3.2.1 Sway‐Bending Modes of Vibration

Essentially, this form is observed when the foundation is very rigid compared to the superstructure. Wind turbines supported on monopiles and jackets supported on piles will exhibit such kind of modes. Figure 3.3 shows a schematic diagram of modes of vibration for monopile supported wind turbines and Figure 3.4 shows schematic diagram of a jacket supported wind turbine system. It is important to note that the first two modes are quite widely spaced – the typical ratio is about four to six times. One of the important points to note is that the foundation is axially very stiff. These analyses can be easily carried out using standard software.

It may be worth noting that the first three modes of vibration of a fixed‐based cantilever beam are given by:

3.1

where α_n is a mode number parameter having the value of 1.875, 4.694, 7.855 for the first, second, and third mode, respectively. EI is the bending stiffness of the beam having length L and M is the mass per unit length of the beam.

3.2.1.1 Example Numerical Application of Modes of Vibration of Jacket Systems

Figures 3.5 and 3.6 shows examples of numerical simulation of a jacket supported offshore wind turbine system where the modes of vibration may be appreciated. This is obtained through eigen solutions. For the twisted jacket (also known as inward battered guided structure [IBGS]), the frequency of first mode and second modes is 0.2678 Hz and the third mode is 1.049 Hz. It may be noted that the first two modes are identical and the task of the designer is to place this in the narrow band of allowable soft‐stiff zone, discussed in Section 2.2 of Chapter 2. Similar observations were also noted for standard jackets on piles, see Figure 3.4. This figure also provides a plan view to visualise the mode shapes. Closed‐form solution for obtaining natural frequencies of jacket is developed in Jalbi and Bhattacharya (2018).

3.2.1.2 Estimation of Natural Frequency of Monopile‐Supported Strctures

Following the concept of target natural frequency (Section 2.2 in Chapter 2), in order to place the natural frequency of the whole system in the desired band, iterations are necessary whereby the different design parameters are altered. It is time consuming to carry out the optimization using software, and as a result, closed‐form solutions are effective. Method to calculate natural frequencies are shown in Chapter 5 and worked‐out examples are carried out in Chapter 6.

As an example, the natural frequency of a monopile supported wind turbine system can be estimated following the method developed by Arany et al. (2015a, b, 2016). This simplified methodology builds on the simple cantilever beam formula to estimate the natural frequency of the tower (fixed‐base assumptions), and then applies modifying coefficients to take into account the flexibility of the foundation and the substructure. This is expressed as

3.2

where C_L and C_R are the lateral and rotational foundation flexibility coefficients, f_FB is the fixed‐base (cantilever) natural frequency of the tower considering the varying stiffness along the length of the tower and transition piece (TP). Effectively, the formulation shows the flexibility provided by the foundation. The method is described in Chapter 5 and example applications are shown in Chapter 6. Arany et al. (2016) provides a detailed derivation of the method, along with validation and verification with 10 case studies of the wind turbine structure.

3.2.2 Rocking Modes of Vibration

Rocking modes of foundation are typical of wind turbines supported on multiple shallow foundations; see, for example, Figure 3.7 where wind turbine structures are supported on multiple bucket‐type foundations. Figure 3.8 shows three other types of scaled model tests where a seabed frame or jacket is supported on shallow foundations. A schematic representation of the rocking modes of vibration are shown in Figure 3.9 where the tower mode may get coupled with rigid modes of vibration. Essentially, there are two types of vibrations: (i) rocking mode of the jacket, which is very rigid; (ii) flexible modes of the tower. Theoretically, for each rocking mode, there can be two flexible modes of the tower. These were observed through scaled model tests and reported in Bhattacharya et al. (2013a, b, 2017a). The foundation may rock about different planes and is dictated by the orientation of the principle axes i.e. highest difference of second moment of area. Figure 3.9 is a simplified diagram showing the modes of vibration where the tower modes can also interact with the rocking modes i.e. the tower may or may not follow the rocking mode of the foundation. Rocking modes of a foundation can be complex, as they interact with the flexible modes of the tower and it will be shown that these need to be avoided.

