Unconventional Propulsion

1

The Specialist Committee on Unconventional Propulsors Final Report and Recommendations to the 22nd ITTC

u 2 0 2 2 Memorial University of Newfou Canada, September 1997. st The 21 ITTC appointed the Specialist Com\ u 2 0 2 2 Universit\u00e0 degli Studi di mittee on Unconventional Propulsors with the February 1998. following membership: \ u 2 0 2 2 Bassin d\u2019Essais des Car\u0 tember 1998. \ u 2 0 2 2 Dr. Neil Bose, (Canada), Chair; Memorial University of Newfoundland, St. John\u2019s. \ u 2 0 2 2 Dr. Michael Billet (USA),2Secretary; TASK SETApFROM THE 21ST ITTC plied Research Laboratory - Penn State, State College, PA. Develop guidelines for carrying out propulsion tests and extrapolating the results to full \ u 2 0 2 2 Dr. Poul Andersen (Denmark); Technical University of Denmark, Lyngby. scale for propellers with ducts, partial ducts, and post-swirl devices, tip plates and z\ u 2 0 2 2 Dr. Mehmet Atlar (UK);preUniversity of Newcastle upon Tyne, Newcastle upondrives. Tyne. \ u 2 0 2 2 Mr. Christian Dugu\u00e9 (France); Bassin 3 INTRODUCTION d\u2019Essais des Car\u00e8nes, Val de Reuil. \ u 2 0 2 2 Dr.Ing. Marco Ferrando (Italy); Universit\u00e0 In order to address extrapolation methodoldegli Studi di Genova, Genova. ogy for unconventional propulsors, the Com\ u 2 0 2 2 Dr. Wenhao Qian (China); Marine Design and Research Institute of China, Shanghai.mittee first reviewed known extrapolation methods currently in use for ships fitted with \ u 2 0 2 2 Dr. Yide Shen (Belgium); University of different types of these devices. A summary of Li\u00e8ge, Li\u00e8ge. these reviews is given in section 4 of this report. At the first meeting of the Committee, Dr. Section 5 then describes the main extrapolation methods in further detail. Michael Billet was elected Secretary of the 1 MEMBERSHIP AND MEETINGS

Committee. Meetings were held as follows:

\

u 2 0 1997.

2

2

\

Many of the extrapolation methods presently in use for powering performance prediction University of Li\u00e8ge, Belgium, February of ships fitted with unconventional propulsors are based on modifications of the ITTC 1978

2

method. The ITTC 1978 formulation was de- for powering prediction of ships fitted with unveloped primarily for single-screw openconventional propulsors. Section 7 gives the propeller ships and established an extrapolation Committee\u2019s Conclusions and Recommendamethodology using interaction parameters such tions. as wake and thrust deduction fraction as well as Reynolds number based friction corrections. Organisations have not only modified the ITTC 4. MODEL SCALE AND FULL SCALE 1978 approach for open propellers based on RESULTS OF VARIOUS EXTRAPOLAexperience, but more importantly in the present TION METHODS FOR UNCONVENcontext, developed new methodology for each TIONAL PROPULSORS particular type of unconventional propulsor treated by using these methods. However, on 4.1 Propeller Boss Cap Fins (PBCF) the horizon are developing unsteady RANS codes, new testing procedures for towing tanks, Recent activities concerning effective boss application of large high-Reynolds number cap designs to improve the efficiency and water tunnels, etc., that can more rigorously cavitation performance of marine screw proaddress these issues. Also, there is still a great pellers have focused on the device known as need for reliable full scale unconventional pro- \u201cPropeller Boss Cap Fins\u201d (PBCF). This wa pulsor performance data to validate any meth- proposed by a group of Japanese inventors, odology. Ogura et al. (1988), as a means of increasing

the efficiency of a ship screw propeller. Almost Devices added to the blade tips, such as all the research and development work on bands, tip fins, bulbs, etc., and those added to PBCF have been reported through a series of the propeller hub, such as boss cap fins, form a technical articles published by the technical class of unconventional propulsor that can departments of the Mitsui OSK Lines Ltd., utilise relatively reliably the ITTC 1978 exWest Japan Fluid Engineering Laboratory Co. trapolation methodology. However, the effect Ltd and the Mikado Propeller Co Ltd. who are of Reynolds number on this methodology the joint patent holders of the device. needs to be resolved as devices that locally modify the flow field to control cavitation or Ouchi (1988) and Ouchi et al. (1988 and improve efficiency are often tested at relatively1989) presented early research and developlow Reynolds number in the towing tank. The ment work on PBCF involving detailed model more extreme cases where the propulsor can betests, flow visualisations and the first full-scale considered part of the ship hull, such as inte- measurements on the 44,979 GT PCC \u201cMergrated ducted propulsors, form a group of un- cury Ace\u201d. These investigations reported a gain conventional propulsors where the ITTC 1978 of 3-7% for the propeller efficiency through the methodology is clearly not adequate. In some \u201creverse PO T \u201d (Propeller Open Test) in a ci cases, unconventional propulsors involving culating water channel, 2 to 2.3 % gain in the pre- and post-swirl vanes, ducts, propeller pods, propulsive efficiency based on self-propulsion etc., have been treated for extrapolation pur- tests and finally a gain of 4% for the power poses by using a modified ITTC 1978 method- output on the actual vessel. Further investigaology. In other cases new methods have been tions were also reported by Ouchi (1989), developed for extrapolation and it is expected Ouchi & Tamashima (1989) and Ouchi et al that the most reliable methods to be developed (1989,1990 and 1992) including 3-D LDV in the future will make use of new methods. measurements in the model propeller slipstream, full-scale measurements with 11 vesSection 6 discusses the rationale behind thesels and cavitation and noise investigations. Committee\u2019s proposals for the future developThe comparative analysis of the sea trial results ment of more general extrapolation methods of 11 ships and their models indicated consid-

3 erable scale effect between the model and actucrease. It was anticipated that the total gain al measurements such that the efficiency gain would be somewhat reduced due to the fricat full scale could be two or three times that at tional drag of the fins. Indeed thrust and torque the model scale. This was mainly attributed to measurements with the same model showed laminar flow around the boss cap and possible 5% gain in efficiency with the PBCF confirmlaminar flow separation at the back of the fins. ing the earlier predictions independently. It was Finally a suitable scale factor, which could be also claimed that it was not apparent from the determined from a large number of tests on respective studies that the PBCF type device models and actual ships in the same manner as had to contain the same number of vanes as the the roughness allowance in the resistance of number of the propeller blades to achieve the ships, was recommended to compensate for themax efficiency gain as claimed in the PBCF scale effect. patent. Ouchi (1988) and Ouchi et al. (1990) also As far as extrapolation methods are conreported on various other aspects of PBCF as- cerned, no particular procedure dedicated to sociated with the extrapolation of the poweringPBCF has been reported in the above referperformance. They indicated that the presenceences. However, it appears that the patent of a rudder significantly reduces the strength ofholding companies have a well-established the hub vortex and hence the gain in propeller procedure to quantify the power saving and efficiency due to PBCF can be reduced by 10- hence the efficiency by the use of PBCF. This 30%. Self-propulsion tests to investigate the procedure is based upon the reverse POT setpropulsive efficiency by PBCF indicated that up in their circulating water channel and the the efficiency gain obtained from the selfexperience gained over a considerable number propulsion tests were similar to the gain obof full-scale measurements. Moreover, private tained from the \u201creverse POT\u201d testscommunication with the with the West Japan Fluid Enrudder. The effectiveness of the PBCF was re- gineering Laboratory Co. Ltd. by this commitported to be hardly affected by the hull wake tee has confirmed that the standard extrapolasuggesting that most of the efficiency gain was tion technique for the performance prediction is due to the propeller flow itself, but it was im- based upon the ITTC 1978 procedure, as brieportant to take into account the effect of the fly reviewed in Section 5.6. rudder. The thrust identity analysis of the selfpropulsion tests, in which the PBCF was conAtlar and Patience (1998) gave a detailed sidered as an appendage to the propeller, dereview of activities on various boss cap designs monstrated that the effect of the PBCF prias well as on PBCF. Amongst them is a recent marily appears as an increase in the relative\u2013 innovative concept, which is known as the Hub rotative efficiency due to the reduction in the Vortex Vane (HVV), jointly developed by the propeller torque due to the PBCF. Potsdam Model Basin, SVA and SCHOTTEL. The HVV is a small vane propeller fixed to the In order to shed more light on the function tip of a cone shaped boss cap within a limiting of PBCF, Gearhart & McBride (1989) carried radial boundary where the tangential velocities out a detailed flow analysis and power measdue to the hub vortex are greater than those of urements with PBCF. The experimental analythe propeller. In this range the small vane prosis of the efficiency gain from PBCF utilised peller diverts the high tangential velocities in the results of LDV measurements in the ARL the direction of the jet, thereby generating adPennstate water tunnel with a model propeller ditional thrust. It is claimed that this mechabehind a hull including its rudder. This connism is different from the mechanism for cluded that 2 % out of a total 6% efficiency PBCF where the fins are located usually gain was due to thrust increase while the rebeyond this limiting radius and the fins themmaining 4% was associated with torque de-

4 selves do not generate thrust. Unlike PBCF, theThe special geometries are in general only limnumber of the vanes of HVV is greater than ited modifications of the propeller. For this reathat of the propeller blade. In their product son the model testing and scaling of results to briefs, Potsdam Model Basin(SVA) (1995) and full scale can in principle be done in the same SCHOTTEL (1995), and in a detailed technical way as for a conventional propeller. In this report by Schulze (1995), the results of LDV section a review of model testing methods and measurements and cavitation details with HVVscaling of the results from such propellers is models are presented. These publications report given. As far as end-plate and similar propelremarkable reductions in the hub vortex cavi- lers are concerned the literature often deals tation by the HVV as well as a successful apwith the propulsors separately, i.e. if and how plication of HVV on the full scale claiming an they improve their efficiency, whereas scaling increase of 3% in the propeller efficiency. and power prediction are not treated. Here, only the latest references or those related to scaling and prediction are given. Atlar and Patience (1998) also reported on a series of experimental investigations with Sparenberg & de Vries (1987) by using various boss caps, which were done in the Emlinear optimisation theory designed and tested a erson Cavitation Tunnel at Newcastle Univer3-bladed end-plate propeller. Open water tests sity, over the last decade. These involved comwere carried out in two different series for two parative efficiency and cavitation performance different rates of revolutions. Correspondingly, measurements with different boss caps through the Reynolds numbers (with respect to chord the reverse open water tests in the presence of length and relative inflow velocity at the 0.85R no rudder and later on with a rudder for some section) varied. There were clearly differences caps, Atlar et al (1998). The cylindrical boss in the measured open-water curves (KQ and η, caps tested were fitted with: fins similar to PBCF; fins with end plates; fins as extension to but only small differences in KT) reflecting the the propeller blades as Trailing Edge Flaps influence of Reynolds number. Over a range of advance ratios, including the design value, the (TEF); an accelerating duct; a decelerating duct and Boss Slots (or holes) (BoS). The latter efficiency was higher for the high Reynolds number, as expected. The tests also included are designed to increase the pressure in the hub measurements with the end-plate propeller vortex core. The comparison of the performances for these caps displayed comparable working behind the model of a tanker, making efficiency savings relative to a standard cone it possible to present the propeller efficiency in the behind condition. shape type except for the cap fitted with the decelerating duct which displayed a large Model testing of an optimally designed amount of efficiency loss. The BoS type cap propeller with two-sided shifted end plates on displayed a superior suppression of the hub vortex cavitation, but also at the cost of a con- the blades was described by de Jong et al. siderable efficiency loss. No particular mention(1992). The tests consisted of open water tests has been made of extrapolation issues in these and cavitation tests. The former were carried out at a rate of rotation sufficiently high to investigations. avoid laminar-flow effects on the propeller blades which was checked by further experiments. In the cavitation tests uniform inflow 4.2 Tip Fins was considered. The interest was focused on the end plate propeller and on comparisons End-plate and other such propellers that with the same propeller without end plates and have modified blade tips, have various names with a reference propeller. For that reason scale like tip-fin, Kappel, winglet, etc., propellers. effects were not considered in greater detail than as outlined above, nor were interaction

5 between propeller and ship or aspects of power procedure to be applied to test results with CLT prediction dealt with. propellers request certain unexpected corrections only known by (the designer) and hence For the special end-plate propeller known not available for model basins. This makes it as the CLT propeller, Hollstein et al. (1997) advisable to rely for the predictions of fulloutline the design of such a propeller for a bulk scale performance of the CLT propeller on the carrier and compare this with data for a sister basis of direct calculations”. In their paper they ship fitted with a conventional propeller. Ac- show calculated open-water characteristics of cording to them no model tests were carried out, their CLT propeller as well as the characterisalthough resistance, open-water and selftics of the conventional propeller. They also propulsion test results for a conventional pro- give trial trip results for the ships with convenpeller are available for the designer of the CLT tional and CLT propellers. propeller. This is because “...the interpolation Table 1 Propulsion coefficients, power and rate of revolutions obtained by model experiments and three different extrapolations and by theoretical predictions (Andersen 1996, table 3).

V = 9.03 m/s w

DMI extrapolation Tip-fin Conv.

ITTC-78

DMI-ITTC

Extrapolation Tip-fin

Conv.

Design*

extrapolation Tip-fin

Conv.

Tip-fin

0.30

0.27

0.27

0.26

0.30

0.27

t

0.16

0.16

0.16

0.16

0.16

0.16

ηR

1.00

1.02

1.00

1.02

1.00

1.02

ηH

1.21

1.16

1.16

1.14

1.21

1.16

-

-

η0

0.63

0.62

0.68

0.66

0.67

0.65

-

-

ηD

0.75

0.73

0.79

0.76

0.81

0.77

-

-

-

-

η D tip fin η D conventional

1.037

1

1.038

1

1.047

1

0.26

Conv.

