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Transverse Thrusters
Chapter 14
Transverse Thrusters Chapter Outline 14.1 Transverse Thrusters 14.1.1 Performance Characterization 14.1.2 Unit Design
379 382 383
Many vessels depend for their effectiveness on possessing a good maneuvering capability in confined waters and sometimes in poor weather. Fig. 14.1 illustrates this in the case of ferries maneuvering effectively in confined waters and berthing stern first into link spans: in some cases, to maintain tight operating schedules. In addition to the specific case of the ferry, many other vessels also require an enhanced maneuvering capability. To satisfy these requirements, several methods of providing a directional thrusting capability are available to the naval architect. One of these is the provision of transverse fixed tunnel propulsion units and another is steerable internal duct thrusters, although of course some propriety designs transcend these two boundaries. While the former type is more common, the various options available and their comparative merits must be carefully considered at the vessel’s design stage. The principal types of units are shown schematically in Fig. 14.2.
14.1 TRANSVERSE THRUSTERS Transverse fixed tunnel thrusters essentially comprise an impeller mounted inside a tunnel, which is aligned athwart the vessel and the essential features of the system are illustrated in Fig. 14.3. It is important to emphasize that the system must be considered as an entity; that is impeller, tunnel, position in the hull, drive unit fairings, tunnel opening fairings onto the hull and the protective grid all need to be evaluated as a complete concept if the unit is to satisfy any form of optimization criteria. Incorrect, or at best misleading results, will be derived if the individual components are considered in isolation or, alternatively, some are neglected in the analysis. Although Figs. 14.2 and 14.3 generally show a transverse propulsion unit located in the bow of the vessel, and in this position the unit is termed a bow thruster, such units can and are also located at the stern of the ship. The bow location is, however, the Marine Propellers and Propulsion. https://doi.org/10.1016/B978-0-08-100366-4.00014-6 © 2019 Elsevier Ltd. All rights reserved.
14.2 Steerable Internal Duct Thrusters Bibliography
385 388
more common and for large vessels. Where enhanced maneuverability in less than ideal situations, such as in the case of a ferry, is required they are often fitted in pairs. For some larger ships, such as cruise liners, more are fitted to enhance the ship-handling capability. The decision as to whether to fit one, two, or more units to a vessel is normally governed by the power or thrust requirement and the available draft. The design process for a transverse propulsion unit located in a given vessel has two principal components: first, to establish the thrust, or alternatively the power, required for the unit to provide an effective maneuvering capability and, second, how best to design the unit to give the required thrust in terms of the unit’s geometry. To satisfy the demand, many manufacturers have elected to provide standard ranges of units covering, for example, a power range of 150 to around 5000 hp. and then select the most appropriate unit from the range for the particular application. Other manufacturers, who perhaps tend today to be in the minority, design a particular unit for a given application. To determine the size of a transverse propulsion unit for a given application, two basic philosophical approaches can be adopted. In both cases, the vessel is considered to be stationary with regard to forward ahead speed. The first approach is to perform a fairly rigorous calculation or undertake model tests, perhaps a combination of both, to determine the resistance of the hull in lateral and rotational motion. Such an exercise would also probably be undertaken for a range of anticipated currents. Additionally, the wind resistance of the vessel would also be estimated either by calculation, typically using a method such as that by Gould (1982); by model tests in a wind tunnel or using CFD modeling. The various wind and hydrodynamic forces on the vessel could then be resolved to determine the required thrust at a particular point on the ship to provide
379
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FIG. 14.1 Ro/Ro ships maneuvering in ferry terminal.
(A)
FIG. 14.3 Transverse propulsion unit—general arrangement in hull.
(B) FIG. 14.2 Types of thruster units: (A) transverse propulsion unit and (B) steerable, internal duct thrusters.
the required motion. Methods such as these, while attempting to establish the loading from first principles, can suffer particularly from correlation problems, scale effects, and not least the cost of undertaking the exercise. Therefore, although these methods are adopted sometimes, more particularly with the azimuthing thruster design problem, it is more usual to use the second design approach for the majority of vessels. This alternative approach either uses a pseudoempirical formulation of the ship maneuvering problem coupled with experience of existing vessels of similar type or is based on a ship maneuvering simulation capability. In essence, the pseudoempirical formulation approach attempts to establish a global approximation to the relationship between turning time, required thrust, and wind speed for a particular class of vessel. An approximation to the turning motion of a ship then can be represented by Eq. (14.1) assuming the vessel rotates about a point as seen in Fig. 14.4:
JP
2 d y ¼ MH + MW + MP dt2
(14.1)
where MH and MW are the hydrodynamic and wind moments, respectively, and MP is the moment produced by the thruster about some convenient turning axis. JP is the polar moment of inertia of the ship and d2y/dt2 the angular acceleration. However, by assuming a constant turning rate, the left-hand side of Eq. (14.1) can be put to zero, thereby removing difficulties with the polar inertia term. That is: MH + MW + MP ¼ 0
(14.2)
In pursuing this pseudoempirical approach, it can be argued that the hydrodynamic moment is largely a function of (dy/dt)2 and the wind moment is a function of the maximum wind moment times sin 2y. The thruster moment is simply the thrust times the distance from the point of rotation and, assuming a constant power input to the unit, is a constant k3. Hence, Eq. (14.2) can be rewritten as 2 dy + k2 sin 2y + k3 ¼ 0 (14.3) k1 dt
Transverse Thrusters Chapter 14
381
FIG. 14.4 Transverse propulsion unit nomenclature: (A) surface area definition and (B) force, moment, and velocity definition.
where the coefficients k1 and k2 depend on the water and air densities (rW and rA), the underwater and above water areas (AU and AA), the vessel’s length (L), wind speed (V), etc., as follows: 9 k1 ¼ 0:5rw AU L3 CMW > = and > ; k2 ¼ 0:5rA AA LV 2 CMA| max
(A)
in which CMW and CMA|max are the water and maximum air moment coefficients, respectively. Consequently, Eq. (14.3) can be rewritten as dy ½ðk2 sin 2y + k3 Þ0:5 ¼ k1 dt from which the time to turn through 90 degrees can be estimated as follows: Z p=2 dt t90 ¼ dy (14.4) dy 0
(B)
Several authors have considered this type of relation for transverse propulsion unit sizing. One such approach due to Ka-Me-Wa (1968) uses a form of Eq. (14.4) to derive a set of approximate turning times for three classes of vessel in terms of the turning time for a quarter of a turn as a function of thruster power and with wind speed as a parameter. The relationship used in this case is " #0:5 0:308CMW rW AU L2pp t90 ¼ kT s 0:5CMA rA AA V 2
FIG. 14.5 Average relationship between turning time and power of unit2: (A) Ro/Ro and ferries; (B) cargo ships; and (C) tankers and bulk carriers. (Reproduced with permission from Ka-Me-Wa, 1976. Steering Propellers. Ka-Me-Wa.)