Three cases are discussed below:

1. Wind turbine supported on symmetric tetrapod foundations such as the one shown in Figure 3.7. A simplified model for analysis is illustrated in Figures 3.10 and 3.11. Research by Bhattacharya et al. (2013a,b) shows that even for the same foundations under each support, there will be two closely spaced vibration frequencies. This is due to different vertical stiffness of the foundation associated with variability of the ground. However, after many thousands of cycles of loading and vibration, these closely spaced vibration frequencies may converge to a single peak.
2. Asymmetric tripod foundation. Example is provided in Figure 3.12 inspired by the concept shown in Figure 1.31 (Chapter 1). Study reported in Bhattacharya et al. (2013a,b) showed that there will two modes of vibration with closely spaced frequencies but with millions of cycles of loading, and these two closely spaced peaks will not converge. This is because the foundation has two different stiffness in two orthogonal planes.
3. Symmetric tripod foundation. In a bid to understand the modes of vibration for a symmetric tripod, tests were carried out on a triangular foundation shown in Figures 3.13 and 3.14. Free vibration tests were carried out and a typical result is shown in Figure 3.15. The mode is like a ‘beating phenomenon’ well known in physics, which is possible for two very closed spaced vibration frequencies with low damping.

Taking into consideration Figures 2.4 and 2.5 (Chapter 2) where the design first natural frequency of the whole system is to be targeted between 1P and 3P, it is important not to have two closely spaced modes of vibration. In practical terms, it is therefore recommended to avoid an asymmetric system. The previous case study also shows that a symmetric tetrapod is better than a symmetric tripod due to higher damping. It may be noted that the beating phenomenon is typical of low damping and two closely spaced modes. Gravity‐based foundation will also exhibit rocking modes of vibration and it may also interact with tower flexible modes. Figures 3.16−3.18 show a schematic diagram of observed modes of vibration from a small‐scale model test. Figure 3.19 shows the changing of natural frequency of foundations with cycles of loads observed in small scale testing.

3.2.3 Comparison of Modes of Vibration of Monopile/Mono‐Caisson and Multiple Modes of Vibration

Deep foundations such as monopiles will exhibit sway‐bending mode, i.e. the first two vibration modes are widely spaced – typical ratio is four to five. However, multiple pod foundations supported on shallow foundations (such as tetrapod or tripod on suction caisson) will exhibit rocking modes in two principal planes (which are, of course, orthogonal). Figure 3.20 shows the dynamic response of monopile supported wind turbine and tetrapod foundation plotted in the loading spectrum diagram obtained from scaled models tests.

3.2.4 Why Rocking Must Be Avoided

Foundations (symmetric or asymmetric) on multiple foundations will exhibit two closely spaced natural frequencies corresponding to the rocking modes of vibration in two principal axes. For a soft‐stiff design, these need to be fitted in a narrow gap, as shown in Figure 3.21. Furthermore, as will be discussed in Chapter 5, owing to soil‐structure interaction (SSI), the two spectral peaks change with repeated cycles of loading. Also, for symmetric tetrapods (but not for asymmetric tripods), these two peaks will converge for sandy deposits. From the fatigue design point of view, the two spectral peaks for multipod foundations broaden the range of frequencies that can be excited by the broadband nature of the environmental loading (wind and wave) thereby impacting the extent of motions. Thus, the system lifespan (number of cycles to failure) may effectively increase for symmetric foundations as the two peaks will tend to converge. However, for asymmetric foundations the system life may continue to be affected adversely as the two peaks will not converge. In this sense, designers should prefer symmetric foundations to asymmetric foundations.

