RISER SYSTEMS 

SCR design methodology is presented in the following flowchart Figure 3.1, “Steel catenary riser design flow chart”.

Figure 3.1 - Steel catenary riser design flow chart

The design checks related to the different design phases of a steel riser system are described in the following sections.

A starting point of the design process is the definition of the riser internal diameter as a function of the fluid characteristics and of the required flow rate. The wall thickness is initially defined to guarantee:

Consideration for corrosion due to internal fluid and to the sea environment shall be performed to identify the actual material characteristics, the required corrosion allowance and the specifications for external coatings (if any).

The assessment of the minimum wall thickness for the riser components can be based on the following codes and standards: API RP 1111, API RP 2RD and ASME B31.4/B31.8, or the DNV codes (Section 1.2).

Static analysis is the fundamental step for preliminary definition of the riser configuration. The design water depth, the maximum static offsets and the heave motions imposed by the floater allow the designer to select the most suitable geometry from the possible alternatives. The definition of the riser total length and the amount of buoyancy (if any) are the next step to achieve a well-balanced initial configuration. Such configuration results into a compromise between the top end tension, the peak combined stresses arising in the touch down point (TDP) or in the bent areas (e.g. arch) and the maximum angle variations at the riser terminations. An adequate margin against the expected dynamic amplification must be accounted for starting from this phase, in order to consider realistic data (e.g. dynamic amplification factor, typically within 1.2-1.3 range, is used to determine the dynamic tensions from the static results).

The mooring system analysis, using numerical or physical models with a statistical representation of the environmental conditions, allows the prediction of the extreme offsets for the floating production system. These values shall be determined for intact or damaged mooring conditions, considering both parallel (current, waves and wind collinear) and transverse weather (current, waves and wind non-collinear) extreme conditions. The transfer functions for the six components of the vessel motion (Response Amplitude Operators or RAOs) in the wave frequency range are usually provided. The top end of the riser is generally subjected to the direct action of waves and current in conjunction with the imposed motion defined by the combination of RAOs with the considered Sea State. The dynamic analysis of the riser is generally based on a time domain approach with non-linear structural and loading models, using both regular and irregular spectral wave conditions.

The main parameters subjected to checks are:

Particular attention has to be paid, under severe storm conditions, to avoid excessive axial compression in the lower part of the riser (close to the TDP), which can lead to uncontrolled deflected shape and involve high bending radius and hence buckling of the pipe.

The need for VIV suppressors should be investigated with respect to VIV effects induced as part of the riser fatigue damage. Should excessive motions be expected, effective suppressors will be selected and applied to the required riser length. Numerical model of the problem is required in order to evaluate the envisaged solutions.

A simplified assessment of VIV is proposed in DNV-OS-F201. This simplified estimate of the induced fatigue damage is computed conservatively by assuming 2D sheared current profiles to apply on the riser (e.g. unidirectional with a magnitude that varies with distance below the sea level).

A lot of developments have been identified in this particular area. Softwares, such as SHEAR7 and VIVA by MIT, OrcaVIV by Orcina are now potentially suitable for handling VIV issues.

The total fatigue damage is assumed being generated by the combined action of the following contributions:

The admissible fatigue life is assumed to be equal to e.g. 10 times the design life for the entire pipe length. Suitable criteria will be defined to couple a particular environmental condition (wave and current) with the corresponding vessel offset. A time-domain approach will be followed to describe the dynamic response of the riser generated by the representative sea-states of the long-term distribution. The total fatigue damage is then evaluated by means of a suitable procedure that shall be aligned with the solution approach, considering a reference S-N (Stress range – Number of cycles to failure) curve like HSE, API X (in RP 2A), DOE or DNV OS C203 curves and the Palmgren-Miner law for summing the partial contributions.

Temporary conditions are generally associated with the life phases of the system before operation. They include:

These phases, together with potentially disconnection and retrieval operations, must be considered within the analysis methodology in order that maximum riser responses in transient static and dynamic conditions are captured. The reference meteocean conditions are identified on the basis of the estimated duration of each operation and of the period of the year selected for the activity.

In conventional flexible pipe, the internal carcass has been designed to provide all of the collapse resistance to the flexible structure. The pipe must be designed to have sufficient collapse resistance with the outer sheath damaged. In this case, the external hydrostatic pressure penetrates the structural layers (pressure and tensile armours) and bears against the fluid barrier.

For deepwater application, a solution to provide additional collapse resistance without adding substantial weight will consist in modifying the pipe structure such that the external hydrostatic will bear outside the pressure armour. Thus both the pressure armour and the carcass will provide internal ring stiffness increasing the collapse resistance.

A further improvement in the collapse resistance for larger diameter and deeper water requires increased ring stiffness of the internal carcass and/or pressure armour.

The axial compressive stiffness of conventional flexible pipe is an order of magnitude lower than the tensile stiffness. When a flexible pipe is subject to axial compression, the tensile armour layers expand radially (i.e. increase of the diameter under compression known as Poisson's effect), with resistance to expansion provided by the helical wrapped tape over the tensile armour and the external sheath itself. The reverse end cap load in deep water is substantial. For example, a 6" ID empty pipe at a water depth of 2000 meters is subject to more than 80 tonnes reversed end cap load.