in which k is the distance of the thruster from the point of rotation nondimensionalized by ship length between perpendiculars, Ts is the propulsion unit thrust, and CMA is a mean wind resistive moment coefficient. The vessels explicitly considered in this approach are ferries, cargo liners, and tankers or bulk carriers and Fig. 14.5 reproduces the results of that prediction. Implicit in this type of
approach is, of course, the coefficient of performance of the unit, which relates the unit thrust T to the brake horsepower of the motor. However, coefficients should not introduce large variations between units of similar types; that is, controllable pitch, constant speed units. While curves such as those shown in Fig. 14.5 can only give a
(C)
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TABLE 14.1 Guide to Thrust Per Unit Area Requirements Ship Type
T/AU (kp/m2)
T/AA (kp/m2)
Ro/Ro and ferries
10–14
4–7
Cargo, ships, tugs
6–10
4–8
Tankers/bulk carriers
4–7
14–16
Special craft (i.e., dredgers, pilot vessels, etc.)
10–12
5–8
rough estimate of turning capability, they are useful for estimation purposes. With ships having such widely differing forms, one with another, due account must be taken in the sizing procedure of the relative amounts of the vessel exposed to the wind and to the water. An alternative approach is to consider the thrust per unit area of underwater or above water surface of the vessel. Table 14.1 shows typical ranges of these parameters, compiled from Ka-Me-Wa (1976) and Lips (1976). When interpreting Table 14.1, one should be guided by the larger resulting thrust derived from the coefficients. This is particularly true of the latest generation of Ro/Ro and Ro/ Pax ferries in which considerable wind exposure is an inherent design feature. Furthermore, in the case of tankers and bulk carriers, the assessment of thruster size by the above water area is not a good basis for the calculation. In the case of the alternative approach, that of using a dynamic simulation capability, this lends itself to a much wider class of simulation scenarios. This is because in many of these techniques not only can the transverse propulsion unit’s characteristics be considered but also a whole range of other contributing factors. Typically, these might include the thrusts of the propellers, perhaps in opposite directions (reversing thrusts); the depth of water and proximity of the quay; water currents and wind speeds. Indeed, all these factors are significant in deciding upon the maneuvering equipment necessary for a particular ship. To achieve this level of simulation, the full equations of motion of the ship must be considered and then solved, albeit with some empirical data derived either from the analysis of other full-scale trials, model test data, or the results of computational fluid dynamic studies. Many of these more advanced ship maneuvering simulation capabilities have reached a stage where reasonably reliable predictions can be developed with relatively modest computational facilities. The question of an acceptable turning rate is always a subjective issue and depends on the purpose for which the vessel is intended and the conditions under which it is expected to operate. Consequently, there is no unique answer to this problem. Hawkins et al. (1965) made an extensive study of several types of maneuvering propulsion
FIG. 14.6 Band of rotation rates versus displacement at zero ship speed. (Reproduced with permission from Beveridge, J.L., 1972. Design and performance of bow thrusters. Mar. Technol.)
devices for the US Maritime Administration, and Fig. 14.6 presents curves based on their work showing measured turning rates as a function of displacement. The band shown in the figure represents turning rates, which have been considered satisfactory in past installations.
14.1.1
Performance Characterization
The usual measure of propeller performance defined by the open water efficiency (0) and given by Eq. (6.2) decreases to zero as the advance coefficient J tends to zero. However, at this condition, thrust is still produced and consequently another measure of performance is needed to compare the thrust produced with the power supplied. Several such parameters have been widely used in both marine and aeronautical applications; in the latter case to characterize the performance of helicopter rotors and VTOL aircraft. The most widely used are the static merit coefficient (C) and the Bendemann static thrust factor (z), which are defined by the following relationships: 9 3=2 > 0:00182T 3=2 KT > > pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ¼ 3=2 C¼ > > = SHP rpD2 =4 p KQ (14.5) T K > > h T i> x ¼ 2=3 ¼ > ; Ps D2=3 ðrp=2Þ1=3 KQ2=3 pð2Þ1=3 > In these equations, the following nomenclature applies: T is the total lateral thrust, taken as being equal to the vessel’s reactive force (i.e., the impeller plus the induced force on the vessel), SHP is the shaft horsepower, Ps is the shaft power in consistent units, D is the tunnel diameter, r is the mass density of the fluid, and KT and KQ are the usual thrust and torque coefficient definitions.