3.3.1 Observed Resonance in German North Sea Wind Turbines

Hu et al. (2014) reported resonance for wind turbines. This is discussed further in Chapter 5.

3.3.2 Damping of Structural Vibrations of Offshore Wind Turbines

As the natural frequency of offshore wind turbines is close to forcing frequencies, damping is critical to restrict damage accumulation and avoid premature maintenance. Therefore, discussion is warranted for issues related to damping, and for practical design purposes can be idealised in fore−aft and side‐to‐side vibrations of offshore wind turbines. The main difference between the sway‐bending (or rocking) vibrations about two axes (X and Y) is that in the along‐wind direction, higher damping is expected due to the high aerodynamic damping caused by the rotating blades interacting with the airflow. On the other hand, for side‐to‐side direction, the aerodynamic damping is orders of magnitudes lower.

A nonoperational (parked or idling) OWT has similar aerodynamic damping in the fore−aft as in the side‐to‐side direction. Since the wind load is acting in the along‐wind direction, the highest load amplitudes are expected in fore−aft motion. This is because for most wind turbines in water depths less than 30 m, wind loading is the dominant load, while for very large diameter monopiles in medium to deep water, wave loading is expected to have equal or higher magnitude. It is worth noting that due to the yaw mechanism of the wind turbine, the along‐wind and cross‐wind directions are dynamically moving and are not fixed; therefore, the foundation is subjected to both cross‐wind and along‐wind loading in all directions during the lifetime of the OWT. One can conclude that analysing vibrations in both directions is important.

Studies considering the damping of the first bending mode either empirically or theoretically include Camp et al. (2004); Tarp‐johansen et al. (2009); Versteijlen et al. (2011); Damgaard and Andersen (2012); Damgaard et al. (2013); Shirzadeh et al. (2013). Based on these studies and other estimates and in the absence of other data, the following assessment of damping ratio contributions is recommended:

• Structural damping: 0.15−1.5%. The value of structural damping depends on the connections in the structure (such as welded connections, grouted connections, etc) in addition to material damping (usually steel) through energy dissipation in the form of heat (hysteretic damping).
• Soil damping: 0.444−1%. The sources of damping resulting from soil‐structure interaction (SSI) include hysteretic (material) damping of the soil, wave radiation damping (geometric dissipation) and, to a much lesser extent, pore fluid induced damping. Wave radiation damping and pore fluid induced damping are negligible for excitations below 1 Hz, and therefore hysteretic damping is dominant for the purposes of this study. The soil damping depends on the type of soil and the strain level.
• Hydrodynamic damping: 0.07−0.23%. Results from wave radiation and viscous damping due to hydrodynamic drag. In the low frequency vibration of wind turbines the relative velocity of the substructure is low and therefore viscous damping, which is proportional to the square of the velocity, is typically very low. The larger contribution results from wave radiation damping, which is proportional to the relative velocity.
• Aerodynamic damping: in the fore−aft direction for an operational turbine 1−6%, for a parking turbine or in the crosswind direction 0.06−0.23%. Aerodynamic damping is the result of the relative velocity between the wind turbine structure and the surrounding air. Aerodynamic damping depends on the particular wind turbine, and is inherent in the popular blade element momentum (BEM) theory for aeroelastic analysis of wind turbine rotors. The magnitude for a particular wind turbine also depends on the rotational speed of the turbine.

The total damping of the first mode of vibration is typically between 1−4% in side‐to‐side vibration, or parked or stopped or idling turbine. On the other hand, the total damping is between 2−8% for an operational wind turbine in the fore−aft direction.

3.4.1 Current Limits on the Rotation at Mudline Level

In terms of SLS criteria, some specific requirements are to be met and the most important being the maximum deflection and rotation (tilt) at the foundation level/pile head (mudline) and at nacelle level. When assessing the total rotation, the initial tilt as well as the accumulated permanent rotation resulting from cyclic and dynamic loading throughout the design lifetime of the offshore wind turbine (OWT) should be considered.