1.00

0.26 1.00

PD [kW]

6720

6980

6410

6650

6270

6570

6850*

7260*

n [1/s]

1.95

1.99

1.98

2.01

1.95

1.99

1.97

2.01

CTh

1.58

1.49

1.49

1.55*

1.46

1.42

1.58

1.55*

* Design includes an increase in resistance (allowance) of 9 per cent For the tip-fin or Kappel propeller with in- basis of model test results. One method was tegrated fins in the tip region (bent blade tips) due to the model basin (Danish Maritime InstiAndersen (1996) carried out a comparative tute) where a correlation allowance and a corstudy with the conventional propeller actually rection to the wake fraction coefficient was apfitted on the ship as a reference. The model plied. The second method followed the ITTC tests consisted of traditional tests, i.e. in addi- 1978 with the exception that a more detailed tion to resistance tests for the ship model, scaling of the influence of friction on the openopen-water, self-propulsion and cavitation testswater propeller characteristics was carried out. for both conventional and tip-fin propeller This procedure was based on the flat plate fricmodels. Three different methods were used fortional resistance coefficient. The third method making the full-scale power prediction on the combined the two others in applying the proce-

6 dure of the model basin (correlation allowance speed, vanished when the air cavity approached and correction to the wake) and the correction its ultimate form, i.e. for Froude numbers to the propeller open-water characteristics. The greater than 3. Hadler and Hecker (1968) indithree methods gave different required power, rectly concurred with this hypothesis, but they demonstrating the need for rational procedurescalculated the Froude number using the total for such estimations (see table 1.). inflow velocity and the immersion of the shaft as the length parameter. Finally, Olofsson (1996) acknowledged an influence of FnD, (calculated with the diameter as the length pa4.3 Surface Piercing Propellers rameter), but he set to 4 the limiting value beyond which the influence disappears. Summarizing, it is widely recognized that the Surface piercing propellers (SPP) can be Froude number does affect the behavior of included into the unconventional propulsor faSPPs, and all authors have suggested the exismily because of their particular mode of operatence of a threshold value that limits this inflution. However, system geometry and compoence. The same general agreement has not been nents for surface piercing propulsion is quite a reached on the minimum value that must be conventional concept for it includes only the attained to avoid scaling problems. This is due classical propeller-shaft-rudder layout. to the different kind of Froude numbers that have been used. Further research is required to Current extrapolation methods are focused identify the minimum value of Fn which must on two items: be achieved during open water tests. Nevertheless, provided that open water tests are per1) extrapolation to full scale of the model formed at Fn beyond the threshold value, the open water performance; Froude number identity can be avoided during 2) extrapolation to full scale of the propeller the tests without affecting the full scale perhull interaction (wake fraction, thrust deformance of the propeller. duction and relative rotative efficiency). According to Shiba (1953), surface tension Comments on item 1). This extrapolation plays its role when the propeller is about to be procedure should take into account the influfully ventilated. Complete ventilation is a ence of two additional parameters, namely the rather sudden phenomenon that can be correFroude number, Fn, and the Weber number. lated to a certain value of J called the critical The influence of the Froude number is relevant advance coefficient JCR. The critical advance as the propeller acts at the interface between air coefficient can roughly be located in the midand water much like hulls. The Weber number dle of the transition region and the sudden drop is a ratio between inertial and surface tension of KT and KQ identify its position. Indeed Shiba forces. Its influence can be easily foreseen for found a correlation between Wn and JCR for a an SPP, which continually pierces the water single propeller. Later work on this matter surface. (Ferrando and Scamardella 1996) acknowledged the influence of the Weber number on The first comments on the role of Fn are the location of JCR, and the existence of a due to Shiba (1953). In his investigation he threshold value of Wn beyond which its influpointed out that gravity affects the shape of the ence disappears. Unfortunately there is not sufair cavity through the Bernoulli equation that ficient evidence to suggest the existence of a can be enforced at the boundary between water unique threshold value. Further research on this and the atmosphere vented cavity. Accordingly, subject is therefore strongly needed. Anyway, the influence of Fn based on the diameter of the provided that open water tests are performed at propeller as the length parameter and nD as the Wn beyond the threshold value, the Weber

7 number identity can be avoided during the testsbig expenditures like extensive tank testing or without affecting the full scale performance of high standard sea trials. There is not a single the propeller. known case study in the open literature providing data for extrapolation purposes. As far as is known, provided that open water tests are performed in agreement with the In some favourable circumstances EHP is above requirements, the performance of SPPs determined by means of towing tests, while for can be scaled in the same fashion as conven- the great majority of applications an effective tional propellers, i.e. by applying a Reynolds horse power estimate is performed using Savitnumber correction only. sky’s (1964) equations or something similar. In both these instances there is not enough data Comments on item 2). The propeller is usu- for a true extrapolation of the propeller-hull interaction. The absence of sea trial results preally located quite far from the hull (i.e. 1 to 2 diameters away) and the projection in the verti-vents the development of a reliable full scale cal plane of the immersed hull surface near the extrapolation. propeller is small. As a consequence, the value of the thrust deduction factor for SPPs is negligible or zero. Furthermore, from a theoretical 4.4 Oscillating Propulsors point of view the phenomenon of augmented resistance for this kind of propulsion can be Oscillating propulsors are still at the researtreated with the same physical model used for ch and development stage. Only very few acconventional propellers because no additional tual applications have been built and tested at thrust affecting devices are involved. Hence thefull scale and all of these have been for small same scaling procedure adopted for conven- craft. Examples are the small boat driven by an tional propulsion can apply also to SPPs. oscillating propulsor and outboard motor designs built by Isshiki et al. (1987), the experiSurface piercing propulsion is mainly em- ments done by a group led by M. Triantafyllou ployed on planing or semi-displacement craft, at MIT, and various oscillating propulsor designs used on human powered craft (e.g. exploiting the ventilation of the transom. As SPPs operate far behind the hull the contribu- Bennett 1996), including examples on human tion of the potential wake is small. On the other powered submersibles (Skidmore et al. 1989). hand, the viscous wake cannot be neglected, As a result, extrapolation methods for powering prediction for oscillating propulsors are because part of the propeller lies in the rare. boundary layer of the hull. Considering that surface piercing propulsion does not include any additional wake modifying device and that Other propulsors that operate with a similar the physics of the wake production process is propulsive mechanism to the oscillating prothe same as in the case of conventional propul- pulsor are trochoidal propellers (a type of cysion there are no obstacles to using the same cloidal propeller), the most recent example of scaling procedure as that used in the case of which is the Whale Tail Wheel (Anon. 1998) conventional propeller arrangements. being fitted to the inland waterways vessel As regards relative rotative efficiency there “Ludwina”. Such propulsors have been studied is no particular reason that prevents the appli- at model scale (e.g. Manen 1973, Bose and Lai 1989, Riijarvi et al. 1994). A special type of cation of traditional scaling methods. oscillating propulsor, that might be described Setting aside the preceding considerations,as a hybrid between a conventional and an osalmost all applications of surface propulsion cillating propulsor, is a propeller with cyclic are in the pleasure craft trade. Generally, the pitch control (e.g. Gabriel and Atlar, 1998). budget for the boat design does not allow for

8 Estimates of the full scale performance of fied ITTC 1978 method have been used to preships fitted with oscillating propulsors have dict full-scale performance from ship model been made by Lai (1990) and Yamaguchi and test results. These predictions have been comBose (1994). These have been based on con- pared with the results from sea trials or from ventional approaches to powering prediction statistical sailing data. using estimates of the hull efficiency to account for propulsor/hull interaction. Yamaguchi and Bose (1994) assumed neligible Table 2 Pre-swirl energy-saving devices. hull/propulsor interaction leading to a hull efficiency of one. The rationale for this estimate Name of pre-swirl device Energy- Energy-saving saving rate % was that as an oscillating propulsor design mechanismmodel test/trial normally has relatively light loading, they are Reaction fin 4-8 / 4-9 PRI usually large devices with relatively large Novel integrated duct 4-5 / 4-6 IPI swept area and large span relative to the ship Inflow compensate nozzleIPI, AFS 6-11 / 8 beam and/or draft. In view of this much of the Simplified compensate IPI, AFS 4-9 / 4-9 nozzle propulsor is not in close proximity to the hull IPI, AFS 4-9 / 4-9 Wake-adapted duct and the hull/propulsor interaction was assumed IPI 4-6 / 3-9 Hydrodynamic fins to be small. For the ship under consideration, Fore-propeller hydrody4-7 / 4-7 IPI this was a conservative estimate as the hull efnamic fin sector ficiency was above one for the comparable PRI 4.5-12 / 5-8 Thrust shaft brackets PRI 5-8 / 5-8 conventional propeller system. Lai (1990), on Sheathed shaft bracket 2-3.7 / 2-3.7 IPI, AFS Hydrodynamic partition the other hand, assumed similar values of the plate wake and thrust deduction fraction as the con- Aperture fin 2-4 / 2-4 PRI ventional propeller system for the oscillating Stern-appended fin 3/3 IPI, AFS propulsor proposal when considering the pow- Flettner rotor at stern post 8/ IPI, AFS 3-4 / Fore-propeller vane-wheelIPI, AFS ering performance of four ships. This was done in the absence of more detailed data as the hull efficiencies for the four vessels varied from 0.92 to 1.12. Since 1984 MARIC has developed more than ten ESDs such as the composite device of 4.5 Pre and Post-Swirl Devices simplified compensate nozzle (SCN) and Costa Since the 1970s the development and appli-propulsion bulb (CPB), thrust shaft brackets cation of hydrodynamic energy-saving devices (TSB), fore-propeller hydrodynamic fin sector (FPHFS) etc. (Qian et al., 1992a, Zhou et al., (ESD) have been demonstrated all over the 1990, Qian et al., 1992b). Eight different types world. Tables 2-4 show energy savings from pre- and post-swirl devices. The propeller in- have been put into use on more than 200 ships over the range from 500 to 150,000 dwt with flow compensate nozzles developed by Schneekluth have been installed on more than an annual 50,000 ton fuel-savings. In all cases 600 ships (Schneekluth, 1986). Powering per- a modified prediction method based on the formance prediction for full-scale ships with ITTC 1978 method has been used in the prediction of the full-scale performance from the pre- and post-swirl devices have been made based on model test results. For most of these ship model test results. Practice indicates that the ITTC 1978 method and the two dimension- the energy-saving rate (ESR) or the predicted al Froude method have been used to predict thespeed of the ships installed with ESDs correlate full-scale performance for ships having a large well with the full scale data. range of block coefficients. At MARIC several different prediction methods based on a modi-

9 Table 3 Post-swirl energy-saving devices. Name of post-swirl devices

Energy- Energy-saving rate % saving mechanismmodel test/trial Fixed guide vane after RRE 4-6 / 5 propeller Vane-wheel RRE 5-15 / 6-12 Rudder-appended thrust RRE 4-5 / 4-8 fins Reaction rudder (asym. RRE 2-4 / 2-4 rudder) 2-4 / 2-4 DVP COSTA propulsion bulb DEP 2-5 / 2-5 Propeller cap fins 5.1 / 8.7 Eddy-eliminating compo- DEP site propeller 5-10 / Integrated duct-vaneRRE wheel 10 / Rudder-appended Flettner PAT rotor

described in section 5.4. Included here are some observations from a set of tests done over a range of Reynolds numbers which include, in particular, some trends at higher than normally tested values of Reynolds number.

Extensive propulsion related tests were recently done on a tugboat model at Bassin d'Essais des Carènes. This tug-boat was a twinscrew vessel with two ducted propulsors. A rudder was placed closely behind each shroud (about 1/3 of the chord of the rudder and less than 1/2 the length of the shrouds behind) so that the interaction between rudder and propulsor had the potential to be strong. The slope of the aft end of the boat was high, so that the propulsors were working in inclined flow conditions. The distance between the shroud and the hull was small which created the possibility of detached Table 4 Pre and post-swirl energy-saving de- flow in some conditions and strong interaction between the hull and propulsors. Due to vices. restricted space, the propulsors were highly loaded. Name of pre and postswirl devices

Energy- Energy-saving saving rate % mechanismmodel test/trial IPI, RRE / 9-18

The model was about 4.5 m long, equipped with a propeller of about 20 cm diameter. The Hydrodynamic fins & speeds tested, based on Froude number scaling, guide vane-wheel were about 1.5 m/s for the transit speed and COSTA propulsion bulb DVP, RRE 4-14 / 4-7.4 about 0.5 m/s for high towing conditions. Open & rudder-appended water, resistance, self-propulsion and varying thrust fins Duct-thrust fins after proDVP, RRE 4 / 4-5 load tests at all advance ratios were performed, peller including measurements of the axial thrust of COSTA propulsion bulb IPI, AFS, 6-12 / 4-12 one shroud. The boat was a twin screw vessel, & simplified compen- DVP which implies that the prediction and sate nozzle extrapolation methods based on open water, Assembly propulsion de- IPI, RRE 5-15 / 5-15 vices resistance and self propulsion tests should not be Composite energy-saving IPI, RRE 8-14 / 8-14 affected by the problems of high wake as would technique be expected on single-screw ships. Additional tests were done to investigate PRI – give a pre-rotation to the propeller inflow some problems related to prediction and IPI – improve propeller inflow extrapolation for ducted propellers and these AFS – alleviate flow separation RRE – recover rotational energy from downstream have relevance also for some other types of DVP – decrease viscous loss after propeller cap unconventional propulsors. These were: DEP – decrease eddy after propeller cap - towing conditions at low speed (from 0.2 PAT – produce additional thrust to 0.4 m/s) to detect any scale effects due to low Reynolds number, especially on the 4.6 Ducted Propulsors shrouds; Some methods for the powering performance - self propulsion tests without rudders of ships fitted with ducted propellers are (transit and towing), to separate effects from

10 interaction of the propulsor with the hull and The open water test of the whole propulsor shows a Reynolds number effect on KT and KQ the propulsor with the rudder; - resistance tests of the hull plus the two that affects the propeller characteristics in the shrouds, without rudders and without speed range up to 4.5 m/s. The extrapolation of propellers; the characteristics of the propeller at the Froude speed (about 0.5 and 1.5 m/s for towing and - resistance tests of the shroud alone without the propeller ("open water test"). transit) would not give the characteristics of the propeller at the higher speed tested. The drag of the shroud in the "open water" The plot of KT of the propeller (KT obtained test (figure 1) clearly shows a separation with the shroud but not including the thrust of phenomenon at low speed and a transition speed. the shroud) versus KQ (figure 2) for the open A speed of at least 2 m/s is necessary to get water tests and the two conditions of load stable values away from this transition. This varying tests (with and without rudders) shows problem seems to be avoided in the behind that over the whole range of advance ratio (from condition where the values are different. bollard pull to almost no thrust) the three curves 30 are almost identical. This indicates that the relative rotative efficiency is very close to 1 with 2 a precision of less than 1%. The dispersion ^ y t 20 i between the curves is due to precision c o l uncertainty in the measurements, the importance e v / g of which becomes relatively smaller for points a r d close to bollard. t 10 c u d

0.5

0

0.4

0

1

2

3

4

5

velocity m/s

Figure 1. Shroud drag. The circles denote the drag from the “open water test”; the squares are the values found from the resistance of the hull tested with and without the shrouds fitted.

r el l e p 0.3 ro p t K 0.2

0.1 0.05

0.06

0.07

0.08

0.09

torque coefficient Kq

0.1

The tests of resistance with the shrouds and without propellers showed that the shroud drag Figure 2. Thrust coefficient of the propeller was about 25% of the total. Comparison of these versus torque coefficient. The open squares are values with those obtained from the open water open water values; the solid squares are the test on the shrouds, showed that the interaction behind condition with the rudder; and the solid term was between 1 and 1.5% (the total drag of circles are the behind condition without the rudder. the hull plus shrouds was 1% to 1.5 % higher than the value obtained from the addition of the naked hull resistance and open water drag of the Nevertheless, looking at the detail of the curves in figure 2 for high loads, there is an shrouds). This interaction term is very small important difference between the curves. despite the geometry. Although the torque coefficient plots against advance coefficient are not shown here, the