Transverse Thrusters Chapter 14
Both of the expressions given in Eq. (14.5) are derived from momentum theory and can be shown pffiffiffi to attain ideal, nonviscous maximum values for C ¼ 2 and z ¼ 1.0 for normal, nonducted propellers. In the case of a ducted propeller with no duct pffiffiffi diffusion, these coefficients become C ¼ 2 and z ¼ 3 2. Clearly it is possible to express the coefficient of merit (C) in terms of the Bendemann factor (z) and from Eq. (14.5), it can easily be shown that: pffiffiffi C ¼ z3=2 2 (14.6)
14.1.2
Unit Design
Having determined the required size of the unit, it is necessary to configure the geometry of the unit to provide the maximum possible thrust. The fundamental decision is to determine whether the unit will be a controllable pitch, constant speed machine or a fixed pitch, variable speed unit: the former type being perhaps the most common amongst larger vessels. In the case of controllable pitch impeller units, the blades are designed as constant pitch angle blades to enable a nominal equality of thrust to be achieved in either direction for a given operational pitch angle. The term nominal equality of thrust is used to signify that equality of thrust is not achieved in practice due to the position of the impeller with respect to the pod and its position in the tunnel. The blades of the controllable pitch units are frequently termed flat-plate blades on account of their shape; although this is not strictly the case since they have an aerofoil cross-sectional shape when viewed normally to their cylindrical sections. In the alternative case of the fixed pitch unit, the radial distribution of pitch angle over the blade can be allowed to vary to develop a suitable hydrodynamic flow regime over the blades: reversal of thrust in this case is achieved by a reversal of rotation of the impeller. Nevertheless, in some instances, flat-plate blades are also used with these types of unit. In both the controllable and fixed pitch cases, the blade sections are symmetrical about their nose-tail lines; that is, the blades do not possess camber. Furthermore, the fixed pitch blade sections need to be bisymmetrical since both edges of the blade have to act as the leading edge for approximately equal times, whereas for the controllable pitch unit a standard National Advisory Committee for Aeronautics (NACA) or other noncambered aerofoil section is appropriate. Transverse propulsion units are a source of noise and vibration largely resulting from the onset of cavitation and the turbulent flow of the water through the tunnel within the vessel. The issue of noise emission is considered in Chapter 10. However, in order to design a unit, which would be able to perform reliably and not cause undue nuisance, this being particularly important in passenger vessels, it is
383
generally considered that the blade tip speed should be kept within the band of 30–34 m/s. Blade design can be achieved by use of either model test data or by theoretical methods. Taniguchi et al. (1966) undertook a series of model tests on a set of six transverse thrusters. These models had an impeller diameter of 200 mm: two had elliptic blade forms while the remaining four were of the Kaplan type and it is this latter type that is of most interest in controllable pitch transverse propulsion unit design. The model tests of Taniguchi et al. considered Kaplan blade designs having expanded area ratios of 0.300, 0.450, and 0.600 in association with a blade number of four. Additionally, there was also a version having a 0.3375 expanded area ratio with three blades. Each of the blades for these units was designed with a NACA 16-section thickness form in association with a nondimensional hub diameter of 0.400 and a capability to vary the pitch ratio between 0 and 1.3. Using these models, Taniguchi et al. evaluated the effects of changes in various design parameters on performance and Fig. 14.7 shows a selection of these results. These highlight the effects of variations in expanded area ratio and pitch ratio; the effects of blade number; and boss ratio. This latter test was carried out with the elliptic blade form unit and the results show that there is little difference between the efficiency () of the elliptic and Kaplan blade forms with the exception that the Kaplan form performs marginally better at all pitch settings. In Fig. 14.7, KT and KQ are the conventional thrust and torque coefficients, respectively, and CF represents the force measured on the simple block hull body containing the tunnel (Fig. 14.8). In these experiments, the efficiency of the unit is defined by ¼
1 KT + CF 3=2 p KQ
(14.7)
The influence of cavitation on these types of blade forms can be seen in results from a different series of flat-plate blades shown in Fig. 14.9. These results, which relate to a blade area ratio of 0.5 and a blade number of four, show how the breakdown of the thrust and torque characteristics occurs with reducing cavitation number for a series of pitch ratios. However, when comparing the results of these tests with those of Taniguchi et al., it should be noted that the test configurations between the results shown in Figs. 14.7 and 14.9 are somewhat different. With regard to theoretical bases for impeller design, several methods exist. These range from empirically based approaches, such as that by van Manen and Superine (1959), to advanced computational procedures of the type discussed in Chapter 8. Indeed, in recent years, the use of boundary element and computational fluid dynamics methods has been particularly helpful in understanding the flow configurations that exist in transverse propulsion units.
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FIG. 14.7 Examples of test data from CP transverse propulsion unit tests: (A) effect of AC/AD; (B) effect of blade number; (C) effect of blade form; and (D) effect of hub diameter. (Reproduced with permission from Taniguchi, K., Watanabe, K., Kasai, H., 1966. Investigations into fundamental characteristics and operating performances of side thrusters. Mitsubishi Tech. Bull.)
The impeller design process is only one aspect of the system design. The position of the impeller in the hull presents an equally important design consideration. Taniguchi et al. (1966) in their extensive model test study examined this problem using simple block models of the hull. In these models the vertical position of the tunnel relative to the base line, the tunnel length, and the effects of frame slope in the way of the tunnel opening could be investigated individually or in combination. Fig. 14.10 shows the effects of these changes at model scale. From these results, it can be seen that these parameters exert an important influence on the overall thrust performance of the unit. As a consequence, it is seen that care needs to be exercised in determining the location of the unit in the hull so as to avoid any unnecessary hydrodynamic losses and also to maximize the turning moment of the unit on the ship system. It will be seen from Fig. 14.10 that these two issues are partially conflicting and, therefore, an element of compromise must be introduced within the design process. Transverse propulsion units are at their most effective when the vessel is stationary in the water with respect to normal ahead speed and tend to lose effectiveness as the
vessel increases its ahead, or alternatively astern, speed. English (1963) demonstrated this effect by means of model tests from which it can be seen that the side thruster loses a significant amount of its effect with ship speeds, or conversely current speeds, of the order of 2–3 knots. The cause of this fall-off in net thrusting performance is due to the interaction between the fluid forming the jet issuing from the thruster tunnel and the flow over the hull surface, due principally to the translational motion of the hull but also, in part, to the rotational motion. Fig. 14.11 shows the effect in diagrammatic form where it is seen that this interaction causes a reduced pressure region to occur downstream of the tunnel on the jet efflux side and this can extend for a considerable way downstream. This induces a suction force on that side of the hull, which reduces the effect of the impeller thrust and alters the effective center of action of the force system acting on the vessel; Chislett and Bjorheden (1966). Considerations of this type led some designers, Schottel (1985), to introduce a venting tube, parallel to the axis of the tunnel, in order to induce a flow from one side of the hull to the other. Fig. 14.12 illustrates the effects of fitting such a device to two different types of vessel
Transverse Thrusters Chapter 14
FIG. 14.8 Taniguchi et al.’s simplified hull form arrangement. (Reproduced with permission from Taniguchi, K., Watanabe, K., Kasai, H., 1966. Investigations into fundamental characteristics and operating performances of side thrusters. Mitsubishi Tech. Bull.)