The Det Norske Verita (DNV) code (DNV 2010a,b) gives typical limits of 0.25° and 0.5° for allowable rotation and states that ‘the deformation tolerances are typically derived from visual requirements and requirements for the operation of the wind turbine’.

Serviceability criteria appear to be ‘turbine manufacturer requirements’ for the operation of the wind turbine, which originates from onshore wind technology and is very similar to a typical tall structure (such as tall buildings). The very low tilt angle requirement for monopile supported structures appears to be especially overcautious in light of floating OWT technology, where tilt angles up to 7° are acceptable.

3.6.1 Ultimate Limit State (ULS)

The ULS is related to the maximum load‐carrying resistance, and can be reached for several reasons: (i) excessive yielding and/or buckling (i.e. loss of structural resistance); (ii) a failure of a component (e.g. brittle fracture of connections); and (iii) a loss of static equilibrium of the structure (whole or part) with a consequent mechanism (e.g. rigid body behaviour, overturning, and capsizing).

As the main aim of a foundation is to transfer all the loads, during its design life, from the wind turbine structure to the ground safely and within the allowable deformations. The design calculations should ensure that the maximum loads on the foundations are much lower than the capacity of the chosen foundation. This calculation is most dependent on the ultimate strength of the soil, i.e. this is a strength type calculation. The first step in design is to estimate the maximum loads on the foundations (predominantly overturning moment, lateral load and the vertical load) due to all possible design load cases – this is covered in Chapter 2 . Subsequently, the loads need to be compared with the capacity of the chosen foundation. This calculation is necessary to avoid failure of foundation; see, for example, Figure 3.27a,b, which shows two cases of ULS failure for monopiles. In the case of Figure 3.27a, the foundation fails due to soil failure around the foundation and eventually through uprooting of the foundation. On the other hand, Figure 3.27b shows the case where the pile fails by forming a plastic hinge where the overturning moment in the monopile exceeds the plastic moment carrying capacity of the pile. For a jacket on small‐diameter piles, this is also very similar. Figure 3.28 shows some of the ULS cases for anchor piles supporting a floating wind turbine, as discussed in Chapter 2 . On the other hand, Figure 3.29 shows the ULS cases for anchor piles where two types of failure mechanisms (overturning i.e. rigid body rotation and sliding/translation) are depicted.

3.6.2 Serviceability Limit State (SLS)

The SLS corresponds to tolerance criteria associated with the regular and normal use of the wind turbine, including: (i) excessive deflection leading to the second order effects modifying the distribution of loads between supported and supporting structures; (ii) excessive vibration jeopardising the functioning of the nonstructural components; (iii) displacements that exceed the limitation of the equipment; (iv) differential settlements of foundation and soil causing intolerable tilt of the wind turbine; and (v) temperature‐induced excessive deformations.

SLS is also directly linked to target natural frequency (eigen frequency) as discussed in Chapter 2 due to possible amplification due to dynamics. For these calculations, prediction of the natural frequency of the whole system (eigen frequency) is necessary. As natural frequency is concerned with very small amplitude vibrations, linear eigen value analysis will suffice and prediction of initial foundation stiffness is necessary. Therefore, determination of stiffness of the foundation is an important design step.

3.6.3 Fatigue Limit State (FLS)

The FLS is related to cumulative damage due to repeated loads. For the case of monopile, this would require predicting the fatigue life of the monopile, as well as effects of long‐term cyclic loading on the foundation. Again, this step requires stiffness of the foundation.

3.6.4 Accidental Limit States (ALS)

ALS considers potential accidental or unexpected loads (e.g. vessel impact) that can lead to loss of global or local structural integrity.

3.7.1 Scour

Scour, the manifestation of which is a local lowering of seabed around foundations, is essentially sediment transport or erosion of soil and is caused by waves and current. Scour can affect the structural stability of foundations by making the foundations unsupported up to the scour depth and is an important design consideration. Scour is usually quantified in terms of its equilibrium depth S and the scour extension Ls as shown in Figure 3.30.