There are three potential modes of failure of the flexible pipe resulting from the reverse end cap load:

If the flexible pipe is not torsionally balanced under axial compression, or has residual torque due to the manufacturing, reeling or installation processes, then the resulting twist can increase the radial expansion of either the outer or the inner tensile armour layer. This would increase the potential for these failure modes to occur. Thus torsion loading must also be considered in evaluating the resistance of a flexible pipe structure to reverse end cap loads.

In order to reduce the tension loads on flexible pipe for deep-water applications, a system approach can be applied. Different flexible pipe structures are used with a mid-line connection between the flexible pipes. The top section is designed for higher tensile capacity and the lower section is designed for high collapse resistance. This approach saves substantial weight over making the entire pipe designed for both the required collapse resistance and top tension capacity. The high tension capacity pipe can be made with carbon fibre / composite tensile armour to result in further weight reduction.

The estimated weight saving versus steel armoured pipe by using a combination of steel structures composed of high tensile capacity top section and high collapse resistance bottom section is 25%, and 50% for high tensile capacity top section with composite armour and high collapse resistance bottom section with steel armour

Figure 3.2, “Non-bonded flexible pipe structure” and Figure 3.3, “Non bonded flexible pipe structure using composite armour material” present the pipe design and capabilities of the composite armour relative to the steel material currently being used. In addition to weight reduction, the composite armour offers the advantage of being essentially inert to corrosion, hydrogen induced cracking and sulphide stress cracking, all potential mechanisms for reducing the service life of a flexible pipe.

Figure 3.2 - Non-bonded flexible pipe structure

Figure 3.3 - Non bonded flexible pipe structure using composite armour material

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The following table provides mechanical properties comparison between steel and composite materials:

The flexible pipe structure must be designed to withstand the following installation loading conditions:

The pipe structure is checked by stress analysis to verify that it can withstand these combined loading conditions. Where necessary, the pipe layer dimensions or materials are modified to assure that the loading is within the allowable utilisation factors prescribed in API specification 17J.

The following figure sets the flexible pipe water depth limitations (typical), and further details on flexible pipe technologies and applications are provided in document Ref. [34].

Figure 3.4 - Flexible pipe water depth capability (TECHNIP-FMC data - 2006)

Note  Maximum water depth based on the following basis, Flowline laid empty Max. laying tension of 425T DAF = 1.25

The analysis is mainly concerned with:

For hybrid riser tower (Section ), the mechanical sizing of the Flowlines and the central core (or stem) pipe is based on burst, hydrostatic collapse and effective tension analysis. A local buckling analysis will be performed based on the global dynamic results in order to check the initial wall thickness evaluation.

The lines are not designed against the buckle propagation criterion because low bending radii are not expected during installation and production phases. Designing the lines to meet the buckle propagation requirement would be over-conservative and would involve higher steel weight, a bigger subsea buoy and therefore higher drag loads on the riser system.

For a hybrid 'S' riser (Section ), steel catenary risers are designed as in Section Section 3.2.2, “Selection of the riser section”. This kind of system also requires the sizing of the tethers. In case of steel tethers, the API-RP-2SK provides the design criteria for both intact and one damaged tether conditions. No compression must occur within the tethers during the field life which would induce detrimental effect by snatch loads. For the composite rope (e.g. polyester) alternative, the design criteria relative to the allowable minimum tether tension, to the maximum tension limits and fatigue given in API RP 2SM must be verified.

The different steel risers can be independently insulated similarly to the SCR system.

In a riser tower concept, syntactic foam in modular forms are used for the purpose of added buoyancy of the riser through its entire length but could also participate to the riser bundle insulation. The minimum foam thickness is given by manufacturing process, mechanical resistance and insulation considerations. The external diameter of the riser tower and the foam modules is governed by:

The subsea buoy’s net uplift force is dictated by:

The Eigen mode analysis is required to prevent the hydrodynamic coupling between the riser system and the FPS 1^st and 2^nd order motions.

The analysis will consider the 1^st and 2^nd Eigen periods of the hybrid riser system. In general, such natural periods would be expected to be:

The dynamic response can be of two different natures: (1) inertial response governed mainly by the excitation period and (2) resonant response at a natural period of the system, governed by the existing damping rate.

The pendulum resonance mode (i.e. low frequency resonance) can be prevented by a proper hydrodynamic design of (1) subsea buoy tension and (2) flexible jumper configuration, to obtain a pendulum frequency higher than the FPS slow drift frequency but lower than the subsea buoy vortex shedding frequency.

The riser tower bending mode can be excited in the high frequency range, which can varied from high frequency (flexible riser VIV) to the wave frequency range (e.g. FPS 1^st order motion, riser tower VIV).

For the subsea buoy, the vortex shedding frequency falls in the low frequency range.

Optimisation of a hybrid riser system response during the design phase is achieved by adjusting the following parameters: