
2. ORCAFLEX 9.7a User’s Manual, ORCINA.
3. Tan, Quiggin, Sheldrake, “Time Domain Simulation of the 3D Bending Hysteresis Behaviour of an Unbonded Flexible Riser”
OMAE2007-29315
4. ORCALAY 3.3b User’s Manual, ORCINA