(Schottel, 1985). More recent research and full-scale practice, however, has shown that the fall off in the effectiveness of bow thrusters with increasing ship speed is attenuated when very large units are employed; typically, of the order of 3 MW and above. Wall effects are important when considering the performance of a transverse propulsion unit. A low-pressure region can be created between the hull surface and the jetty wall when in the presence of a jet from a bow thruster unit. This induces suction between the wall and the hull to occur, which, in the case of an idealized flat-plate at about three jet diameters from the wall, experiences suction of the order of the jet thrust. Such a magnitude, however, decays rapidly with increasing separation distance such that at about six jet diameters the suction is only about 10% of the jet thrust. The tunnel openings need to be faired to some degree to prevent any undue thrust losses from the unit and to minimize the hull resistance penalty resulting from the discontinuity in the hull surface. However, the type of fairing required to enhance the thrust performance of the transverse propulsion unit is not the same as that required to minimize hull resistance during normal ahead operation and, consequently, a measure of compromise is required in the design
385
process within this context. This highlights the compromise necessary in designing the opening fairings to suit the nominally zero speed thrusting condition as well as minimizing the hull aperture resistance at service speeds. Indeed, Holtrop and Mennen (1982) made some regression-based estimates of the aperture size on ship resistance as discussed in Chapter 12. This compromise can normally be achieved provided that the ship’s service speed is below around 20– 24 knots; however, if a service speed at the top end of this range or above is contemplated, then consideration should be given to the fitting of tunnel orifice doors. Fig. 14.13 shows such a case in which doors have been provided to minimize the hull frictional resistance. Furthermore, it can be seen that the hinges on the doors are aligned such that the axis about which door opening occurs approximately aligns with the flow streamlines generated over the bulbous bow, which, in this case, lies to the left of the picture. Such an alignment can be particularly useful in minimizing the ship’s frictional resistance should a door actuating mechanism fail and the door cannot be closed after use. A further advantage of doors fitted to the ends of thruster tunnels is that the turbulent noise generated by the water passing over the tunnel orifice is considerably reduced at service ship speeds when compared to normal thruster openings. Additionally, within the context of noise generation, the traditional shape of controllable pitch transverse propulsion unit blades has been trapezoidal, when viewed in plan form as seen in Fig. 14.9; however, this is not conducive to quietness of operation. The application of moderate skew to the controllable pitch impeller blades of thruster units, Fig. 14.14, has, by helping to control the effects of cavitation, given a further degree of control in minimizing the noise generated by these units. This can be particularly beneficial to passenger ship operation or other ship types where the accommodation is located in the vicinity of the thruster units and where the vessel may be required to maneuver in port while people are still sleeping. Furthermore, careful attention to the hydrodynamic fairing of the pod strut and body and to the changes of section that occur within the tunnel space make a significant difference to the noise generation potential of the unit.
14.2 STEERABLE INTERNAL DUCT THRUSTERS These types of thruster, sometimes erroneously referred to as pump jets, are particularly useful for navigating a ship at slow speed, as well as for the more conventional docking operations. In the case of research ships, for example, when undertaking acoustic trials of one kind or another, it is sometimes helpful not to be dependent upon the main propellers to drive the ship. This is because although the propellers may have been designed to be subcavitating at the speeds
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FIG. 14.9 Effect of cavitation on KT and KQ for a Kaplan blade form.
of interest for scientific measurements, there will still be turbulence noise generated by the flow over the propulsor and its supporting arrangements. Consequently, to have a propulsor driving the ship in the sense of a tractor at the bow of the ship may be helpful since the noise and disturbance of propulsion can be largely removed from the scientific measurement positions or towed arrays. With this type of thruster, the water enters the system through an intake located usually in the vicinity of the ship’s bow and then passes through an impeller-driven pump from where the water is exhausted to the outlet. At the outlet, located on the bottom of the ship, the efflux from the pump
passes through a vectoring ring comprising a cascade of horizontally aligned deflector vanes, which impart a change of direction to the water flow. The cascade can be rotated to any desired direction in the horizontal plane, generally through the full 360 degrees, and resulting from the change in direction of the flow velocities a thrust force can be generated in the desired direction. When free running it is often possible, assuming that the unit has been sized properly and the ship is not too large, to drive the vessel at speeds of the order of 5 knots or so. In the case of berthing, then due to the azimuthing capability of the thruster unit, an additional directional degree of control is afforded.
FIG. 14.10 Effects of tunnel location, frame shape, and entrance radius on model scale: (A) tunnel length series; (B) bottom immersion series; (C) tunnel entrance shape series; and (D) hull frame inclination. (Reproduced with permission from Taniguchi, K., Watanabe, K., Kasai, H., 1966. Investigations into fundamental characteristics and operating performances of side thrusters. Mitsubishi Tech. Bull.)
FIG. 14.11 Transverse propulsion unit jet interactions with forward ship speed.
FIG. 14.12 Effect of AST vent (Schottel, 1985).
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BIBLIOGRAPHY
FIG. 14.13 Doors fitted to a set of transverse propulsion unit openings.
FIG. 14.14 Use of skew with transverse propulsion unit blades.
Beveridge, J.L., 1972. Design and performance of bow thrusters. Mar. Technol. Chislett, M.S., Bjorheden, O., 1966. Influence of Ship Speed on the Effectiveness of a Lateral Thrust Unit. Hydroog Aerodynamisk Laboratorium (Rep. No. Hy-8 Lyngby). English, J.W., 1963. The design and performance of lateral thrust units for ships—hydrodynamic considerations. Trans. RINA 103 (3). Gould, R.W.F., 1982. The Estimation of Wind Loads on Ship Superstructures. Maritime Technology Monograph No. 8, RINA. Hawkins, S., et al., 1965. The use of maneuvering propulsion devices on merchant ships. Robert Taggart Inc. (Report RT-8518, Contract MA-3293). Holtrop, J., Mennen, G.G.J., 1982. An approximate power prediction method. ISP 29. Ka-Me-Wa, 1968. Steering Propellers with Controllable Pitch. Ka-Me-Wa. Ka-Me-Wa, 1976. Steering Propellers. Ka-Me-Wa. Lips, 1976. Lips Transverse Tunnel Thrusters. Lips Publ. Schottel, 1985. Modern Lateral Thrusters with Increased Performance— The Anti-Suction Tunnel—A-S-7. Schottel Nederland B.V. Taniguchi, K., Watanabe, K., Kasai, H., 1966. Investigations into fundamental characteristics and operating performances of side thrusters. Mitsubishi Tech. Bull. van Manen, J.D., Superine, A., 1959. Design of screw propellers in nozzles. ISP 6 (55).