11 torque coefficient in the bollard condition for theonly is plotted in figure 3) for open water and open water and behind condition (KQ = 0.0915) behind condition with and without the rudder was the same with a precision of less than 0.2 % shows again an effect of the rudder. for both propellers. However, KQ becomes 0.5 0.0945 in the bollard condition with rudders (an increase of 3%). Also, the KT of the propeller t 0.4 c u shows little change from 0.422 in open water to d f o 0.3 0.414 in the behind condition without rudder (a t n e decrease of 2%), while there is an increase to i c 0.2 i f 0.438 (plus 4%) with the rudder. Clearly, the f e o rudder has a "postswirl" effect on the propeller c 0.1 t s rotor, increasing the thrust by 6% and the torque u r h t 0 by 3% between the two behind conditions. As a result of this postswirl effect, thrust is increased -0.1 more than torque leading to an increase in 0.2 0.25 0.3 0.35 0.4 0.45 efficiency. These arguments together with figure thrust coefficient of the propeller 2 show that the effect of the rudder on the relative rotative efficiency is far more important than the difference between operating in open Figure 3. Thrust coefficient of the duct versus water and the behind condition. thrust coefficient of the propeller. The open squares are open water values; the solid The plots against advance coefficient (see squares are the behind condition with the figure 4 for thrust coefficient against advance rudder; and the solid circles are the behind coefficient) show also that a KT identity for the behind condition with rudder, based on the rotor condition without the rudder.

alone is meaningless for low advance ratio: KT Direct comparison changes from 0.438 in the bollard pull condition the to 0.426 for JS = 0.3 while KT in open water has a betwwww.kriso.re.kr/ITTC/index.htmleen maximum value of 0.422. A KT identity without open water and behind condition with rudder gives a confusing picture which at first sight rudder is also problematic as a value of KT = 0.414 is obtained in open water for J0 = 0.25 (i.e.seems to show that there is no effect on the less than the open water value and leading to a shroud at transit speeds as a result of operating in negative wake). A KQ identity could be done in the hull wake and in an inclined flow, but that this latter case, as the bollard KQ is the same in the effect at towing speeds is large. This is a false conclusion since comparing the open water the open condition and the condition without rudder, but the slope of the KQ versus JS is too results with the behind condition curve without small to be precise at least for small J. When a the rudder shows a shift in the thrust of the plot was drawn (not presented here) showing KTshroud. However, the thrust values at bollard for of the propeller versus KQ, for the open water these two conditions are within 1%. Everything condition and behind condition only, but with appears as if the shroud is "seeing" a flow having less speed than the propeller rotor itself, that is to additional points from propulsion tests at smaller say as if the wake fraction "seen" by the shroud Reynolds number (simulating towing at full is a little bit larger than that "seen" by the scale speeds of 2 to 4 knots), there was a big discrepancy between some of these last points propeller rotor. Is this due to a boundary layer effect of the hull on the pressure distribution of due to scale effects and/or large errors in the shroud? Direct comparison of the open water measurement of very small forces. results with the behind condition with rudder gives then a confusing picture simply due to the The two plots of KT of the duct versus KQ and fact that for the same flow around and in the KT of duct versus KT of the propeller (the latter

12 shroud, the rudder gives a shift on thrust and torque of the propeller rotor (postswirl effect) 4.7 Partial Ducts which is an effect as large as the effect of the hull on the shroud. But the postswirl effect Wake equalizing ducts were developed by doesn't change the flow around the shroud as theSchneekluth (1986). They consist of two nozzlebollard thrust of the shroud indicates. shaped half ring ducts which are installed on both sides of the stern ahead of the propeller. These latter points have further implicationsTheir diameters are about the same as the radius on the identification of wake fraction from the of the propeller and their chord is smaller than the diameter. Sometimes, only one duct is fitted thrust identity. From the plots of thrust coefficient against advance coefficient (one of to the stern on one side of the propeller. which is shown in figure 4) the following can be seen. Comparing open water and behind The wake behind single-screw ships is noncondition with rudder gives a wake fraction of homogeneous (i.e. there are very small velocities about 10% for J around 0.3-0.4 (towing at the top of the propeller disc) and it is assumed condition). However, in the case without rudder,that improving the homogeneity of the wake will wake fraction is below 3%. The value of 1-w is improve the propulsion efficiency (essentially surprisingly more than 1 in this region of the open water propeller efficiency). Optimising advance ratio. There seems to be a difference of the angle of the partial duct to the stern under regime between these low J values and the load conditions (accelerating the flow in the top region of transit. part of the disc and slowing it in the lower part) is said to improve the homogeneity of the wake. 0.45 r e l l e p o r p f o t n e i c i f f e o c t s u r h t

0.4 0.35 0.3

Power savings of 5 to 10 % have been reported (Schneekluth 1986), despite the fact that this appendage creates an additional viscous drag. On the other hand, this partial duct can create a thrust and reduce eventual separation.

The extrapolation of powering performance of wake equalising ducts fitted on large single0.2 screw ships are a typical example of the 0.15 difficulty of scaling data from model tests of 0 0.2 0.4 0.6 0.8 1 1.2 short appendages. The reason is that these advance coefficient J devices are usually fitted to very large ships, which means a very low towing speed in the tank to comply with Froude number scaling of Figure 4. Thrust coefficient of propeller versus the full scale ship speed. For a given ship speed, advance coefficient. The open squares are Froude number scaling imposes a reduction of open water values; the solid squares are the the model speed for a smaller scale model; behind condition with the rudder; and the whereas the maximum model size remains solid circles are the behind condition withoutbasically the same for each facility. the rudder. Reynolds effects can act in the opposite way: In the region of transit speeds, the effect of the rudders is about 0.04-0.05 on the advance - separation which occurs on the model may coefficient of the rotor. The KT identity based not occur, or may occur over a smaller region, upon the total KT may work because of a at full scale, hence a partial duct effective at cancelling effect between the rudder and the reducing this separation on the model may be shroud. less effective at full scale; 0.25

13 - the boundary layer is relatively far smaller at full scale than on the model, hence the size and inclination of the duct optimised for the model may not be optimal on the full scale; - friction of the partial duct is exaggerated on the model due to the very low Reynolds number on such small appendages (values of Reynolds number on the appendages may be about 50,000 only).

underestimated even for a conventional propeller. - The propeller open water efficiency actually decreased slightly with the duct for this model, in contradiction to the principle of operation of wake-equalising ducts. - The wake fraction and thrust deduction fraction changed slowly, but did not reach a constant value over the range of speed tested. In these tests the ducts reduced suction and increased wake so then the increase in propulsive efficiency came from an increase in hull efficiency.

Friesch and Johannsen (1994) did tests in a large cavitation tunnel (HYKAT) on one model of the size normally used in model towing tank tests at speeds from 3.4 to 5.5 m/s in order to investigate these Reynolds number effects. As In conclusion, partial ducts may result in published results over a large Reynolds number energy saving at full scale, but this was not, and range are rare, some of the implications of this probably cannot be proven by model tests at work are described; however the Committee Froude speed by use of the present testing cautions that published records from many moreprocedure. These tests done at higher Reynolds tests of this type are needed for different types ofnumber in a large cavitation tunnel showed an unconventional propulsor in order to reliably erratic behaviour of the energy saving with test identify true trends. The equivalent Froude speed speed. for this model was about 1.2 m/s and comparisons were made with towing tank tests No conclusion could be drawn from the done at this speed. Their results are as follows: Froude scale speed tests as they were affected by unrealistic levels of separation and other Reynolds number effects. Although not - The wake equalising duct did not reduce the necessarily true for all partial ducts, the savings drag of the ship (without the propeller). indicated in this set of model tests were - There was a reduction of power required forapparently due to increased hull efficiency; the propulsion, but the behaviour of it was very open-water efficiency seemed to decrease a bit Reynolds number dependant: from 2.5 % at and the wake was not more homogenous. Tests Froude speed; to 9.6% at 3.5 m/s; to only 6% where separation is occurring on the hull without at 5.5 m/s. the propeller and where this is reduced/ suppressed with the propeller, can lead to an - The propeller characteristics were affected underestimation of suction (thrust deduction) very little by the presence of the ducts (slight and an overestimation of wake fraction (i.e. an increase of thrust and torque coefficients), overestimation of hull efficiency) when analysed using ITTC 1978 methods. but, alarmingly for the testing community, they became stable for speeds higher than 3.5 More generally, it is difficult to reliably m/s only (at which speed the torque predict full scale performance of this type of coefficient was reduced by 20% compared device from model tests at equivalent Froude with the value obtained at Froude speed for a speed in the towing tank and even tests at higher comparable thrust coefficient). This shows speeds in the cavitation tunnel can give uncertain that open water tests can be used, but that trends. To address these issues, and to bridge the Reynolds number effects on Reynolds number range, further reliable the correlations are needed between: tests done in characteristics of the propeller may be

14 large cavitation tunnels at even higher speeds; towing tank tests done for some equivalent speeds at the lower end of the cavitation tunnel range; and full scale trials results.

done by Rains et al. (1981). They presented a semi-empirical approach to estimate the total drag of a podded drive and comparative power estimations. The investigation also included a model test programme with a DD-963 class destroyer model fitted with 3 different drive Tests performed on a large tanker in the GTH at the Bassin d’Essais des Carènes at full systems: a conventional twin drive; a nonspeed (11 m/s) with partial ducts optimised in azimuthing twin tractor; and a pusher type podded propulsor. The tractor type drive has its model tests at Froude speed showed a propeller ahead of the pod while this is recompletely different effect of the duct on the versed in the pusher type. The model tests inwake in the two situations. This implies that optimisation of the inclination of the duct shouldvolved resistance tests with and without the propulsion appendages, open water tests of the be done from the results of high Reynolds propeller in isolation, which was the same for number tests. Propulsive tests showed no energy all three propulsion systems, and selfsaving for the high Reynolds number tests. propulsion tests. In the latter tests, only torque Specifically: for each propeller was measured and the pro- towing tank tests showed energy savings of pulsive coefficients analysed appear to be 1-2% depending on the propeller tested and based on the torque identity method. Although the performance comparison of the three prospeed simulated; pulsion systems was presented in terms of the - there was a difference in nature of the wake appendage drag, effective power and delivered (without the partial duct) between tests at thepower on the full scale, no specific reference Froude speed and those at higher Reynolds was made as to how they were extrapolated to full-scale. number, in particular the occurrence of a region of high vorticity in the region where the duct was to be fitted; Minsaas (1988) investigated a tractor type non-azimuthing Z-drive, with a steering flap at - by subtracting the wakes, the effect of the the tail of the drive strut, as a promising alterpartial duct appeared to be completely native to water jet propulsion for speeds up to different from the two tests at differing 55 knots. Neglecting the cavitation and prospeeds, the smoothing effect being cancelledpeller induced drag, semi-empirical formulae at higher Reynolds number; were given to calculate the viscous drag of the unit due to the body, the strut within the slip- pressure fluctuations were improved with stream and strut outside the slipstream. In these the duct fitted and reproduced in the higher formulae emphasis was placed on the considerReynolds number tests similar signals to the able scale effect on the resistance estimation as well as on its validity for only smooth surfaces. full scale. Reynolds number effects in tests performedLater Halstensen and Leivdal (1990) described at speeds of about 1 m/s may have more effect the development and full scale application of on the predictions than not using the Froude this type of podded drive, called “SpeedZ”, to scaling identity, especially for ship testing at Fjellstrand catamarans “Sleipner” and “Draupner” and to Westamarin SES catamarans “Suvery small Froude numbers. per Swede” and “Super Dane”. The study reported on the open water tests done with sub cavitating and partially cavitating propellers 4.8 Z - Drives (Podded Propulsors) fitted to the propulsion unit in the cavitation tunnel at MARINTEK, as well as towing tank An early investigation into the power pretests with the naked hull. Although no details diction of vessels fitted with podded drives was

15 of the power performance prediction were dis- screws. cussed, the complexity of the resistance prediction at full scale due to the effect of the proThe final group of tests, which were perpeller was reported. From this study it appears formed at MARIN, involved open water tests, that the full-scale estimations were made basedhull resistance tests with the same hull and selfon the naked hull resistance tests from the propulsion tests. However, the propeller used towing tank and the efficiency of the total pro- was a new design developed by the Krylov pulsion unit obtained from the cavitation tunnel. Ship Research Institute which was customised The latter included a resistance correction for to the Azipod and hull. Two sets of open water the full scale. tests were done with this propeller, with and without the Azipod housing. The calculated net A recent study on the power performance efficiency of the Azipod in the open water tests, aspects of azimuthing podded drives was pre- which was subject to low Reynolds number sented by Kurimo et al. (1998) on the devel- effects, was found to be slightly lower than that of the propeller alone at the corresponding adopment of two 14 MW Azipod units for the cruise ship “Elation”. The study reported on an vance coefficient. The hull resistance tests were done without the Azipod units. Following the extensive model test programme done in a wind tunnel, towing tank and cavitation tunnel self-propulsion tests a value of 5-7% of power at three institutions. The experiments done in saving was predicted on the full-scale when the VTT towing tank involved open water tests compared with the original vessel fitted with a with a stock propeller, hull resistance tests andconventional propulsion system. Although no self-propulsion tests. The purpose of the open details were given, it was indicated that the water tests was threefold: to determine the in- prediction was carried out according to MARIN’s “Standard” method for thrusters. teraction coefficients between the propeller and the unit; to compare the performance of two pusher and one tractor type unit; and to provide In this study, the use of a special dynaa basis for the analyses of the propulsive coef- mometer for the open water tests highlighted ficients. The open water tests were done using the importance of the complex interaction bea special dynamometer that could separately tween the Azipod housing and the propeller for measure the thrust of the propeller and resis- a tractor type unit when the propeller thrust and tance of the Azipod housing. Based on these the resistance of the pod housing were meastests and separate wind tunnel experiments, ured separately. Therefore it was argued that which involved resistance measurements of the adopted scale effect correction methods various components of the propeller housing, aused in the extrapolation of the thruster propultractor type Azipod was optimised. The resis- sion test results often concentrate on the scaltance tests were performed by using an earlier ing of the measured “resistance” of the housing. hull model which had appendages for conven- This may be acceptable for the pusher type tional twin screw drive. However these apAzipod, but not for the tractor type due to the pendages were removed and the optimised complex interaction between the propeller and Azipod housing was fitted. These tests were pod housing which results in an increased local performed with and without the housing such pressure field behind the working propeller. that the resistance of the housing could be ob- Therefore, emphasis was placed on the gap tained. In the self-propulsion tests, the total between the Azipod housing and the propeller thrust of the each Azipod unit was measured. Itsince low gap sizes would increase the resiswas not possible to measure the propeller thrust tance of the housing and the measured thrust of and the resistance of the pod separately due to the propeller, if they were separately measured, the small scale of the model. The measured de- although the net thrust of the azipod unit would livered power indicated a saving of 3-4% over be the same. Based on this argument, it was the same hull driven by conventional twin claimed that the difference between the total