The scour depth depends on the current speed and there is a threshold current speed at which undisturbed sand far from the pile just starts to move. If the current speed is lower than the threshold, scour does not develop and the seabed remains stable. However, if the current speed increases scour starts to develop. Table 3.2 summarises empirical formulas on scour depth based on literature and design standards. Caution must be exercised while using these expressions for large diameter monopiles as they may not have been calibrated.

Publication	Flow condition	Relationships
Sumer et al. (1992) and adopted in DNV‐OS‐J101 (2014)	Steady Current KC → ∞
Sumer et al. (1992) adopted in DNV‐OS‐J101 (2014)	Waves KC ≥ 6

The equations given in Table 3.2 are governed by Keulegan‐Carpenter (KC) number given by the following equation:

where U_max is the maximum value of the orbital velocity at the bed, T is wave period, and D_P is the pile diameter. DNV code recommends using linear theory to find the orbital velocity at the bed.

KC number is a nondimensional number representing the relative importance of drag forces over inertia forces for objects in an oscillatory fluid flow. Scour protection measured are often taken to avoid unplanned maintenance.

3.7.2 Corrosion

Monopiles located in offshore marine environments are susceptible to corrosion, which has implication on long‐term performance and fatigue life. Current design standards such as the DNV and literature suggest different methods to account for in design:

1. Corrosion allowance, i.e. extra wall thickness, added during design to compensate for any reduction in wall thickness during design life. The corrosion allowance is based on the corrosion rate, which is a minimum of 0.1 mm per year following DNV.
2. Use a coating system such as zinc paint.
3. Cathodic protection is a technique to prevent the corrosion of steel surface by making the surface to act as a cathode of an electrochemical cell.

3.7.3 Marine Growth

The flora and fauna (such as bacteria, plant, and animal life) surrounding an offshore submerged structure may causes marine growth, see Figure 3.33. This will have implications on the structural design as it adds weight, influences the geometry and surface texture, and affects the corrosion rate. Marine growth can be due to tubeworms, barnacles, corals, and sea anemones, which colonise on a structure shortly after installation. The thickness and rate of growth depend on numerous factors such as salinity, oxygen content, pH, current, and temperature, age, orientation, and depth of the structural component below sea level. Structurally, the extra weight needs to be addressed relative to the total mass the foundations and when performing dynamic analysis of the whole system. The increased thickness will also increase the hydrodynamic loads. In the absence of field data, the DNV recommendations provided in Table 3.3 may be used as a guidance. The values are based on Norwegian and UK waters and are provided to give the reader an indication about the size and impact of marine growth on the structural performance of offshore wind turbine foundations.

3.12.1 Installation of Foundations

Installation of large monopiles is currently carried out in the following ways: (i) hydraulic or vibration hammers from above water; (ii) underwater hydraulic hammers; (iii) in different ground conditions, drill‐drive‐drill operation; (iv) drill and grout. In comparison to drilling operations, driving a pile is faster and is relatively less weather sensitive. However, driving may damage the pile head and there may be issues with verticality. As expected, drill‐drive is slower than driving and is used if the driving operation didn't reach the required penetration depth. There are some calculations that engineers must carry out, and these are discussed next (Figure 3.36).