Transverse Thrusters Chapter Outline 14.1 Transverse Thrusters 14.1.1 Performance Characterization 14.1.2 Unit Design
379 382 383
Many vessels depend for their effectiveness on possessing a good maneuvering capability in confined waters and sometimes in poor weather. Fig. 14.1 illustrates this in the case of ferries maneuvering effectively in confined waters and berthing stern first into link spans: in some cases, to maintain tight operating schedules. In addition to the specific case of the ferry, many other vessels also require an enhanced maneuvering capability. To satisfy these requirements, several methods of providing a directional thrusting capability are available to the naval architect. One of these is the provision of transverse fixed tunnel propulsion units and another is steerable internal duct thrusters, although of course some propriety designs transcend these two boundaries. While the former type is more common, the various options available and their comparative merits must be carefully considered at the vessel’s design stage. The principal types of units are shown schematically in Fig. 14.2.
14.1 TRANSVERSE THRUSTERS Transverse fixed tunnel thrusters essentially comprise an impeller mounted inside a tunnel, which is aligned athwart the vessel and the essential features of the system are illustrated in Fig. 14.3. It is important to emphasize that the system must be considered as an entity; that is impeller, tunnel, position in the hull, drive unit fairings, tunnel opening fairings onto the hull and the protective grid all need to be evaluated as a complete concept if the unit is to satisfy any form of optimization criteria. Incorrect, or at best misleading results, will be derived if the individual components are considered in isolation or, alternatively, some are neglected in the analysis. Although Figs. 14.2 and 14.3 generally show a transverse propulsion unit located in the bow of the vessel, and in this position the unit is termed a bow thruster, such units can and are also located at the stern of the ship. The bow location is, however, the Marine Propellers and Propulsion. https://doi.org/10.1016/B978-0-08-100366-4.00014-6 © 2019 Elsevier Ltd. All rights reserved.
14.2 Steerable Internal Duct Thrusters Bibliography
385 388
more common and for large vessels. Where enhanced maneuverability in less than ideal situations, such as in the case of a ferry, is required they are often fitted in pairs. For some larger ships, such as cruise liners, more are fitted to enhance the ship-handling capability. The decision as to whether to fit one, two, or more units to a vessel is normally governed by the power or thrust requirement and the available draft. The design process for a transverse propulsion unit located in a given vessel has two principal components: first, to establish the thrust, or alternatively the power, required for the unit to provide an effective maneuvering capability and, second, how best to design the unit to give the required thrust in terms of the unit’s geometry. To satisfy the demand, many manufacturers have elected to provide standard ranges of units covering, for example, a power range of 150 to around 5000 hp. and then select the most appropriate unit from the range for the particular application. Other manufacturers, who perhaps tend today to be in the minority, design a particular unit for a given application. To determine the size of a transverse propulsion unit for a given application, two basic philosophical approaches can be adopted. In both cases, the vessel is considered to be stationary with regard to forward ahead speed. The first approach is to perform a fairly rigorous calculation or undertake model tests, perhaps a combination of both, to determine the resistance of the hull in lateral and rotational motion. Such an exercise would also probably be undertaken for a range of anticipated currents. Additionally, the wind resistance of the vessel would also be estimated either by calculation, typically using a method such as that by Gould (1982); by model tests in a wind tunnel or using CFD modeling. The various wind and hydrodynamic forces on the vessel could then be resolved to determine the required thrust at a particular point on the ship to provide
379
380
Marine Propellers and Propulsion
FIG. 14.1 Ro/Ro ships maneuvering in ferry terminal.
(A)
FIG. 14.3 Transverse propulsion unit—general arrangement in hull.
(B) FIG. 14.2 Types of thruster units: (A) transverse propulsion unit and (B) steerable, internal duct thrusters.
the required motion. Methods such as these, while attempting to establish the loading from first principles, can suffer particularly from correlation problems, scale effects, and not least the cost of undertaking the exercise. Therefore, although these methods are adopted sometimes, more particularly with the azimuthing thruster design problem, it is more usual to use the second design approach for the majority of vessels. This alternative approach either uses a pseudoempirical formulation of the ship maneuvering problem coupled with experience of existing vessels of similar type or is based on a ship maneuvering simulation capability. In essence, the pseudoempirical formulation approach attempts to establish a global approximation to the relationship between turning time, required thrust, and wind speed for a particular class of vessel. An approximation to the turning motion of a ship then can be represented by Eq. (14.1) assuming the vessel rotates about a point as seen in Fig. 14.4:
JP
2 d y ¼ MH + MW + MP dt2
(14.1)
where MH and MW are the hydrodynamic and wind moments, respectively, and MP is the moment produced by the thruster about some convenient turning axis. JP is the polar moment of inertia of the ship and d2y/dt2 the angular acceleration. However, by assuming a constant turning rate, the left-hand side of Eq. (14.1) can be put to zero, thereby removing difficulties with the polar inertia term. That is: MH + MW + MP ¼ 0
(14.2)
In pursuing this pseudoempirical approach, it can be argued that the hydrodynamic moment is largely a function of (dy/dt)2 and the wind moment is a function of the maximum wind moment times sin 2y. The thruster moment is simply the thrust times the distance from the point of rotation and, assuming a constant power input to the unit, is a constant k3. Hence, Eq. (14.2) can be rewritten as 2 dy + k2 sin 2y + k3 ¼ 0 (14.3) k1 dt
Transverse Thrusters Chapter 14
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FIG. 14.4 Transverse propulsion unit nomenclature: (A) surface area definition and (B) force, moment, and velocity definition.
where the coefficients k1 and k2 depend on the water and air densities (rW and rA), the underwater and above water areas (AU and AA), the vessel’s length (L), wind speed (V), etc., as follows: 9 k1 ¼ 0:5rw AU L3 CMW > = and > ; k2 ¼ 0:5rA AA LV 2 CMA| max
(A)
in which CMW and CMA|max are the water and maximum air moment coefficients, respectively. Consequently, Eq. (14.3) can be rewritten as dy ½ðk2 sin 2y + k3 Þ0:5 ¼ k1 dt from which the time to turn through 90 degrees can be estimated as follows: Z p=2 dt t90 ¼ dy (14.4) dy 0
(B)
Several authors have considered this type of relation for transverse propulsion unit sizing. One such approach due to Ka-Me-Wa (1968) uses a form of Eq. (14.4) to derive a set of approximate turning times for three classes of vessel in terms of the turning time for a quarter of a turn as a function of thruster power and with wind speed as a parameter. The relationship used in this case is " #0:5 0:308CMW rW AU L2pp t90 ¼ kT s 0:5CMA rA AA V 2
FIG. 14.5 Average relationship between turning time and power of unit2: (A) Ro/Ro and ferries; (B) cargo ships; and (C) tankers and bulk carriers. (Reproduced with permission from Ka-Me-Wa, 1976. Steering Propellers. Ka-Me-Wa.)