16 thrust of the unit and the thrust of the propellerrepresents one of the earliest applications of in the tractor type of unit should not be related inviscid blade design theory and, as a result, to Reynolds number dependent viscous effects introduces the problem of scaling. An exalone. Also, more detailed methods for the ex- trapolation method developed by the ITTC trapolation of test results from tractor type de- community has been widely used for commervices are recommended to be developed. cial ships having an open screw propeller to predict full-scale powering performance. The Mewis (1998) has discussed recent devel- ITTC 1978 methodology uses resistance tests opments in large podded drives with a specific of the ship model, open-water tests of the emphasis on the range of their efficiency im- model open propeller, and model-scale selfprovements and difficulties in power perfor- propulsion tests, to determine interaction coefficients, which are then scaled to predict fullmance prediction. This relates mainly to the tractor types, a factor which was also pointed scale performance. However, this methodoloout by Kurimo (1998). A part of the difficulties gy and the scaling of its interaction coefficients has been associated with the different configu- are based on particular commercial ship datarations used in open water tests and the treat- bases. The direct application of this methodolment of the thrust force measured during theseogy to complex propulsors such as ducted protests. In order to demonstrate these difficulties,pellers or pump jets, water jets and various open water tests of a tractor type Azipod drive novel forms of open screw propellers continues to be an unresolved issue. have been carried out in three different configurations as reported by the Propulsion Committee of this Conference. Any propulsor configuration and its performance are highly dependent upon the hyKurimo (1998) presented the results of sea drodynamic characteristics of the ship trials with the cruise ship “Elation” which werehull/marine vehicle. Energy caused distortions carried out in the Gulf of Finland in December are present in the ingested flow due to the skin 1997. The trial results involved the presenta- friction drag of the ship hull/marine vehicle tion of speed measurements, cavitation obser- and upstream appendages. It is in this envivations, pressure pulse measurements and ma-ronment that wake-adapted complex propulsors noeuvring tests. The trials indicated that at the excel. Complex propulsors provide more design options with which to meet the many addifull power of 2 x 14MW, the vessel achieved approximately 0.55 knots greater speed than tional performance requirements in these comthe mean value of the corresponding speeds of plex wakes. These new requirements include, the six previous sister ships, which have con- but are not limited to: 1. minimization of unventional twin-screw diesel electric drives. De- steady forces, 2. reduction in radiated noise, 3. spite this, the scatter between the predictions improved efficiency for heavily loaded propellers, 4. reduction of cavitation and eliminafrom the different model basins, which used different extrapolation methods, was large andtion of specific cavitation types, 5. improved designers were cautioned to treat test results off-design performance, and 6. enhanced shipmaneuvering characteristics. with great care to avoid optimistic expectations. The principles employed and the problems encountered in designing either an open screw 5 DESCRIPTION OF EXTRAPOLATION propeller or a complex propulsor are similar. METHODS AND TEST PROCEDURES However, the design methods employed are different for the two cases. First of all, the design of a wake-adapted complex propulsor 5.1 Momentum Methods having a duct and/or multi-blade rows treats the propulsor as a unit. Therefore, its modelThe propulsion of a ship or marine vehicle

17 scale testing procedure and scaling methodology will be different than that developed for an A relationship to determine required rotor open screw propeller. One present-day methodthrust is determined by integrating the moof designing complex propulsors includes a mentum in the stream wise direction between a combination of a vortex lifting-surface method station located far upstream of the propulsor with a computational fluid dynamics based and a station located after the body and can be through-flow analysis method.These two expressed asT = m (∆V ) where T is the rotor methods allow for hull interactions, viscous the and is V effects and blade-to-blade row effects to be in- thrust, m is the mass flow rate ∆ integrated change in stream wise momentum. cluded as well as three-dimensional effects such as streamline curvature, radial pressure An important assumption in this relationship is gradients, and secondary flow (Kerwin et al. that free-stream static pressure exists at each station. (1994) and Schott (1996)). An incoming velocity profile near the plane Interest in multi-blade row propulsors has of the propulsor and the predicted hull drag codeveloped because of increasing efficiency efficient are necessary to start the design procdemands for high-speed surface ships and submerged vehicles. The attractiveness of ess. This has been traditionally obtained from ducted propulsors for these applications is due model scale tests with corrections for variations in Reynolds number. In some cases, analytical to the ability to design a propulsor having a lower blade relative velocity and a higher effi- predictions are also being used due to more reliable computational fluid dynamic procedures ciency than can be achieved by an open proand their validation to specific geometries peller in these applications. The similarities (Larsson et al. (1998), Arabshahi et al. (1998), which exist between ducted propulsors and axial-flow compressors and liquid pumps have Zierke et al. (1997) and Stern et al. (1996). lead to the application of momentum analysis It is also necessary to determine the total methods. Nowhere are different testing methodologies more evident between an open pro- drag of the hull, which includes increments due peller and a complex multi-blade row marine to appendages, etc. The flow field solution repropulsor than in reviewing a momentum basedquires the calculation of the frictional drag on all the various propulsor components. It must design methodology. be remembered that the pressure drag of the afterbody, which is initially included only as a Early efforts in developing momentum based design methodology are given by Wisli- bare-body drag coefficient, is substantially modified by the addition of the propulsor. cenus (1960 and 1968), Henderson et al. (1967), and Bruce et al. (1974). This method However, this effect is initially estimated from the modified pressure distribution with the prodetermines the type, size, and design of any multiblade row propulsor from an optimisation pulsor/hull combination. of the mass flow through the propulsor to It must be emphasised that this momentumachieve design goals. An optimised blade-row spanwise-circulation distribution is obtained based design procedure is applied where there from minimising the energy losses in the pro- is incoming vorticity in the flow. The approach uses an inviscid calculation of the flow field pulsor and in the discharge jet. A detailed and is coupled with an energy analysis through analysis of the flow field at various stations through the propulsor follows by solving the the propulsor. Therefore, all the energy losses momentum, continuity and energy equations. through the propulsor are calculated and the efficiency of the rotor becomes the hydraulic The upstream boundary conditions are the mass flow rate, the momentum and the kinetic ener- efficiency, which is the ratio of the energy placed in the fluid to the shaft energy. Solugy of the ingested flow.

18 tions to the flow field are obtained by using the scale ship drag which must include appendage Streamline Curvature Method (Treaster, 1969).drag and propulsor/hull interactions. Current Today, this has been extended to a RANS soattempts to do this utilise CFD procedures, larlution using computational fluid dynamics ge water tunnel tests, and an extended analysis modeling. of propulsion data. This design procedure for complex propulsors does not require open water tests for ex- 5.2 Other General Methods trapolation methodology to predict full-scale speed and rpm. In fact, testing of these wakeOther extrapolation methods are being deadapted complex propulsors in uniform flow veloped using self-propulsion testing, the phisignificantly reduces propulsor efficiency and losophy of which could be considered for unintroduces corrections that are not well defined. conventional propulsors. For these methods an Therefore, only model-scale propulsion test open water test is not necessary, only a load data is required for extrapolation to full scale. varying self-propulsion test is required. This model test is conducted over a range of over/under loading conditions. This is usually A modern MARIN method based on a form achieved by varying the rotor rpm at constant factor (1+k) concept has been developed for ship velocity. Resistance test data or calcula- some ship configurations for the extrapolation tions are only necessary for the design. of propulsion test data to full scale. The form factor is the ratio of viscous resistance to the At these model scale conditions, the effi-flat plate drag based on the ITTC 1957 formula. In general, this factor can be determined for ciency can be defined as each hull form using low rpm self-propulsion measurement data. The scale effect on resis( RTm −FD )Vm effectivepower = ηD = tance (FD) is determined from the standard 2π qm nm deliveredpower equation including an incremental resistance coefficient (CA). The measured relationship and between the thrust coefficient (KT) and the apparent advance ratio (JV) from the propulsion 1 2 tests is corrected for both wake and propeller FD ={ (1 +k)(C fm −C fs ) −C A } ρ m Sm Vm 2 blade friction scale effects to predict power and From the model scale KT and JV curve, correc- rpm. This is effectively an ITTC 1978 method tions are made to account for Reynolds numberwith the wake fraction found from a statistical differences to full-scale first by skin friction method based on previous test results rather corrections to the propulsor components than directly through open water tests of the ( ∆K T , ∆K Q ) and secondly, by corrections to actual model components. the wake ∆ ( n ). It is very important to note that these corrections cannot account for separatingAn extensive analysis of typical combatant/auxiliary Naval ship data has been conductflows. ed by the Naval Surface Warfare Center, Carderock Division, to validate model test exIn summary, the model scale propulsion test provides a KT versus JV curve. This curve trapolation procedures (Karafiath, 1997). The must be corrected for Reynolds number effectsdeveloped prediction methodology uses a correlation allowance (CA) to account for differboth on the propulsor and on the hull which produces a wake scale effect. The weak point ences between ship and model roughness and of this or any other extrapolation methodology for other variables that influence the powering prediction. Thus, the propulsion test is conremains the determination/prediction of the full ducted at an overload condition with an added

19 tow force (FD) to overcome the effects of an of the propeller are scaled following ITTC additional frictional resistance and achieve 1978 procedure while no scale effect correction equivalent thrust loading. This is accomplishedis applied to the thrust/drag produced by the by varying the propeller rpm at a constant ve- pre-swirl stator. locity. One basic assumption of this method is that the efficiency of the model propeller is the Van et al. (1993) state that this method same as that of the ship propeller; this is approvides a better full scale prediction than the propriate for small wake fractions. This is ITTC 1978 procedure. This statement is supagain an ITTC 1978 type method, but disported by the close agreement between model penses with the form factor, open water tests and full scale predictions presented, that is not and other corrections. found by using the ITTC 1978 procedure directly. Both of the above methods have limited application for unconventional propulsors since It is well known that the ITTC 1978 correthe previous test data and correlation coeffi- lation procedure fails to correctly scale the percients, needed for the methods to work, do not formance of unconventional propulsion sysexist. tems, and this is due to two main causes. The first one is that whatever device is applied to the hull, it generally has a longitudinal dimension which is much less than that of the hull 5.3 Performance Prediction Methods For itself. Given the usual testing speeds, this proShips With Pre-Swirl Stators duces a Reynolds number corresponding to a laminar flow on the device. The second cause Van, Kim and Lee of KRISO (1993) proposed two alternative procedures for perfor- is that the model hull has a boundary layer that mance prediction for ships fitted with pre-swirldiffers from the full scale one both in thickness stators. The authors explicitly state that the twoand in velocity distribution. Therefore, the inprocedures, named A and B respectively, basi- teraction between the hull and the special devically follow the ITTC 1978 method. In answer ce can seldom be correctly reproduced at model scale if it strongly modifies the flow around the to a request for clarification Van pointed out that at their towing tank the form factor methodhull. is not used. Accordingly these procedures are In this particular case of the pre-swirl stator, not true variations of the ITTC 1978 correla- we can assume that the flow around the stern is tion methodology as they follow the 2-D ITTC not overly affected. This can be argued fol1957 approach. The two procedures are dis- lowing the principle of operation of the precussed separately. swirl stator that is designed to produce a rotational speed component opposite to that inMethod A. In this method, the propeller duced by the propeller. For this reason, testing and the stator are considered as a propulsion the stator together with the open propeller insystem and are tested together. This assump- sures a higher Reynolds number on the stator tion implies that in both open water and self- itself due to the propeller-induced velocities. propulsion tests, the thrusts of propeller and Probably this is the main reason of the better stator are measured simultaneously and their agreement of Method A performance predicsum is used as the thrust of the propulsion sys- tion with the model results. tem. The hull resistance is scaled according to the ITTC 1957 method (i.e. a form factor is not Method B. This procedure does not require used). the joint test of the stator and the propeller because the stator is tested with and considered part of the hull. On the other hand, it requires In his correspondence with the Committee, that a double set of resistance and selfVan stated that the open water characteristics

20 propulsions test are done with and without the General remarks. Both of the proposed methods are basically 2-D procedures that can stator. be regarded as variations of the ITTC 1957 corThe scaling process is again the two dimen- relation approach, but in principle these techsional approach of the ITTC 1957 method with niques could be used also with the ITTC 1978 an exception made for the determination of the3D correlation procedure. full-scale wake, which is performed by means of the following formula that closely resembles From a theoretical point of view, the prothat suggested by the ITTC 1978 correlation posed methods are acceptable when applied to procedure: special propulsive devices that do not considerably alter the flow around the hull. In other words, these methods are suitable for scaling wS =(tMO +0.04) the performance of propulsive devices whose C FS +C A + (wMO −tMO −0.04) +(wMS − wMO ) effect is mainly confined to altering the proC FM peller inflow, without affecting the pressure field around the hull. Since these methods adwhere: dress the scaling problem only from the poten= ship wake wS tial flow point of view, they are not suited to wMO = model wake without stator treat such devices that could produce a considwMS = model wake with stator erable variation of the pressure field acting on tMO = model thrust deduction without stator. the hull. Actually, a hull pressure-modifying device could alter the characteristics of the while the standard ITTC 1978 ship wake is: flow around the model hull, e.g. the extent of laminar separation. In this case, the proposed wS =(t +0. 04) methods will probably produce results as inaccurate as those of the unmodified ITTC 1978 (1 + k)C FS + ∆CF +(wM −t −0.04) procedure. The lack of sea trials results pre(1 + k) CFM vents a practical evaluation of the capability of these methods to correlate model test results The major difference compared with the with the actual performance of the ship. Thus, ITTC 1978 formulation is the term (wMS – wMO). further effort is required to validate these Since in the opinion of Van et al. (1993) the methods. main effect of the stator is the increase of the angles of attack of the propeller blade sections, From an experimental point of view, if the the stator action can be considered to be a two proposed techniques are equally reliable, it mainly potential phenomenon. Thus, the differ-appears that Method B would be preferable. ence in wakes with and without stator can be Actually, the testing procedure related to directly transferred to full scale. Method B requires additional resistance and

self-propulsion tests; this is its major drawback, If the flow on the hull is not overly affected but the procedure is straightforward and does by the presence of the stator, this assumption not require any special test rig. On the contrary, looks reasonable and this procedure is accept- an ad hoc test rig is necessary to simultaneable. The same would not be true in the case of ously measure the thrust of the propeller and of other devices that accelerate or decelerate the the stator as required by Method A. The time flow on the hull, like ducts, partial ducts etc. consumption of this technique is comparable Van et al. (1993) state also that Method B ex- with that of standard tank practice, but the rehibits a good agreement with power savings quired special equipment would not be feasible obtained at model scale. or available in all of the towing tanks.

21 5.4 ITTC 1978 Modified Methods for Ducted Propellers.