3.12.1.2 Predicting the Increase in Soil Resistance at the Time of Driving (SRD) Due to Delays (Contingency Planning)

During pile installation, often driving has to be stopped and restarted, for various reasons such as changes of cushion or hammer or the addition of a follower. Such delays typically last from a few hours up to a few days. At the design stage, it is necessary to estimate both the SRD during continuous driving and also the amount of set‐up (increase in pile driving resistance) that may occur during delays, to ensure that the hammers taken offshore are sufficient to meet all eventualities. This is a critical design consideration for offshore pile installation projects. Empirical relationships have been proposed for quantifying set‐up – see Table 3.4 for one such case. Set‐up is primarily dependent on three mechanisms or physical processes – consolidation, stress relaxation, and soil ageing. Theoretically, all these mechanisms/processes start acting as soon as the pile penetrates the ground. It is, however, still uncertain which of the above dominate in a particular soil condition, how long each process continues, and the contribution of each component to the observed overall set‐up. There is widespread evidence showing that the capacity of most driven piles increases with time after driving; see Bhattacharya et al. (2009b).

Reference	Equation	Type of soils	Comments (all times are in days)
Skov and Denver (1988)	where Qt and Q0 are the pile capacities at time t and t0 respectively.	Sand and clay	For sand: A = 0.2, t0 = 0.5 For clay: A = 0.6, t0 = 1.0
Huang (1988)	Qt = QEOD + 0.236{1 + loge(t)(Qmax − QEOD)}	Soft ground soil of Shanghai	Qt = Pile capacity at time t Qmax = Maximum pile capacity QEOD = Pile Capacity at the end of driving (EOD).
Guang‐Yu (1988)	Qt = QEOD(0.372St + 1)	Piles driven in soft soils.	Q14 = pile capacity at 14 days and St is the sensitivity of soil.
Svinkin (1996)	Qt = A. QEOD. t0.1	Sand	For upper bound; A = 1.4 For lower bound; A = 1.025

3.12.1.3 Buckling Considerations in Pile Design

Piles can never be assumed to be straight and free from residual stresses due to construction processes. Apart from out‐of‐line straightness (see Figure 3.37), a pile may be thin walled and therefore vulnerable to local buckling (see Figure 3.38). This has been observed in a number of cases where offshore piles have collapsed during driving due to progressive closure of the internal dimensions – the initiating mechanism being local buckling. Thus, the design method has also to consider the interaction between local and global buckling. Figure 3.39 shows a typical offshore installation and Figure 3.40 shows the pile stick‐up. Once the pile is in the sleeve, it is important to check the buckling potential of the pile under the action of the lateral forces due to the wave loading and the hammer weight. Often attachments are used for transportation and lifting, see Figure 3.41, and it is necessary to evaluate the impact of such attachments on pile driving. Details of buckling considerations can be found in Bhattacharya et al. (2005) and Aldridge et al. (2005).

3.12.2 Installation of Suction Caissons

The installation of suction caissons occurs in two distinct stages.

3.12.2.2 Second Stage

In this stage of installation, pumping the water (contained in the cavity that exists between the caisson lid and the soil plug) is carried out. The act of pumping water out of this void is twofold: the first is to increase the downward force, effectively pulling the caisson into the sea bed. The second is to create an upward hydraulic gradient in the soil mass that loosens the soil around the caisson annulus and on the inside of the caisson skirt. This suction needs to be limited to avoid boiling of the sediment on the inside of the caisson. This will still further install the caisson over that achieved through self‐weight installation. In some instances, caissons may be installed without applying suction (stage one only); this is achieved by using a sufficiently large surcharge. A simplified figure showing the installation process is shown in Figure 3.42.

The major reference for the installation load for suction caisson have come from a series of papers produced by Houlsby and Byrne (2005a,b). These estimations are based on theory and corroborated against experimentally obtained data. Both papers provide a conclusive guide as to caisson installation in sand and clay.

3.12.3 Assembly of Blades

The blades can be assembled in two ways: (i) each blade is individually attached see Figure 3.43 (ii) star type: All the blades are assembled and lifted and installed. Figure shows from London Array wind farm see Figure 3.44.

For jackets or seabed frame supported on multiple piles, there are two methods:

1. Subsea templates are first laid in the seabed, and piling is done at the correct location using the template (also known as pre‐piling). The jackets are then transported and the jacket leg‐pile connection grouted.
2. Jackets are transported and then laid in the seabed using mudmats. The piles are then driven.