in which k is the distance of the thruster from the point of rotation nondimensionalized by ship length between perpendiculars, Ts is the propulsion unit thrust, and CMA is a mean wind resistive moment coefficient. The vessels explicitly considered in this approach are ferries, cargo liners, and tankers or bulk carriers and Fig. 14.5 reproduces the results of that prediction. Implicit in this type of
approach is, of course, the coefficient of performance of the unit, which relates the unit thrust T to the brake horsepower of the motor. However, coefficients should not introduce large variations between units of similar types; that is, controllable pitch, constant speed units. While curves such as those shown in Fig. 14.5 can only give a
(C)
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TABLE 14.1 Guide to Thrust Per Unit Area Requirements Ship Type
T/AU (kp/m2)
T/AA (kp/m2)
Ro/Ro and ferries
10–14
4–7
Cargo, ships, tugs
6–10
4–8
Tankers/bulk carriers
4–7
14–16
Special craft (i.e., dredgers, pilot vessels, etc.)
10–12
5–8
rough estimate of turning capability, they are useful for estimation purposes. With ships having such widely differing forms, one with another, due account must be taken in the sizing procedure of the relative amounts of the vessel exposed to the wind and to the water. An alternative approach is to consider the thrust per unit area of underwater or above water surface of the vessel. Table 14.1 shows typical ranges of these parameters, compiled from Ka-Me-Wa (1976) and Lips (1976). When interpreting Table 14.1, one should be guided by the larger resulting thrust derived from the coefficients. This is particularly true of the latest generation of Ro/Ro and Ro/ Pax ferries in which considerable wind exposure is an inherent design feature. Furthermore, in the case of tankers and bulk carriers, the assessment of thruster size by the above water area is not a good basis for the calculation. In the case of the alternative approach, that of using a dynamic simulation capability, this lends itself to a much wider class of simulation scenarios. This is because in many of these techniques not only can the transverse propulsion unit’s characteristics be considered but also a whole range of other contributing factors. Typically, these might include the thrusts of the propellers, perhaps in opposite directions (reversing thrusts); the depth of water and proximity of the quay; water currents and wind speeds. Indeed, all these factors are significant in deciding upon the maneuvering equipment necessary for a particular ship. To achieve this level of simulation, the full equations of motion of the ship must be considered and then solved, albeit with some empirical data derived either from the analysis of other full-scale trials, model test data, or the results of computational fluid dynamic studies. Many of these more advanced ship maneuvering simulation capabilities have reached a stage where reasonably reliable predictions can be developed with relatively modest computational facilities. The question of an acceptable turning rate is always a subjective issue and depends on the purpose for which the vessel is intended and the conditions under which it is expected to operate. Consequently, there is no unique answer to this problem. Hawkins et al. (1965) made an extensive study of several types of maneuvering propulsion
FIG. 14.6 Band of rotation rates versus displacement at zero ship speed. (Reproduced with permission from Beveridge, J.L., 1972. Design and performance of bow thrusters. Mar. Technol.)
devices for the US Maritime Administration, and Fig. 14.6 presents curves based on their work showing measured turning rates as a function of displacement. The band shown in the figure represents turning rates, which have been considered satisfactory in past installations.
14.1.1
Performance Characterization
The usual measure of propeller performance defined by the open water efficiency (0) and given by Eq. (6.2) decreases to zero as the advance coefficient J tends to zero. However, at this condition, thrust is still produced and consequently another measure of performance is needed to compare the thrust produced with the power supplied. Several such parameters have been widely used in both marine and aeronautical applications; in the latter case to characterize the performance of helicopter rotors and VTOL aircraft. The most widely used are the static merit coefficient (C) and the Bendemann static thrust factor (z), which are defined by the following relationships: 9 3=2 > 0:00182T 3=2 KT > > pffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ¼ 3=2 C¼ > > = SHP rpD2 =4 p KQ (14.5) T K > > h T i> x ¼ 2=3 ¼ > ; Ps D2=3 ðrp=2Þ1=3 KQ2=3 pð2Þ1=3 > In these equations, the following nomenclature applies: T is the total lateral thrust, taken as being equal to the vessel’s reactive force (i.e., the impeller plus the induced force on the vessel), SHP is the shaft horsepower, Ps is the shaft power in consistent units, D is the tunnel diameter, r is the mass density of the fluid, and KT and KQ are the usual thrust and torque coefficient definitions.