(1984). In the case of the open water screw test, the equivalent open water velocity in which the screw works in the nozzle must be known. The velocity change is expressed by a ∆J correction

Some ITTC 1978 modified methods have 1−τ U been proposed for extrapolation of powering (1 + 1 + τCT ) ∆J = N = J performance of ducted propellers. Stierman nD 2τ (1984) presented three methods: where the nozzle is considered as an appendage; where where τ is the thrust ratio TP/TT. The purpose of the nozzle is treated as a part of the propulsion the ∆J correction is to find the correct subdiviunit; or where the screw, nozzle and hull are sion of the wake fraction into a potential and a treated as three interacting objects respectively. viscous part. Nozzle as an appendage of the hull. The In this method, the influence of the nozzle nozzle is primarily considered as a flow regu- on the screw is barely taken into account. A lator. The resistance test is done with the noz- nozzle limits the radial outflow of a screw and zle behind the hull and the open water test withtherefore, the KT/KQratio will not be measured the screw alone. correctly during the open water test.

The thrust deduction fraction is defined as Nozzle as a part of the propulsion unit. Here the open water test is done with the screw TP − R H + N + nozzle system and the resistance test is done t= TP with the naked hull. The thrust deduction fracwhere Tp is the propeller thrust and RH+N is the tion is defined using the naked hull resistance, and the total screw + nozzle thrust resistance of the hull with nozzle. The wake fraction is found by propeller thrust identity V − VA T + T N − RH t = P w= TP + TN V where V is the model speed and VA is the advance velocity of the screw. The total resisIn this case, VA in the defining formula of the tance of the hull with nozzle is extrapolated by wake fraction is the entrance velocity into the subtracting the estimated model nozzle resis- screw + nozzle system. The naked hull resistance from the measured total resistance, scal- tance is extrapolated according to the ITTC ing the resistance of the naked hull according 1978 guidelines. The scaling of the screw to the ITTC 1978 method, and adding again the thrust and torque from the open water test is estimated full size nozzle resistance. It is as- performed with the same ∆KTP and ∆KQ correcsumed that the formula to estimate the nozzle tions as proposed by the ITTC 1978. The nozresistance does not introduce too large errors zle thrust coefficient KTN should also be scaled. because the nozzle resistance only amounts to aThe resistance difference ∆K is roughly estiTN few percent of the total resistance. Note that mated using a flat plate friction line and a forthis is in contrast to the tests described in sec- mula according to Hoerner. The interaction cotion 4.6. efficients are extrapolated using the ITTC 1978 method or by considering the effective power The scaling of the propeller characteristics of the screw + nozzle system in their normal is performed in exactly the same way as outposition and in a position far behind the hull. lined in the ITTC 1978 method. The interaction coefficients are extrapolated either according to The objections against the method are twothe ITTC 1978 or by the method of Stierman fold:

22 (1) the nozzle thrust must necessarily be measThis method has been used to predict the ured during the self-propulsion test; (2) the action of the screw + nozzle system performance is of ships fitted with energy saving different in open and behind conditions. devices such as the simplified compensate nozIn the behind condition the nozzle produces a zle; COSTA propulsion bulb; thrust shaft brackets; fore-propeller hydrodynamic fin seclarger thrust due to the contracting inflow at the stern. For instance, in an open water condition tor; COSTA propulsion bulb and rudderτ = TP/TT = 0.90 may be found, while τ = 0.70appended thrust fins; composite energy-saving in the behind condition. Such an effect can be technique; etc. seen also on a twin screw ship even with no The size of MARIC’s towing tank is contracting inflow, perhaps due to boundary layer or inclined flow influences. Due to this 70*5*2.5m. The length of the geosim ship discrepancy, it is incorrect to determine the models were 3.5 to 4.5m. The diameters of wake fraction by using the total thrust identity single and twin-propeller models Dm are bigger than 0.12m and 0.11m respectively. The folaxiom. lowing method is used for the energy saving Screw, nozzle and hull as three interacting devices described in section 4.5 and tested at objects. To assess the interaction between theMARIC. The method utilises resistance, selfscrew and the nozzle, an open water test is car- propulsion and open water tests. The open waried out with the ducted screw. The resistance ter test is done with the propeller alone; the resistance and self-propulsion tests are done with test is performed with the nozzle behind the and without the energy saving devices. This hull to take into account the interaction between the nozzle and the hull. The thrust de- testing procedure is described in the following paragraphs. duction and wake fraction are defined as in method 1, by making use of the hull + nozzle Resistance test. The resistance test of the resistance, the screw thrust measured during the self-propulsion test, and the entrance ve- single-screw ship model is usually done locity VA into the screw disk. The open water without the appendages such as bilge keel etc. velocity VA in which the screw acts is the sum Measurements are made of the total model reof the translation velocity of the screw + nozzlesistance Rm, towing speed Vm and the water temperature tm at the same loaded condition as system and the nozzle induced velocity. The the propulsion test. To get a reliable value of screw must be out of the nozzle. The KTP/KQ ratio remains constant, but the entrance veloc- (1+k) more data are taken in the low speed area ity is changed. This means a shift of the J-axis than elsewhere. For the power prediction of the of the open water diagram with a ∆J-correction,full-scale ship with the above mentioned enerstanding for the dimensionless nozzle induced gy saving devices, the resistance test of the ship model is usually done with and without velocity. The nozzle induced velocity can be calculated by momentum methods, and vortex the above mentioned energy saving devices to compare their influence on the resistance perring or sheet methods. formance. The resistance of the hull and the nozzle is separately extrapolated, as described in method Open-water test. An open water test is done for the propeller model used in the pro1. The propeller characteristics are scaled using pulsion test. In this test the propeller submerthe well-known ∆KT and ∆KQ corrections. gence is bigger than its diameter. The advance speeds vary from 0 to the case of zero thrust while the rotational speed is kept constant 5.5 The Modified Full-Scale Performance during the test. The results of the open-water Prediction Method (MARIC Method e.g. test are used in the analysis of the selfZhao et al., 1988).

23 propulsion test results. The Rn of the propeller model in the open-water test should reach the and critical value 3*105 as far as possible. For the power prediction of the full-scale ship with the nS = nm / λ1/ 2 above mentioned energy saving devices, the open-water test is usually carried out without the energy saving device. The following equations can be easily derived: Self-propulsion test. When performing the self-propulsion test, the model is tested at a 2 minimum of 5 different speeds. At each speed PES = RTSVS andRTS =0.5CTS ρ SVS SS 3.5 the propelling forces Zm can be varied by PDm = 2πQmnmλ ρ S / ρ m changing the propeller revolution rate to enable the skin-friction correction FD to vary. The When the ITTC 1978 method is used thrust produced by the propeller should meet the following condition: CTS =(1 +k)C FS +CW +∆C F

Tm (1 −tm )+Zm = Rm

and

The ship self-propulsion point is given by the = CW CTm −C Fm (1 +k) following condition: 1/ 3 3 ∆C F = 105(K S / LWL ) −0.6410− 1 2 { (1+k) −C FS (1+k) −∆C F } Z m = FD = ρ m SmVm C FM −6 2 K S =150×10 m The self-propulsion points can be calculated 2 when Vm, nm, Tm, Qm, Zm have been recorded. C Tm= Rm / 0.5ρ mVm Sm For the power prediction of the full-scale ship with the above mentioned energy saving de- and the friction coefficient is the ITTC 1957 vices, the self-propulsion test of the ship model line. is usually carried out with and without the energy saving device in order to compare their When the Froude method is used influence on the propulsion performance. CTS =C FS +C R +∆C F + The ship model speed Vm is first corrected C R =CTm −C Fm for blockage (a procedure following the Emerson method is used – see the Resistance Comwhere,∆C F + is the resistance correction coefmittee report of the 19th ITTC, 1990). The testing procedure of the ship model with the ficient. It can be adjusted to keep CTS cal cul ated from both methods nearly the same. Normally energy saving device is the same as that without the energy saving device. The wetted it can be taken from the following empirical surface area of the energy saving device is not formulae: taken into account in the transformation of the results to the full scale. 10 3∆C F + =0.75−0.00352LWL

[

]

[

]

The basic transformation. The following or formulae can be derived from the kinematical and dynamical similarities 3.5 PE =(RTm −FD )Vmλ ρ S / ρ m VS = λ

1/ 2

Vm

24 where, FD is the resistance correction of the rical stern and twin-skeg forms, the ship’s preself-propulsion point. It can be estimated from rotating efficiency η n will lead to a propulsion the following formulae: efficiency in the form FD =0.5ρ mVm Sm[(1 +k)C Fm −(1 +k)C FS −∆C F ] 2

FD + =0.5 ρ mVm Sm [C Fm −C FS −∆C F + ] 2

η

n

= n0 / n

where, n is the revolution rate of the propeller behind the asymmetrical stern and n0 is the corDetermination of the propulsion coefficient.responding nominal revolution rate. This can The coefficients KTm and KQm are obtained from be obtained by taking the average value between the revolution rates of right-hand and the self-propulsion test as follows left-hand turning propeller models in the selfpropulsion tests. Or it can be the revolution rate 2 4 K Tm =Tm / ρ mnm Dm of this propeller model in the open-water test 2 5 K Qm =Qm / ρ mnm Dm with VA as its axial inflow speed when its thrust reaches the value in the following equation

( (

) )

Based on the thrust identity method J0m, KQ0m RVS TVA n0 Q0 (1 − t ) and η 0m can be read from the propeller model = =η 0η nη Rη H ηD = − nQ n Q n Q w ( ) 2 2 1 π π characteristics whose scale effect correction 0 0 has been made with KTm as the input data. Fol- The prediction of the full-scale performance for lowing this, the propulsion coefficients can single screw ships. The scale effect of the protherefore be calculated as follows pulsion factors recorded in the self-propulsion test revised at MARIC by the following formulae η Rm =K Q 0 m / K Qm 1−wm =J 0 mnmDm / Vm 1−tm =(Rm −FD )/ Tm

η Hm =(1−tm )/ (1−wm )

′

η Dm

=η 0 mη Hmη Rm

′ whereη Dm should be coincident with η Dm (in the above equation η Dm = PE / PDm ).

[CFS (1+k)+∆CF ] [CFm (1+k)]

wS =tm +(wm −tm )

tS =tm −0.08834+0.01262LWLm η RS

= ηRm +0.08645−0.01236LWLm

J0s, KQ0s and η0S can be read from the load coefficient of the full-scale propeller KTS/JS2 chart with the KTS/JS2 value as the input.

Basic transformation for twin-screw ships. SCTS With Tm and Qm values for both propeller mod- K TS / J S 2 = 2 2 2D (1−tS )(1−wS ) els as the input data, the power and the thrust deduction fraction can be calculated. For determining wmη, 0m, and η Rm the respective av- The following can then be calculated erage values for both propeller models should (1 −wS )VS be used. ns = J 0S D Basic transformation for ships with asym3 PDS = 2πρ S D 5nS K Q 0S / η RS metrical stern form or twin-skeg form. For ships with a big tangential component in the η HS =(1−tS )/ (1−wS ) propeller inflow, such as those with asymmet- η DS = PE / PDS or

25 Since the presence of a rudder would weaken the swirl in the propeller slipstream, it is essential to place the rudder behind the propeller The additional increment of the bilge keel in these tests. By using this configuration, the measurements of thrust and torque are taken and the air resistance of the superstructure over the water level can be estimated as 4-6% of PDS..for the propeller with and without the PBCF. The rate nS is proportional to (1.04-1.06)1/3. TheBased on the measured values of the torque, it values of CPand CN obtained from this method is possible to estimate the delivered power and are close to 1, so PD, ns and VS can be used di- hence the ratio of the delivered power with and rectly to predict the full-scale performance. without the PBCF. The measurement of this ratio will be the gain (improvement rate) in For the prediction of the full-scale perfor- propeller efficiency (∆ηpm) due to the PBCF in the model scale. mance for twin-screw ships with shaft brackets η DS

=η 0Sη RSη HS

In the POT set up one should bear in mind that the presence of the boat housing and the ts =tm rudder behind the propeller will affect the proη RS =η Rm peller advance velocity. Therefore the necessary correction should be made in the advance coefficient in terms of the wake fraction caused For twin-skeg ships wm should be correctedby the boat housing and rudder, and comparibased on experience. The full-scale performan-son of the efficiencies should be made at the ce prediction method and its procedure are thecorrected advance coefficient for the same thrust. same as those for single-screw ships. wS =wm

By considering the scale effects mentioned earlier in Section 4.1, the propeller efficiency 5.6 Propeller Boss Cap Fins (PBCF) gain in the model scale (∆ηpm) is related to the improvement rate of the propulsion efficiency Although no particular extrapolation method specific to PBCF has been reported in on the full-scale (∆ηs) as shown, for example, the open literature, it appears that the PBCF by Ouchi & Tamashima (1989), in figure 5. patent holding companies have a well estab- This guidance implies that the power saving or lished procedure to quantify the efficiency gainefficiency gain expected at the full-scale will be 2 to 3 times greater than the model scale by the use of PBCF. This procedure is based upon the so-called “reverse POT” (Propeller predictions. Open Tests) set-up and experience gained over In the above outlined procedure the proa considerable number of full-scale measurepeller and PBCF are considered as a unit proments. Furthermore, private communication by this committee with the West Japan Fluid En- pulsor and the effect of the PBCF is included in gineering Laboratory Co. Ltd has confirmed the propeller open water characteristics. The that the extrapolation for ship powering per- prediction of the efficiency gain due to PBCF formance is based upon the ITTC 1978 proce- does not require self-propulsion tests, but relies dure, Nishimoto (1998). These are briefly re- heavily on the accumulated knowledge and experience gained with PBCF on the full-scale. viewed in the following. In the reverse POT configuration the propeller is placed downstream of the open water propeller boat in order to provide a free flow field for the development of the hub vortex.