Figure 3.45 shows installation photograph for Germany's Alpha Ventus offshore Wind farm in deeper waters.

3.12.4 Decommissioning

Decommissioning of energy infrastructure is increasingly becoming an important design consideration from the view of sustainability. This section of the chapter provides a case study of the decommissioning of Lely Wind Farm (Netherlands) where the monopile is also extracted from the ground. All the self‐explanatory photos are provided. In this context, it is necessary to provide an overview of the current legislation on offshore installations.

For the United Kingdom, decommissioning of offshore structures is regulated by UK law and by the OSPAR Decision 98/3 on the Disposal of Disused Offshore Installations (OSPAR Commission 1998). However, the removal of offshore piles is not required under legislation. Typically, the piles are cut below the seabed, at a suitable distance appropriate, often around three metres below the natural seabed.

3.13.1 Monopiles

Monopiles have been predominantly used to support WTGs in water depths up to 30 m. However, there are discussions with regard to the use of monopiles in deeper water depths termed as ‘XL’ monopile. Preliminary calculations suggest that 10 m diameter monopiles weighing 1200 t may be suitable for 45 m water depth, of course dependent on ground conditions. However, the use is uncertain due to the following:

• No codified cyclic design to predict long term tilt;
• Lack of redundancy in foundation system and therefore chance of single‐point failure;
• Installation costs and lack of adequate specialised vessels;
• Connection between foundation, transition piece, and the tower.

Some of these aspects are described below in further details.

1. Lack of redundancy. Monopiles are ‘overturning moment’ resisting structures, and there are two main components: (i) overturning moment arising from the thrust acting at the hub level; (ii) overturning moment due to the wave loading. Also, these two moments can act in two different planes and will vary constantly, depending on the time of the day and time of the year. Monopiles are rigid piles and the foundation collapse can occur if the soil around the pile fails, i.e. there would be rigid body movement. If the foundation starts to tilt, it is very expensive to rectify.
2. Cyclic (rather dynamic) design of monopile. The response of monopile under cyclic/dynamic load is not well understood and there is a lack of guidance in codes of practice. If cyclic design is incorrect, monopiles can tilt in the long term. If the tilt is more than the allowable limit, the turbine may need a shutdown. Monopile design is usually (also wrongly) carried out using API design procedure calibrated for flexible pile design, where the pile is expected to fail by plastic hinge.
3. Issues related to installation of monopiles. Large monopile installation requires suitable vessel availability, as well as specialised heavy lifting equipment. Other issues are noise refusals, buckling of the pile tip, drilling out, and grouted connections. If the site contains weak rock (siltstone/sandstone/mudstone) and where the local geology shows bedrock or hard glacial soils at shallow depths, drive‐drill‐drive techniques may be required, with subsequent increases in cost and schedule. It must be mentioned here that driving reduces the fatigue life.

3.13.2 Jacket on Flexible Piles

There has been a considerable interest in jacket‐type structures for deeper‐water applications but is perceived being expensive due to the amount of steel required. However, jackets supported on piles can be considered as a safe solution due to excellent track record of good performance in offshore O&G industry. Offshore O&G industry has been using long flexible piles (diameters up to 2.4 m), which are easy to drive, and necessary vessels are readily available (relatively as opposed to vessels to install monopiles). This aspect will drive down the TIC (time in construction) costs regarding piling, and also, large vessels are not required for pile installation. However, there are costs associated with jacket installation. One of the requirement is the optimisation of the jacket so as to consume minimum steel. There are two types of jacket – normal jacket or twisted jacket. The advantage of the twisted jacket over normal jacket is fewer number of joints and therefore less fatigue issue.

3.13.3 Jackets on Suction Caissons

Jackets on suction caissons need to be designed so that rocking modes of vibration are avoided.