Transverse Thrusters Chapter 14
Both of the expressions given in Eq. (14.5) are derived from momentum theory and can be shown pffiffiffi to attain ideal, nonviscous maximum values for C ¼ 2 and z ¼ 1.0 for normal, nonducted propellers. In the case of a ducted propeller with no duct pffiffiffi diffusion, these coefficients become C ¼ 2 and z ¼ 3 2. Clearly it is possible to express the coefficient of merit (C) in terms of the Bendemann factor (z) and from Eq. (14.5), it can easily be shown that: pffiffiffi C ¼ z3=2 2 (14.6)
14.1.2
Unit Design
Having determined the required size of the unit, it is necessary to configure the geometry of the unit to provide the maximum possible thrust. The fundamental decision is to determine whether the unit will be a controllable pitch, constant speed machine or a fixed pitch, variable speed unit: the former type being perhaps the most common amongst larger vessels. In the case of controllable pitch impeller units, the blades are designed as constant pitch angle blades to enable a nominal equality of thrust to be achieved in either direction for a given operational pitch angle. The term nominal equality of thrust is used to signify that equality of thrust is not achieved in practice due to the position of the impeller with respect to the pod and its position in the tunnel. The blades of the controllable pitch units are frequently termed flat-plate blades on account of their shape; although this is not strictly the case since they have an aerofoil cross-sectional shape when viewed normally to their cylindrical sections. In the alternative case of the fixed pitch unit, the radial distribution of pitch angle over the blade can be allowed to vary to develop a suitable hydrodynamic flow regime over the blades: reversal of thrust in this case is achieved by a reversal of rotation of the impeller. Nevertheless, in some instances, flat-plate blades are also used with these types of unit. In both the controllable and fixed pitch cases, the blade sections are symmetrical about their nose-tail lines; that is, the blades do not possess camber. Furthermore, the fixed pitch blade sections need to be bisymmetrical since both edges of the blade have to act as the leading edge for approximately equal times, whereas for the controllable pitch unit a standard National Advisory Committee for Aeronautics (NACA) or other noncambered aerofoil section is appropriate. Transverse propulsion units are a source of noise and vibration largely resulting from the onset of cavitation and the turbulent flow of the water through the tunnel within the vessel. The issue of noise emission is considered in Chapter 10. However, in order to design a unit, which would be able to perform reliably and not cause undue nuisance, this being particularly important in passenger vessels, it is
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generally considered that the blade tip speed should be kept within the band of 30–34 m/s. Blade design can be achieved by use of either model test data or by theoretical methods. Taniguchi et al. (1966) undertook a series of model tests on a set of six transverse thrusters. These models had an impeller diameter of 200 mm: two had elliptic blade forms while the remaining four were of the Kaplan type and it is this latter type that is of most interest in controllable pitch transverse propulsion unit design. The model tests of Taniguchi et al. considered Kaplan blade designs having expanded area ratios of 0.300, 0.450, and 0.600 in association with a blade number of four. Additionally, there was also a version having a 0.3375 expanded area ratio with three blades. Each of the blades for these units was designed with a NACA 16-section thickness form in association with a nondimensional hub diameter of 0.400 and a capability to vary the pitch ratio between 0 and 1.3. Using these models, Taniguchi et al. evaluated the effects of changes in various design parameters on performance and Fig. 14.7 shows a selection of these results. These highlight the effects of variations in expanded area ratio and pitch ratio; the effects of blade number; and boss ratio. This latter test was carried out with the elliptic blade form unit and the results show that there is little difference between the efficiency () of the elliptic and Kaplan blade forms with the exception that the Kaplan form performs marginally better at all pitch settings. In Fig. 14.7, KT and KQ are the conventional thrust and torque coefficients, respectively, and CF represents the force measured on the simple block hull body containing the tunnel (Fig. 14.8). In these experiments, the efficiency of the unit is defined by ¼
1 KT + CF 3=2 p KQ
(14.7)
The influence of cavitation on these types of blade forms can be seen in results from a different series of flat-plate blades shown in Fig. 14.9. These results, which relate to a blade area ratio of 0.5 and a blade number of four, show how the breakdown of the thrust and torque characteristics occurs with reducing cavitation number for a series of pitch ratios. However, when comparing the results of these tests with those of Taniguchi et al., it should be noted that the test configurations between the results shown in Figs. 14.7 and 14.9 are somewhat different. With regard to theoretical bases for impeller design, several methods exist. These range from empirically based approaches, such as that by van Manen and Superine (1959), to advanced computational procedures of the type discussed in Chapter 8. Indeed, in recent years, the use of boundary element and computational fluid dynamics methods has been particularly helpful in understanding the flow configurations that exist in transverse propulsion units.
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FIG. 14.7 Examples of test data from CP transverse propulsion unit tests: (A) effect of AC/AD; (B) effect of blade number; (C) effect of blade form; and (D) effect of hub diameter. (Reproduced with permission from Taniguchi, K., Watanabe, K., Kasai, H., 1966. Investigations into fundamental characteristics and operating performances of side thrusters. Mitsubishi Tech. Bull.)
The impeller design process is only one aspect of the system design. The position of the impeller in the hull presents an equally important design consideration. Taniguchi et al. (1966) in their extensive model test study examined this problem using simple block models of the hull. In these models the vertical position of the tunnel relative to the base line, the tunnel length, and the effects of frame slope in the way of the tunnel opening could be investigated individually or in combination. Fig. 14.10 shows the effects of these changes at model scale. From these results, it can be seen that these parameters exert an important influence on the overall thrust performance of the unit. As a consequence, it is seen that care needs to be exercised in determining the location of the unit in the hull so as to avoid any unnecessary hydrodynamic losses and also to maximize the turning moment of the unit on the ship system. It will be seen from Fig. 14.10 that these two issues are partially conflicting and, therefore, an element of compromise must be introduced within the design process. Transverse propulsion units are at their most effective when the vessel is stationary in the water with respect to normal ahead speed and tend to lose effectiveness as the
vessel increases its ahead, or alternatively astern, speed. English (1963) demonstrated this effect by means of model tests from which it can be seen that the side thruster loses a significant amount of its effect with ship speeds, or conversely current speeds, of the order of 2–3 knots. The cause of this fall-off in net thrusting performance is due to the interaction between the fluid forming the jet issuing from the thruster tunnel and the flow over the hull surface, due principally to the translational motion of the hull but also, in part, to the rotational motion. Fig. 14.11 shows the effect in diagrammatic form where it is seen that this interaction causes a reduced pressure region to occur downstream of the tunnel on the jet efflux side and this can extend for a considerable way downstream. This induces a suction force on that side of the hull, which reduces the effect of the impeller thrust and alters the effective center of action of the force system acting on the vessel; Chislett and Bjorheden (1966). Considerations of this type led some designers, Schottel (1985), to introduce a venting tube, parallel to the axis of the tunnel, in order to induce a flow from one side of the hull to the other. Fig. 14.12 illustrates the effects of fitting such a device to two different types of vessel
Transverse Thrusters Chapter 14
FIG. 14.8 Taniguchi et al.’s simplified hull form arrangement. (Reproduced with permission from Taniguchi, K., Watanabe, K., Kasai, H., 1966. Investigations into fundamental characteristics and operating performances of side thrusters. Mitsubishi Tech. Bull.)