26 9 Container

8

s

) 7 (% t s e T6 l a ri T 5 ip h S y b 4 in a G

Ferry

Pcc Container

Container Pcc

2

3

Reefer

Pcc Pcc

2 Container

1 0 0

1

2

3

Gain by Model Test (%) ∆ =η=pm

Figure 5. Correlation of Efficiency Gains by PBCF between Model and Actual Ship, Ouchi & Tamashima (1989)

correction to the propeller characteristics due to difference in blade drag at model and full scale is developed for conventional propellers. Moreover, it relies on data for one representative profile only: at 0.75R. Hence due to differences in geometry, velocity and load distributions this method of scaling is not believed to be accurate enough for tip fin propellers. Instead the blade was subdivided into streamwise strips and the sectional drag estimated using the theoretically calculated velocities and a simple, flat-plate estimate of the frictional resistance which was dependent on mainly the local Reynolds number. To secure a fair comparison this procedure was applied to both the tip fin and the conventional comparator propellers. By this procedure the corrections turned out to be bigger for the tip fin propeller, i.e. it is more sensitive to scale effects. Unfortunately, no full scale tests have been done, so no confirmation of this scaling procedure can be made. 5.8 Extrapolation From an Extended Analysis of Self Propulsion Test Data

An alternative to the above procedure, Under this category are grouped some which makes use of self-propulsion tests, has been used by the West Japan Fluid Engineeringmethods that were developed with the aim of improving the powering prediction process by Laboratory Co Ltd. and is also reported implicitly by Ouchi (1988). In this procedure the avoiding and replacing some of the assumptions inherent in their use. The common theme PBCF is considered to be an appendage and then the usual open water characteristics of thebehind these methods is that the use of the propeller are predicted from standard series towing resistance is considered meaningless or data. The effect of the PBCF is included in the misleading when attempting to predict full self-propulsion factors, which are obtained by scale power. means of the thrust identity method. The perAbout ten years ago, some researchers formance of the ship with PBCF is predicted based on the standard ITTC 1978 procedure. (Holtrop 1990), started to make full scale powering predictions using self propulsion and open water tests as the sole source of experimental data. This was the first step in the di5.7 ITTC 1978 Modified Method for Tip rection of dispensing with towing resistance Fins data, but this method was still consistent with The only extrapolation method reported in traditional propulsion analysis since the model detail in the literature for end plate and similarresistance had to be reconstructed from model propellers is that reported by Andersen (1996).self propulsion test results. Since the tip fin propeller is only a slight modiBriefly described here are two methods fication of a conventional propeller the ITTC 1978 procedure has been followed with only a which fall into this category, but which have minor modification. The standard ITTC 1978 not yet been developed to a stage suitable for

27 general application. These methods, however, The second method described in this sechave promise in the development of useful tion, Schmiechen’s (1991) “rational theory”, is tools for the full scale powering prediction for a method for ship powering analysis based on vessels equipped with unconventional propul- results from model self-propulsion load varysors. ing tests alone. Stand alone resistance and propulsor open water tests are avoided. Iannone (1997a, 1997b), starting from the ITTC 1978 methodology framework, introTwo overload tests are done at the same duced a new propulsive analysis through the steady speed, but different values of the overseparation of the self propulsion flow into its load. Care must be taken to ensure that the viscous and potential components. To this end speed is steady to avoid significant acceleraonly open water and British self propulsion tion/inertia forces in the longitudinal momentests are required. In particular, to evaluate the tum equation. The following parameters are viscous component of the self propulsion flow measured for the two tests: speed, shaft rotathe innovative concept of self propulsion formtional speed, shaft thrust, torque and the towing factor KSP is introduced. force. These are designated In the first version of this methodology the V1 N1 T1 Q1 F1 self propulsion form factor was derived from V2 N2 T2 Q2 F2 model thrust measurements at low speed during self propulsion tests. This procedure, however, produced rather uncertain values of KSP, due to The thrust, T, and torque QP, are assumed the small values of measured thrust and the un-to vary as a quadratic function of the rotational avoidable presence of laminar flow or separa- speed N such that tion. 2 T =TO N +TH NV

2 Iannone’s investigations (1997b) revealed Qp =QP 0 N +QPH NV that a sort of thrust deduction phenomena occurs during self propulsion tests. Furthermore it which is equivalent to them being linear funcwas appreciated that the self propulsion form tions of the ship advance coefficient JH. The factor at service conditions depends not only subscripted terms on the RHS of these equaon hull trim at speed, but also on speed itself tions are found from the experimentally meas(Iannone 1998). As a consequence the predicured values as follows tion method was refined (Iannone 1998) by determining the thrust deduction factor and the T1 N2V2 −T2 N1V1 self propulsion form factor by tests at the serviT0 = 2 2 N1 N2V2 −N2 N1V1 ce speed. Moreover a new data reduction 2 2 method was proposed (Iannone 1998), leading N1 T2 −N2 T1 TH = 2 to different relationships among propulsive 2 N1 N2V2 −N2 N1V1 characteristics measured during British self Q N V −Q2 N1V1 propulsion tests. QP 0 = 21 2 2 2 N1 N2V2 −N2 N1V1 After some satisfactory applications to 2 2 N1 Q2 −N2 Q1 single and twin screw hulls, Iannone claims QPH = 2 2 that the method, though still in a refinement N1 N2V2 −N2 N1V1 phase, is promising for its suitability to be used as an extrapolation tool for full scale powering From the steady state form of the longitudinal predictions for unconventional propulsors. momentum equation

28 0 =TE +F −R =T (1−t) +F −R

K PL =K PLPO +K PLPI J p +K PLP2

Jp

2

2 and by assuming that the thrust deduction fraction can be modelled as a linear function of the The coefficients of this parabola are found in ship advance coefficient the following sequence from the coefficients already calculated:

t =tH J H

K PPO =2π KQPO

the value of tH is found as tH =

T2 +F2 −T1 −F1

K PPH =2π KQPH

ö

æ T2 V2 T1 −V1 çç DN − DN è 2 1

1/ 2

K PLPO

æ2 ö 3/ 2 =K PPO −ç ÷ KTO èπ 1/ 2

where D is the propeller diameter. The resistance is found from

æ tH V1 ö ç +F1 R =T1 ç1 − è DN1

K PLP1 =K PPH

K æ2 ö 3 1/ 2 −ç KTO KTH − T 0 2 2 èπ

To obtain KPLP2 use is made of the zero thrust condition when the equation

K T = K T 0 +K TH J HT =0 and the advance coefficients of the two steady states are found as yields in sequence J H1 =

V1

; J H2 =

V2

DN1 DN2 The following coefficients are then found based on Schmiechen’s definitions KT 0 =

T0 ρD

K QP 0 =

K TH =

; 5

K QPH =

QP 0

ρD

TH

;

4

ρD

3

QPH ρD

4

J HT =

KTO KT H

K PLT =K PPO +K PP H J HT and a cubic equation for the propeller advance coefficient in the zero thrust condition J PT =

K PPH

−

2 PPLP1

+

4 ( K PLT −K PLPO ) 2

KTH π J PT π J PT Schmiechen then assumes that the power coefficient representing the power losses of the propeller, the difference between the shaft As JPT occurs on both sides of this equation, a power supplied to the propeller and the jet solution here is to assume a value for JPT on the power of the propeller in coefficient form RHS and calculate a new value for JPT on the LHS. This procedure can be iterated until the K PL = K PP −K PJ RHS and LHS values fall within a pre-set limit. From this the coefficient KPLP2 can be found can be represented by a quadratic equation in from the advance coefficient of the propeller, JP.

29 K PLP 2 =

2

( K PLT −K PLPO −K PLP1J PT )

2

J PT

various parameters in Schmiechen’s “ thrust deduction theorem”

Using the above basis, the propulsive perfor1/ 2 τ =(1+CT ) −1 mance of the vessel can be found over a range of ship advance coefficients, JH, which can be assumed. In sequence the thrust and power co-thrust deduction fraction t = tH JH efficients are found as 2 é τt ù ê(1 +τ )t − KT = KTO +KTH J H 2 ë χ = K PP =K PPO +K PPH J H (1 −t) An iteration is then necessary to obtain the energy wake working value of the propeller advance coefficient at this ship advance coefficient and this is we χ (1 −w) = − 1 done using the loss parabola. w w K PL =K PLPO +K PLP1J P + K PJ =K PP −K PL JP =

K PJ KT

−

2 KT

π

K PLP2 J P 2

2

and the various propulsive efficiencies: propeller efficiency

2

KP J

η TP

KT J P

=

K PP

In this procedure also, a value of Jp is assumed jet efficiency and a new value is calculated; the process is iterated until the desired accuracy is achieved. From here the following propulsive coefficients are obtained K QP = K QL =

K pp 2π K PL 2π

wake fraction w = 1−

JP JH

8 KT π

JP

2 = 1 1 +(1 +CT ) 2

pump efficiency η JP

=

η TP η TJ

effective thrust coefficient

thrust loading coefficient CP =

η TJ

2

CE =

K T (1 −t) æ JH

2

hull efficiency η

ET

=

(1 − t) (1 −w)

T

E çç =ρ D 2V 2 è

ö

30 total propulsive efficiency

η EP =η ETη TP configuration efficiency

η = EP η JP

azimuthing unit and the propeller is considered as a whole and the propulsive coefficients are related to the interaction between the whole propulsion unit and the hull. In this case the system’s net thrust must include the drag forces on the housing. The resistance test with the hull is performed without the propulsion unit fitted.

æ PE ö çç =P è J

In the second approach, the azimuthing unit is considered as an open propeller in isolation and the propulsive coefficients are related to energy wake the interaction between the propeller and the hull appended with the housing. Therefore the propeller thrust and torque will not include the we = w – χ (1-w) effect of the drag forces on the housing and the resistance test must be performed with the housing unit fitted, but without the propeller in place. and normalized pressure level in the wake 2 2 In both of the above approaches the probC p =(1 −w) ( χ +2χ ) . lem arises as to how to scale the drag of the housing unit due to the low Reynolds number Although Schmiechen has not done a whole at model scale which results in relatively large set of evaluations for a series of ships, he has drag forces. In the first approach, a scale effect presented comparisons between model and full correction can be applied to the model KT valscale parameters for detailed tests done on the ues obtained from an open water test with the Meteor. For this vessel, the following paramecomplete unit where the system’s net thrust is ters were found to be similar for the model and measured. In the second approach, this correcfull scale: KT, KQL, ηJP, CE, ηEP and ηEJ. tion can be applied to the appended hull resisSmall variations between model and full scale tance based on the resistance results. The final occurred in: KQ, w, ηTJ, ηTP , CP, ηET, CE, ηEJ power prediction is expected to be the same for and t. both cases. However, the total propulsive efficiency will be different due to the difference in the resistance for the naked and appended hull. 5.9 Z - Drives (Podded propulsors) This is actually a matter of definition of propulsive efficiency. In some cases a stock proAllied with the recent upsurge in the num- pulsor unit may be used instead of a geometriber of applications of podded drives, there is cally similar model of the actual propulsor. growing concern with the differences in testing procedures and extrapolation methods being In a third approach, open water tests of the used for these propulsors, particularly for trac-podded drive as a whole unit are also done. tor type units. The extrapolation methods in useHowever, the thrust and torque measurements have not been published in the open literature. are taken at the shaft excluding the drag effects Also there is a lack of full-scale data for these of the housing and are corrected to the fullsystems due their short application history. Thescale. The resistance tests with the model hull following outline approaches to the testing andare done without the podded drive while the extrapolation for podded propulsors have beenresistance of the drive housing and the scale suggested. effects are estimated numerically.

η EJ

In the first approach the housing of the In a fourth approach, in addition to resis-

31 tance and self-propulsion tests, two open water tests: one with the propeller alone; the other As a result of its work, the Committee feels with the propeller plus the housing together that where there is weak interaction between (whole unit), are used for the evaluation of the the unconventional propulsor and the ship hull, podded drive. In this approach, the podded then methods developed on the core of the drive is assumed to be an appendage and its ITTC 1978 approach can give levels of accuradrag is converted to the full-scale without any cy for extrapolation of full scale ship powering correction for scale effects. The thrust deduc- performance that are of the same magnitude as tion fraction is calculated from the system’s those obtained for conventional propulsors. (net) thrust and the total resistance with the Weak interaction occurs with devices such as pods. The wake fraction is obtained based on a tip fin and tip plated propellers, propeller boss cap fins, and other devices that are only a small KQ identity and no correction is made to the full scale. The open water test results for the whole modification to a conventional propeller. Efpodded drive are corrected for the difference infective modifications do need to be made to Reynold’s number between the open water testaccount for differences in blade frictional drag resulting from differences in Reynolds number and the self-propulsion test. Further corrections are also made to the open water test results for between model and full scale and a method to the full scale by using ∆KT and ∆KQ values do this for tip fin propellers is referred to in based on open water test results with the pro- section 5.7. However, where there is strong interaction between the device and the hull, then peller alone. methods based on the ITTC 1978 approach are not adequate. Strong interaction occurs with all types of ducted and partially ducted propellers, 6 DISCUSSION OF EXTRAPOLATION pre- and post-swirl devices, z-drives, etc. The METHODOLOGIES reason for this is that these latter unconvenIn section 4, this report has given an over- tional propulsors have large, complex and strongly modifying effects on the flow around view of methods of extrapolation used in the past for different types of unconventional pro- the hull. In addition, the exact physical mechapulsor. Problems arising during testing and ex- nisms by which some of these devices interact with the hull is not always clear. As a result, trapolation have been highlighted in some cases. In section 5, these methods of extrapola-the testing and analysis leading to extrapolation tion, together with other candidate extrapola- does not lend itself to being broken down into the pieces designated by the ITTC 1978 procetion methodologies, have been discussed in dure (i.e. separate resistance, open water profurther detail. pulsor, and self-propulsion tests and their analysis). Most of these interactions between To date, many extrapolations for unconventional propulsors have used methods heav- propulsor and hull are strongly Reynolds number dependant. ily based on the ITTC 1978 methodology. However, the modifications to the ITTC 1978 Benchmark work is necessary to investigate method proposed and used for many of these unconventional propulsors are different: i.e. these flow phenomena and to identify trends in there are as many or more modifications to the powering performance as the Reynolds number is increased. Examples of some of the types of ITTC 1978 method as there are types of unconventional propulsor. This situation is less test that are necessary are described in sections 4.6 and 4.7 for models fitted with ducts and than ideal as there is always a question as to whether any power saving predicted is a fun- partial ducts. To do this work it is necessary to damental characteristic of the device or a func- use the largest facilities available and to avoid tion of the testing, analysis and extrapolation strict adherence to Froude scale speeds. New test techniques need to be developed in some method in use.

32 cases. A heavy reliance needs to be made on the flow velocities would be made at these loself-propulsion tests with geometrically scaled cations and these would be used to identify the models of the unconventional propulsor and thrust of the device. (Note that this is analohull arrangements. Use can and should also be gous to the method of analysis proposed for made of RANS-type CFD calculations to inwaterjets by the Waterjets Group of the 21st vestigate trends in certain detailed flow beITTC.) The thrust deduction effect could be haviour as the Reynolds number of a particular found from an integration of local pressures set up is increased. over the afterbody. CFD methods would be used to supplement velocity values from actual To scale or extrapolate powering perform- test results and to investigate trends at higher ance, it is necessary to know the trends in per- Reynolds numbers. In contrast, the reason for formance as Reynolds number is increased andlooking at an extended analysis of selfthis can be found from these specially designedpropulsion test data is with the plan that powand perhaps expensive tests. Full scale trials ering performance information can be extracted results from ships fitted with unconventional from these tests in a more macro manner (than propulsors are needed for the same cases that by using the integrations referred to above), but have been studied in depth at model scale. On- by avoiding as far as possible the less realistic ce this base of knowledge is established, more aspects of the assumptions present in convenroutine work might be possible following a tional methods of extrapolation. The Commitmore simple test and extrapolation procedure. tee recommends that work is continued on the development of new general methods for exIn carrying out propulsion tests for uncon- trapolation of the powering performance of ventional propulsors, the Committee recom- ships fitted with unconventional propulsors mends the use of extensive load varying tests along these lines. done at high values of Reynolds number. In working towards a guideline for extrapolation, The published performance of unconventhe Committee recommends that the data from tional propulsor systems is sometimes clouded these tests be used in combination with an for a number of reasons. Often the advantage of analysis using one or more of: a momentum fitting an unconventional propulsor design is analysis, as described in outline in section 5.1; not actual energy saving, but the suppression of the results from RANS-type CFD calculations; cavitation or the reduction of unsteady pressure and/or an extended analysis of self-propulsion pulses and vibration. There may be no efficientests, some perhaps rather simplistic examplescy gain in some situations. In addition, unconof which are described in section 5.8. The aim ventional propulsors can be fitted to ships is to avoid assumptions made in the convenwhere the initial design of the conventional tional ITTC 1978 type analysis which are propeller is not optimum. This latter leads to an known to be less than realistic. An example of apparently large propulsive efficiency gain one of these is the use of a propeller open waterwith the device, whereas a proportion of that test (often done at one Reynolds number) with gain is the result of more focussed design effort results from the propeller in the behind condi- being placed on the propulsion design in gention (often done at a different Reynolds num- eral. ber) to identify a wake fraction. A second example is the empirical methods used for the scaling of that wake fraction from model to full 7 DRAFT CONCLUSIONS AND RECscale. OMMENDATIONS A true momentum analysis would identify General technical conclusions streamtubes passing from ahead of to behind the propulsor unit. Detailed measurements of

33

•

•

•

•

•

•

•

propulsors when applied to ships fitted with The exact physical mechanism by which some unconventional propulsors interact tip fin and tip plated propellers, propeller boss cap fins, surface piercing propellers with the hull is not clear. and other devices which are modifications of a conventional propeller. Extrapolation of full scale powering for ships fitted with unconventional propulsors Recommendations to the Conference may be developed from extensive load varying tests and in addition one or more of momentum analysis, CFD computations, or • For powering prediction it is recommended that ship models fitted with unconventional an extended analysis of self propulsion test propulsors, such as propellers with ducts, data. partial ducts, pre- and post- swirl devices, z-drives, etc., should be tested as a unit and Use of extrapolation methods for unconnot broken down into component tests of ventional propulsors based on modificathe hull, propulsor and rotor and stator tions to the ITTC 1978 method show simicomponents. lar levels of variation of the powering prediction found between methods as the level • Extrapolation methods of full scale powering for unconventional propulsors should of the power saving expected with the debe done using self-propulsion load varying vice. In other words, the accuracy of these tests of the geometrically similar ship extrapolation methods is in most cases of model and geometrically similar propulsor. the same order of magnitude as the level of power saving of the device under analysis. • It is recommended that for self-propulsion There is a shortage of accurate data from tests with unconventional propulsors the effect of the rudder on the propulsion system full scale trials supporting extrapolation be considered due to the influence of the predictions made for unconventional prodownstream swirl on the rudder. pulsors. Extrapolations cannot be reliably made of• To reduce/ eliminate the scaling of flow separation effects during self-propulsion self-propulsion test data if flow separation tests these tests should be done at higher which occurs on the unconventional prolevels of Reynolds number than can be pulsor and/or the ship hull is not scaled corachieved by rigid adherence to Froude rectly. It is recognised that most methods of ex- number scaling. trapolation currently in use are modifications to a greater or lesser degree of the Recommendations for future work ITTC 1978 method. It is recommended that • It is recommended that work is continued the ITTC 1978 method is only used with on the development of new general caution as a guideline for extrapolation of methods for the extrapolation of model test model test results to full scale ship powerresults for unconventional propulsors. The ing prediction for unconventional propulextrapolation methods should be validated sors. against extensive model tests and full scale Extrapolation methods for unconventional trials. propulsors based on modifications of the • Extensive tests for unconventional propulITTC 1978 method are expected to give sors over a wide range of Reynolds numpowering predictions to the same level of bers have shown that some usual assumpaccuracy as that achieved for conventional

34 tions of trends in performance are false. peller with high aspect ratio blades, Marine There is a need for detailed tests over a Technology, Vol. 26, No. 3, 192-201. wide range of Reynolds numbers to be done on ship models fitted with each different Bruce, E. P., Gearhart, W.S., Ross, J.R., and Treaster, A.L., 1974, The Design of type of unconventional propulsor. Where Pumpjets for Hydrodynamic Propulsion, possible such test results should be comProceedings of Fluid Mechanics, Acoustics, pared with CFD computations and full and Design of Turbo-Machinery, NASA, scale trials. Washington, D.C. REFERENCES

Ferrando, M. & Scamardella, A. 1996 Surface Piercing Propellers: Testing Methodologies, Andersen, P., 1996, A Comparative Study of Result Analysis and Comments on Open Conventional and Tip-Fin Propeller PerWater Characteristics. In Proceedings Small formance. In Proc. 21st Symp. on Naval Craft Marine Engineering Resistance & Hydrodynamics. Trondheim, Norway. Propulsion Symposium, pp.5-1 – 5-27. YpAnon., 1998, Whale tail hits the road, Marin silanti: University of Michigan News, No. 65. Friesch, J. & Johannsen, C. 1994 Propulsion Optimization Tests at High Reynolds Numbers, SNAME Transactions, Vol. 102, pp.1 – 21

Arabshahi, A., Beddhu, M., Briley, W., Chen, J., Gaither, A., Gaither, K., Janus, J., Jiang, M., Marcum, D., McGinley, J., Pankajakshan, R., Remotigue, M., Sheng, C., Gabriel, R. and Atlar, M., 1998, Calculation of Sreenivas, K., Taylor, L., Whitfield, D., the performance of a ship propeller with 1998, A Perspective on Naval Hydrodycyclic blade pitch control, International namics Flow Simulations, Twenty-Second Shipbuilding Progress, 45 (443), 201-223. Symposium on Naval Hydrodynamics, Washington, D.C., August. Gearhart, W. S., McBride, M. W., 1989, “Performance Assessment of Propeller Boss Cap Fin Type Device”, 22nd ATTC, St Atlar, M., Hannah, J., Takinaci, A.C., and John’s., Newfoundland. Korkut, E., 1998, “Effect of various boss caps on the efficiency and cavitation per- Hadler J.B. & Hecker R. 1968 Performance of partially submerged propellers. In Proformance of a propeller.”, Intl Symposium ceedings, Seventh Symposium on Naval Honoring Tarik Sabuncu On the Occasion th Hydrodynamics (ed. R.D.Cooper & S.W.. of His 75 Birthday, Istanbul Technical Doroff), pp. 1449-1493. Arlington: Office University. of Naval Research – Department of the Navy. Atlar, M., Patience, G., 1998, “An Investigation into Effective Boss Cap Designs to Eliminate Hub Vortex Cavitation”, Halstensen, O. and Leivdal, PA., 1990. The development of the speedZ propulsion PRADS’98, The Hague. system. The 7th high speed surface craft Bennett, J., 1996, Of fiddles and fins, Wooden conf., London. Boat, 132, September/October, 40-43. Henderson, R.E., McMahon, J.F. and Wislicenus, G.F., 1967, A Method to Design Bose, N. and Lai, P.S.K., 1989, The experimental performance of a trochoidal pro-

35 Pumpjets, Underwater Missile Propulsion, Compass Publications. Kerwin, J. E., Keenan, D.P., Black, S.D. and Diggs, J.G., 1994, A Coupled VisHollstein, H.J., Gomez, G.P. & Gonzalezcous/Potential Flow Design Method for Adalid, J. 1997 Bulk carrier speed trialWake-Adopted Multi-Stage, Ducted Proswith a Sistemar CLT propeller. The Naval pulsors using Generalized Geometry, Architect, February 1997, pp. 33-6. Transactions of the Society of Naval Architects and Marine Engineers, New York, Holtrop J., 1990, Are Resistance Tests Indispp. 2-28. pensable?, Proceedings 20th ITTC, Madrid, Kurimo, R. et al.,1998. Azipod propulsion for Vol. 2, pp. 454-459. passenger cruisers: Details of the hydrodynamic development and experience on the Iannone L., 1997a, Power Performance Analypropeller design for “Fantasy”-class cruise sis and Full-Scale Predictions without liners. NAV & HMSV Intl. Conference, Towing Tests, Proceedings of Nav & Sorrento. HSMV 1997, Sorrento, Italy, 18 – 21 March. Iannone L., 1997b, Self-Propulsion Flow Com- Kurimo, R., 1998. Sea trials experience of the first passenger cruiser with podded propulponents Analysis for Power Performance sors. PRADS’98, The Hague. Predictions, Proceedings Eight Congress of the International Maritime Association of Mediterranean (IMAM), Istanbul, Turkey. Lai, P.S.K. 1990 Oscillating foil propulsion, Ph.D. thesis, University of Glasgow. Iannone L., 1998, Fattore di Forma e Risucchio Larsson, L., Regnstrom, B., Broberg, L., Li, Dnelle Condizioni di Servizio dall’Analisi A, and Janson, C-E, 1998, Failures, Fantadel Flusso di Autopropulsione, INSEAN sies, and Feats in the Theoretical/Numerical (Rome model Basin) Report No. 1996-61, Prediction of Ship Performance, TwentyRome, Italy. Second Symposium on Naval Hydrodynamics, Washington, D.C. August. Isshiki, N., Hashimoto, K. and Morokawa, H. 1987 Studies on fin ship propelled by oscillating tail fin, Proceedings International Manen, J.D. van, 1973, Non-conventional proMarine Engineering Conference, Chinese pulsion devices, International Shipbuilding Society of Naval Architecture and Marine Progress, Vol. 20, No. 226, 173-193. Engineering, Shanghai. Mewis, F., 1998. Podded drives im vormarsch de Jong, K., Sparenberg, J.A. Falcão de Camhydrodynamische aspekte. Schiff & Hafen, pos, J.A.C. & van Gent, W. 1992 Model 11. Testing of an Optimally Designed Propeller with Two-Sided Shifted End Plates on the Blades. In Proc. 19th Symp. on Naval Hy- Minsaas K.J. 1998. Propulsion systems for high speed marine craft. MARINTEK A/S drodynamics, pp. 461-75. Washington DC: Report, May. National Academy Press. Nishimoto, H., 1998, Private Communications Karafiath, G, 1997, U. S. Navy Ship-Model with West Japan Fluid Engineering LaboPowering Correlation and Propeller RPM ratory Co., Ltd. Predictions, Propellers/Shafting ’97, Society of Naval Architects and Marine EngiOgura, M., Koizuka, H., Takeshita, H., Kohno neers.

36 Y., Ouchi, K., Shiotsu, T., 1988, “A Screw Propeller Boss Cap with Fins”, UK Patent Application , GB 2194295 A. Olofsson, N. 1996 Force and Flow Characteristics of a Partially Submerged Propeller. Doctoral Thesis. Goteborg: Chalmers University of Technology Department of Naval Architecture and Ocean Engineering. Ouchi, K. 1988, “Research and Development of PBCF (Propeller Boss Cap Fins) to Enhance Propeller Effciency”, The Motor Ship 10th. International Marine Propulsion Conference, London

Propeller Efficiency -”, Naval Architecture and Ocean Engineering, Vol. 28, Papers Edited by Society of Naval Architects of Japan. Published by Ship and Ocean Foundation, Japan 1992. Ouchi, K., Tamashima, M., Arai, K., 1992 “Propeller Noise Reduction caused by PBCF”, PRADS’92, Newcastle upon Tyne. Potsdam Model Basin (SVA), Propeller Design, 1995, “TVV and HVV Propellers - an Innovative Concept developed by SVA in Cooperation with SCHOTTEL”, Postdam Model Basin (SVA) Product Brief.

Qian, W.H. et al. 1992a The exploratory deOuchi, K., Ogura, M., Kono, Y., Orito, H., velopment of simplified compensate nozzle Shiotsu, T., Tamashima, M., Koizuka, H., and its composite device. Proceedings of 1988, “ A Research and Development of the 2nd International Symposium on ProPBCF (Propeller Boss Cap Fins) - Impeller and Cavitation, Hangzhou, China provement of Flow from Propeller Boss -”, Journal of Society of Naval Architects of Qian, W.H. et al. 1992b A new type of ship Japan, Vol. 163. energy-saving device—fore-propeller hydrodynamic fin sector, Journal of Kansai Ouchi, K., 1989, “Research and Development Society of Naval Architects of Japan, No. of PBCF (Propeller Boss Cap Fins) – Im218. provement of Flow from Propeller Boss –“, The Proceedings of the International Sym- Rains et al. 1981. Hydrodynamics of podded posium on Ship Resistance and Powering ship propulsion. Journal of Hydronautics, Performance’89 (ISRP), Shanghai. Vol. 15, Nos 1-4. Ouchi, K., Tamashima, M., 1989, “Research Riijarvi, T., Li, J., Veitch, B., Bose., N., 1994, and Development of PBCF – New and Experimental performance and comparison Practical Device to Enhance Propeller Effiof performance prediction methods for a ciency -”, PRADS’89. trochoidal propeller model, International Shipbuilding Progress, Vol. 41, No. 426, Ouchi, K., Tamashima, M., Kawasaki, T., Koi113-136. zuka, H., 1989, “A Research and Development of PBCF (Propeller Boss Cap Fins) – Savitsky D. 1964 Hydrodynamic Design of 2nd Report: Study on Propeller Slipstream Planing Hulls, Marine Technology, Vol. 1 and Actual Ship Performance-”, Journal of No. 1, Oct. 1964 Society of Naval Architects of Japan, Vol 165. Schmiechen, M., 1991, Proceedings of the 2nd International Worshop on the Rational Ouchi, K., Tamashima, M., Kawasaki, T., KoiTheory of Ship Hull-Propeller Interaction zuka, H., 1990, “ Research and Developand its Applications, VWS, The Berlin Ship ment of PBCF (Propeller Boss Cap Fins) – Model Basin, HEFT 56, June 13-14. Novel Energy-Saving Device to Enhance

37 Schneekluth, H. 1986 Wake equalizing duct, The Naval Architect, April, 1986 / A pamphlet from Schneekluth Hydrodynamik, Entwicklungs- und Vertriebsges. MBH Schott, C.G., 1996, Design and Analysis of Axial Flow Turbo-Machinery Blades in Steady Incompressible Flow by a Combination of Momentum and Singularity Methods, Master of Science Thesis, The Pennsylvania State University. SCHOTTEL, 1995, “TVV and HVV Propeller, A Joint Development of SCHOTTEL and SVA”, SCHOTTEL-Werft Product Brief.

Van, S-H, Kim, M-C. and Lee, J-T, 1993 “Some remarks on the powering performance prediction method for a ship equipped with a pre-swirl stator-propeller system”, Korea Research Institute of Ship and Ocean Engineering, Daejeon, Korea, ITTC 1993, Vol. 2. Wislicenus, G.F., 1968, Pumping Machinery for Marine Propulsion, ASME Symposium on Pumping Machinery for Marine Propulsion.