(Schottel, 1985). More recent research and full-scale practice, however, has shown that the fall off in the effectiveness of bow thrusters with increasing ship speed is attenuated when very large units are employed; typically, of the order of 3 MW and above. Wall effects are important when considering the performance of a transverse propulsion unit. A low-pressure region can be created between the hull surface and the jetty wall when in the presence of a jet from a bow thruster unit. This induces suction between the wall and the hull to occur, which, in the case of an idealized flat-plate at about three jet diameters from the wall, experiences suction of the order of the jet thrust. Such a magnitude, however, decays rapidly with increasing separation distance such that at about six jet diameters the suction is only about 10% of the jet thrust. The tunnel openings need to be faired to some degree to prevent any undue thrust losses from the unit and to minimize the hull resistance penalty resulting from the discontinuity in the hull surface. However, the type of fairing required to enhance the thrust performance of the transverse propulsion unit is not the same as that required to minimize hull resistance during normal ahead operation and, consequently, a measure of compromise is required in the design
385
process within this context. This highlights the compromise necessary in designing the opening fairings to suit the nominally zero speed thrusting condition as well as minimizing the hull aperture resistance at service speeds. Indeed, Holtrop and Mennen (1982) made some regression-based estimates of the aperture size on ship resistance as discussed in Chapter 12. This compromise can normally be achieved provided that the ship’s service speed is below around 20– 24 knots; however, if a service speed at the top end of this range or above is contemplated, then consideration should be given to the fitting of tunnel orifice doors. Fig. 14.13 shows such a case in which doors have been provided to minimize the hull frictional resistance. Furthermore, it can be seen that the hinges on the doors are aligned such that the axis about which door opening occurs approximately aligns with the flow streamlines generated over the bulbous bow, which, in this case, lies to the left of the picture. Such an alignment can be particularly useful in minimizing the ship’s frictional resistance should a door actuating mechanism fail and the door cannot be closed after use. A further advantage of doors fitted to the ends of thruster tunnels is that the turbulent noise generated by the water passing over the tunnel orifice is considerably reduced at service ship speeds when compared to normal thruster openings. Additionally, within the context of noise generation, the traditional shape of controllable pitch transverse propulsion unit blades has been trapezoidal, when viewed in plan form as seen in Fig. 14.9; however, this is not conducive to quietness of operation. The application of moderate skew to the controllable pitch impeller blades of thruster units, Fig. 14.14, has, by helping to control the effects of cavitation, given a further degree of control in minimizing the noise generated by these units. This can be particularly beneficial to passenger ship operation or other ship types where the accommodation is located in the vicinity of the thruster units and where the vessel may be required to maneuver in port while people are still sleeping. Furthermore, careful attention to the hydrodynamic fairing of the pod strut and body and to the changes of section that occur within the tunnel space make a significant difference to the noise generation potential of the unit.
14.2 STEERABLE INTERNAL DUCT THRUSTERS These types of thruster, sometimes erroneously referred to as pump jets, are particularly useful for navigating a ship at slow speed, as well as for the more conventional docking operations. In the case of research ships, for example, when undertaking acoustic trials of one kind or another, it is sometimes helpful not to be dependent upon the main propellers to drive the ship. This is because although the propellers may have been designed to be subcavitating at the speeds
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Marine Propellers and Propulsion
FIG. 14.9 Effect of cavitation on KT and KQ for a Kaplan blade form.
of interest for scientific measurements, there will still be turbulence noise generated by the flow over the propulsor and its supporting arrangements. Consequently, to have a propulsor driving the ship in the sense of a tractor at the bow of the ship may be helpful since the noise and disturbance of propulsion can be largely removed from the scientific measurement positions or towed arrays. With this type of thruster, the water enters the system through an intake located usually in the vicinity of the ship’s bow and then passes through an impeller-driven pump from where the water is exhausted to the outlet. At the outlet, located on the bottom of the ship, the efflux from the pump
passes through a vectoring ring comprising a cascade of horizontally aligned deflector vanes, which impart a change of direction to the water flow. The cascade can be rotated to any desired direction in the horizontal plane, generally through the full 360 degrees, and resulting from the change in direction of the flow velocities a thrust force can be generated in the desired direction. When free running it is often possible, assuming that the unit has been sized properly and the ship is not too large, to drive the vessel at speeds of the order of 5 knots or so. In the case of berthing, then due to the azimuthing capability of the thruster unit, an additional directional degree of control is afforded.
FIG. 14.10 Effects of tunnel location, frame shape, and entrance radius on model scale: (A) tunnel length series; (B) bottom immersion series; (C) tunnel entrance shape series; and (D) hull frame inclination. (Reproduced with permission from Taniguchi, K., Watanabe, K., Kasai, H., 1966. Investigations into fundamental characteristics and operating performances of side thrusters. Mitsubishi Tech. Bull.)
FIG. 14.11 Transverse propulsion unit jet interactions with forward ship speed.
FIG. 14.12 Effect of AST vent (Schottel, 1985).
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BIBLIOGRAPHY
FIG. 14.13 Doors fitted to a set of transverse propulsion unit openings.
FIG. 14.14 Use of skew with transverse propulsion unit blades.
Beveridge, J.L., 1972. Design and performance of bow thrusters. Mar. Technol. Chislett, M.S., Bjorheden, O., 1966. Influence of Ship Speed on the Effectiveness of a Lateral Thrust Unit. Hydroog Aerodynamisk Laboratorium (Rep. No. Hy-8 Lyngby). English, J.W., 1963. The design and performance of lateral thrust units for ships—hydrodynamic considerations. Trans. RINA 103 (3). Gould, R.W.F., 1982. The Estimation of Wind Loads on Ship Superstructures. Maritime Technology Monograph No. 8, RINA. Hawkins, S., et al., 1965. The use of maneuvering propulsion devices on merchant ships. Robert Taggart Inc. (Report RT-8518, Contract MA-3293). Holtrop, J., Mennen, G.G.J., 1982. An approximate power prediction method. ISP 29. Ka-Me-Wa, 1968. Steering Propellers with Controllable Pitch. Ka-Me-Wa. Ka-Me-Wa, 1976. Steering Propellers. Ka-Me-Wa. Lips, 1976. Lips Transverse Tunnel Thrusters. Lips Publ. Schottel, 1985. Modern Lateral Thrusters with Increased Performance— The Anti-Suction Tunnel—A-S-7. Schottel Nederland B.V. Taniguchi, K., Watanabe, K., Kasai, H., 1966. Investigations into fundamental characteristics and operating performances of side thrusters. Mitsubishi Tech. Bull. van Manen, J.D., Superine, A., 1959. Design of screw propellers in nozzles. ISP 6 (55).