Wislicenus, G.F., 1960, Hydrodynamic and Propulsion of Submerged Bodies, J. Am. Rocket Soc., Vol. 30, No. 12, pp. 1140 – Schulze, R., 1995, “SVA- Nabenkappenflossen 1148. fur Schiffspropeller”, SVA (Postdam Model Basin) Report No: 2218. Shiba H. 1953 Air-Drawing of Marine Propellers. Transportation Technical Research Institute, Report No. 9. Tokyo: the UnyuGijutsu Kenkyujo.

Yamaguchi, H. and Bose, N. 1994 Oscillating foils for marine propulsion, Proceedings of the Fourth International Offshore and Polar Engineering Conference, Osaka.

Zhao, H.H. et al. 1988 Experimental research Skidmore, J.E., Lueschen, J.D., Renzo, J.A. on the scale effect of ship geosim, Proand Landrum, C. 1989 Design of a human ceedings of Ship Performance Symposium powered wet submersible for competition, of China Shipbuilding Engineering Society. Oceans ‘89, Seattle. Sparenberg, J.A. & de Vries, J. 1987. An Optimum Screw Propeller with Endplates. In- Zhou, Z. M. et al. 1990 Energy-saving shaft bracket suitable for multi-screw vessels, ternational Shipbuilding Progress, vol. 34, China Shipbuilding, No. 2, (in Chinese). no. 395, pp. 124-33. Zierke, W.C. (editor) and 18 authors, 1997, A Stern, F., Paterson, E. G., Tahara, Y., 1996. Physics-Based Means of Computing the CFDSHIP-IOWA: CFD Method for SurFlow Around a Maneuvering Underwater face Ship Boundary Layers and Wakes and Vehicle, Technical Report No. TR 97-002, Wave Fields, IIHR Report #381. Applied Research Laboratory, The PennStierman, E.J. 1984 “Extrapolation methods for sylvania State University. ships with a ducted propeller”, International Shipbuilding Progress, Vol.31, No. 356. Treaster, A. L., 1969, “Computerized Applica- . tion of the Streamline Curvature Method to the Indirect, Axisymmetric Turbomachine Problem,” ARL TM 514.2491 -16, The Pennsylvania State University.

1

on pressure fluctuation on the oscillating foil surface has been published. Such data would be invaluable for calibration for CFD work. There is also a need for some research and development work on control mechanisms for full scale oscillating propulsors, but this is by Pengfei Liu, Institute for Marine Dynamics outside of the scope of the ITTC. This report demonstrates an excellent job Building on the success of the workshop by the Committee on such a difficult and comprehensive subject in such a short time. It held by the Propulsion Committee to compare is a pity that this Specialist Committee will not RANS and panel methods to predict propeller performance I suggest a workshop on CFD continue. Continuation of the work by this Committee is important, because the methods modeling of propellers operating in ducts. Many issues could be addressed in such a of predicting, both numerically and experimentally, the performance capabilities ofworkshop. Both experimental and CFD predictions of unsteady pressure distribution on oscillating propulsors, tip fins on screw the inner surface of a duct are difficult to find propellers and others are still not yet fully developed. There is a lot of remaining work to in the literature, although some authors have presented the steady pressure distribution be done. inside the duct without experimental Benchmark data including A bigger database covering more aspects verification. of the performance of energy saving devices, measurements of pressure fluctuation on the duct’s surface would be very useful for and alternatives or modifications to the calibration and verification of panel methods conventional propulsors would be valuable. Experiment data for calibration and validation and RANS codes. A workshop would focus efforts in this field. of computer codes for unconventional propulsors is very scarce. For example, it is difficult to find experimental data on tip fin or end plate propellers. Such data need to include not only the propeller shaft thrust and torque by M.X. van Rijsbergen (MARIN) coefficients but the blade pressure distribution In the discussion on extrapolation of as well. As stated in the report, the oscillating propulsors, the report of the development of oscillating foil propulsion systems is in its infancy. Little experiment datacommittee concentrates on propulsor-hull

2 interaction. MARIN would however like to ask The figure above shows that the torque attention for scale effects on the oscillating coefficient decreases with 30 to 60% for propulsor itself. Reynolds numbers in the range from 2·10 to 7·10 , dependent on the blade angle motion and the condition. The torque coefficient of the MARIN has been thoroughly involved in the development of the first prototype Whale wheel is dependent on both the lift and drag coefficients of the blades. Tail Wheel (see e.g. van Manen et al.,1996). This is a cycloidal propulsor which can operate in a trochoidal blade motion. From open water tests at scale factor 2 and several rotation rates, The observed decrease in torque coefficient it appeared that mainly the torque coefficient isof the Whale Tail Wheel occurs in the same considerably sensitive to the Reynolds number region of Reynolds numbers where a (see the figure below). significant decrease in drag coefficient of streamlined foils has been reported (see e.g. Hoerner, 1965). The Reynolds number is defined as: V Rc

Re =

Trochoidal propellers can operate at the condition where at some point in the blade path, where V is the average resultant velocity the entrance velocity to the blade equals zero or relative to the blade, c is the chord length and ν practically zero, causing extremely low is the kinematic viscosity of water. The torque Reynolds numbers. Special precautions during testing and for extrapolation of results may coefficient is defined as: have to be taken. Little is known on this issue. ν

KQ =

Q

Due to the sometimes large angles of attack of the flow on the blade, in combination with low Reynolds numbers, leading edge separation where Q is the torque delivered by the two may occur, which is also likely to be prone to main shafts (corrected for friction), ρ is the density of water, n is the rotation rate, D is the scale effects. diameter of the wheel and L is the span of the References: blades. The torque coefficient is normalised relative to the value at the lowest Reynolds van Manen, J.D. and van Terwisga, T., “A new number for each condition and blade angle way of simulating Whale Tail Propulsion”, 21 motion (designated ‘bam A’ or ‘bam B’). Naval Hydrodynamics, Symposium on Reynolds effect on torque coefficient Trondheim, June 1996 Whale Tail Wheel 2

4

ρn D L

120%

bollard pull 'bam A' design condition 'bam A'

100%

design condition 'bam B'

Hoerner, S.F., “Fluid-dynamic drag”, Hoerner fluid dynamics, 1965

80%

KQ 60%

by Hassan Ghassemi, Institute for Marine Dynamics

40%

20%

0% 2.0E+05

3.0E+05

4.0E+05

5.0E+05

Re

6.0E+05

7.0E+05

Of the many different unconventional propulsors considered by the Committee, podded propulsors are of increasing practical

3 interest at the present time. Due to their better maneuverability, regular inflow to the propeller,(a) reduction of cavitation and vibration, compactness, different types of podded propulsion systems (pulling, pushing and shuttle) are becoming popular in current construction using electrical propulsion in place of conventional propulsion systems.

The fourth general technical conclusion is: “There is a shortage of accurate data from full scale trails supporting extrapolation predictions made for unconventional propulsors”. My question is: Was any such data available to the Committee; did the committee validate extrapolation procedures with full-scale data?

Some recent computational research on a pulling type of podded propulsion system at the The fourth recommendation to the Institute for Marine Dynamics (IMD) raised (b) conference is: “To reduce the scaling issues concerning the best approach to take in of flow separation effects, selfmaking reliable predictions. For example, the propulsion tests should be done at drag of the strut/pod may be affected by the higher levels of Reynolds number than downstream propeller wake and its modeling. can be achieved by adherence to Froude Thus it is important to find a reasonable vortex number scaling.” My comment is: In wake model to calculate the drag of strut/pod in this context turbulence stimulators the propeller slipstream using a panel method. should account for this. I would like to By carrying out experiments, Halstansen and have recommendations for the use of Leivdal showed that strut/pod drag is turbulence stimulators included in the compensated by recovery of rotational energy committee report. in the slipstream. What method does the Committee recommend for the strut/pod drag of the pulling type of podded propulsion in the vortex wake generated by the propeller? by Friedrick Mewis, HSVA

The hydrodynamics of different types of First I would like to give my thanks to the podded propulsion systems (pushing, pulling committee for it’s really good work. In chapter and shuttle) may change the overall 5.9 (page 240) are described very well the hydrodynamic performance of the vessel. Can possible methods for using open water test the Committee comment on these differences results for podded drives. and how they may influence the approach to making performance predictions? Some factors are mentioned which affect the measured values of propeller thrust, propeller torque and unit thrust (system thrust) • Ghassemi, H. and Allievi A. 1999: Fluid Analysis and Hydrodynamic Performance and additionally four different approaches to extrapolation methods are described. of Conventional and Podded Propulsion Systems, Oceanic Engineering International, Unfortunately, the estimation of the Vol. 3 No. 2. propeller thrust in the case of pulling pods is very difficult and the correct measurement is impossible. The reason for that fact is the gap between the propeller hub and the pod housing. by Jaakko V. Pylkkanen (VTT) The pressure in this gap influences to a substantial level the measured propeller thrust. First of all I would like to express my This substantial inner force in the gap is a appreciation to the Committee for its result of the high pressure behind the working competent report.

4 propeller, the pressure change due to the propulsor are becoming more common and stagnation point of the pod shaft coupled with experiments are being conducted in water the large gap area. The width of the gap, whichtunnels to measure the flowfield. can vary in different model tests, and in different towing tanks, influences the measured ! $"%" & thrust too. Having been involved to some extent The error can be up to 10% dependent on [Riijarvi et al. 1994] with testing and the propeller loading, the pod-geometry, the computational work on trochoidal propellers, width of the gap and other reasons. The we agree that Reynolds number scale effects measured unit thrust is not affected by this are very important in the extrapolation of open problem because the force in the gap is an innerwater and self propulsion data on these force of this unit. My proposal is to use the propellers. As the discussor explains, large measured unit-thrust as the basis for the angles of attack do occur on the blades of a estimation of the so called “small figures” like trochoidal propeller, both during each thrust deduction fraction, effective wake revolution and at low advance ratios of fraction, propulsor efficiency… in pod operation. Work has shown, in fact, a sort of propulsion only. critical advance ratio below which stall effectively reduces both the thrust and torque coefficients of the propeller. A multiple stream tube theory [Bose, 1987] using experimental lift/drag coefficients at different Reynolds numbers for the foil ! " # sections in a quasi steady manner can be used to predict the performance of a trochoidal First of all, we would like to thank Dr. Liu propeller in a way that models this behaviour for his very supportive comments on the Committee’s efforts and his discussion. The and this critical advance coefficient, at least in most difficult problem for the Committee was a qualitative way. A more accurate method to identify reliable model/full-scale data for anywould require experimental characteristics for unconventional propulsor. In almost every casethe foil sections during dynamic stall from oscillating foil tests for the appropriate these databases are incomplete for the calibration and validation of computer codes; conditions and over a range of Reynolds however, it is very important to note that only numbers. model-scale data exists for most References: unconventional propulsors.We could not identify any experimental database that Riijarvi, T., Li, J., Veitch, B.J., Bose, N., 1994, includes details of the local steady/unsteady pressure and velocity fields associated with the Experimental performance and comparison of unconventional propulsor. In most cases the performance prediction methods for a trochoidal propeller model, International unconventional propulsor was designed to improve powering thus only nominal powering Shipbuilding Progress, Vol. 41, No. 426, pp. measurements were made. At this time the 113-136. Committee could not identify a specific reference that presents unsteady duct pressureBose, N., 1987, Rotary foil propellers, Papers of the Ship Research Institute, Tokyo, Japan, measurement data. Vol. 24, No. 5, pp. 45-67. We agree that a workshop on CFD modeling of propellers operating in ducts is needed. These types of unconventional

5 As far as the performance prediction of different types of pods is concerned: whether it The committee would like to thank Dr.is a pusher or a tractor type, both Ghassemi for his questions. configurations are unconventional, the latter being less conventional and hence more Although the methods for predicting the complex to handle. drag of the pod housing is not the main concern However, if the performance prediction is of this committee, Dr. Ghassemi is asking for the committee’s opinion on this issue. He is based on model test procedures, it would not particularly interested in the pulling (or tractor)make much difference to look for fundamental differences in the approaches depending on the type in which case the pod housing will be in type of the propulsor. This is because the the wake of complex propeller flow. current procedures, which are not in the public domain although they are outlined in Section As reported in Section 4.8 of this committee’s report, studies due to Rains et al. 5.9 based on the best of our knowledge, utilize standard tests of the resistance open water and (1981) and Minsaas (1988) include some self-propulsion. In any case, these tests will be information on the drag characteristics of performed for both types in a similar fashion, pusher and tractor type propulsors. In the method for the tractor type, Minsaas neglects although the scale effect correction for the pod the effect of cavitation and propeller induced drag (due to low Rn) will be more complex for drag, and provides a semi-empirical formulae the tractor type. for the viscous drag of the housing components depending upon whether the component is in or However, our recommendation to the outside the propeller’s slipstream. ITTC as outlined in Section 7, is that the On the other hand, amongst the reported extrapolation of podded propulsors should be computational methods, one can mention done based on a self-propulsion test which studies e.g. Dumez (1997) (reported in Le includes the hull, pod and propeller as unit and magazine du Bassin d’Essais des Carenes) and not by breaking the test up into components Korpus et al (1998) (reported in PRADS `98). such as open water, resistance etc. The former of these studies is based on the panel technique while the later utilized a References: RANS technique. There is also a recent useful work reported by Vartdal et al (1999) Dumez, F-X, 1997. PODS, some encouraging (published in the CFD `99 workshop, Norway) conclusions, Le magazine du Bassin d’Essais comparing these two procedures on a tractor des Carenes, No. 7, January. type pod drive. Korpus, R. et al., 1998. Hydrodynamic design of integrated propulsor/stern concepts by In the light of this information, the Reynolds – Averaged Navier – Stokes committee feels that panel methods are adequate to model the potential flow effects techniques. PRADS ’98, The Hague taking into account the influence of the propeller's vortex wake. However, the major Vartdal, L., et al, 1999. On the use of CFD contribution due to the viscous effects still Methods in Developing the Azipull Podded remains as the problem which will require Propulsion System, CFD Conference ’99, complex boundary layer and even separated Ulsteinrik, Norway flow models. For these, the committee recommends use of RANS techniques, as Minsaas, K.J., 1988. Propulsion Systems for successfully applied by Korpus et al (1998), High Speed Marine Craft, MARINTEK A/S including the effect of the propeller. Report, May issue. !

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The following briefly summaries the verbal reply given at the conference. (a)

Where full scale data was available, estimates of energy efficiency gains were made using this data and the results presented in the report reflect this. An example of this is included in the efficiencies quoted in tables 2.4.

(b)

The effectiveness of turbulence stimulators in controlling separation is extremely uncertain and cannot take the place of tests done at increased levels of the Reynolds number. !

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The following briefly summarises the verbal reply given at the conference. We thank the discusser for his comments and we agree that the gap between propeller and pod in tractor units has a great influence on the measured thrust values of the propeller. Again we recommend self propulsion tests to be used on their own as the basis of extrapolation of full scale power.The intermediate values in this process, such as thrust deduction fraction, wake, etc., may not have the same meaning, or values, as in the powering prediction for conventional propulsion systems.