1966-Vanmanen-research on the Manoeuvrability and Propulsion of Very Large Tankers

8/12/2019 1966-Vanmanen-research on the Manoeuvrability and Propulsion of Very Large Tankers

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^tto Sixth Symposium

& NAVAL HYDRODYNAMICS

PHYSICS OF FLUIDS, MANEUVERABILITY AND OCEANPLATFORMS, OCEAN WAVES, AND SHIP-GENERATED

WAVES AND WAVE RESISTANCE

Sponsored by the

OFFICE OF NAVAL RESEARCH

and

DAVIDSON LABORATORYSTEVENS INSTITUTE OF TECHNOLOGY

September 28 - October 4, 1966Washington, D.C.

RALPH D. COOPERSTANLEY W. DOROFFEditors

ACR-136

OFFICE OF NAVAL RESEARCH-DEPARTMENT OF THE NAVY

Washington, D.C.


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RESEARCH ON THE MANOEUVRABILITYAND PROPULSION OFVERY LARGE TANKERS

J. D. van Manen, M. W. C. Oosterveld, and J. H. WitteNetherlands Ship Model BasinWageningen, The Netherlands

ABSTRACT

This pa pe r deals with the re sul ts of inves tigat ions into the ma noe uvr abili ty and pro puls ion of lar ge ta nk er s, equipped with an advance d ste rnar ran gem en t for propulsio n and ship- cont rol . Author s suggest toeliminate partly or completely the rudder, provide the ship with an extr em el y ci gar -sh ap ed afterbody having a shrou ded pro pe lle r with alar ge hub- to -d ia me te r ra tio and fit both bow and st er n th ru st er s to theship. The appl icat ion of cargo -pu mp- dri ven lat er al th ru st er s workingon the eje ctor prin cipl e is dis cus sed .

1. INTRODUCTION

The requirements for the manoeuvrability and propulsion of very largeta nk er s, to be built in the nea r future, open the question whether conventionalsolutions for the stern arrangement still have to be maintained.

The increasing displacement of tankers results in an increase of the required shaft horsepower. The nonuniformity of the flow at the propel le r and thehigh required shaft horsepower may lead in many cases to vibration troublesand cavita tion-e rosion damage on the propel ler bl ades . The increasing size oftankers and the difficulty of developing port areas large enough to accommodatethem will lead to higher requirements with respect to the effectiveness of theship-control devices. The effectiveness of a rudde r strongly depends on theship speed. As speed de cr ea se s, the effectiveness of the rudder becomes le ssand less.

In an attempt to provide large tankers with propulsion devices with superiorcavitation and propeller-induced vibration characteristics in addition to a highpropulsive efficiency and to provide tankers with a greater degree of manoeuvrability, the authors suggest to eliminate the rudder, provide the afterbody of anextremely cigar-shaped stern with a shrouded propeller having a large hub-to-diameter r at io and fit both bow and st ern th ru st er s to the ship. In Fig. 1, theinstallation of the propeller in a clear plastic nozzle on the afterbody of a tanker

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van Manen, Oosterveld and Witte

Fig. 1 - Ar ran gem ent of pro pel lerin noz zle for a ta nk er mod el

model as suggested by the authors is shown,thruster fitted to the model.

Figure 2, shows the bow and stern

The draft, propeller diameter, efficiency, propeller cavitation, propeller-induced vibration, construction of the rudder and so on, are all design factors,

influenced favourably by the solution indicated by Figs. 1 and 2. To obtain acomparison between the propulsive performance and the manoeuvrability of atanker with a conventional stern arrangement and the one according to Figs. 1and 2, tests were carried out with models representing a 90,000-ton deadweighttanker . The .re sul ts of these experiments are presented and discussed in thi spaper.

REGULATING VALVE PUMP MOTOR .

NOZZLE

WATER INTAKE

=0= fiteFig. 2 - Bow and st er n th ru st er ar ra ng em en ts fi t ted to the mode l

240


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Manoeuv rabili t y and Pro pul sio n of Ver y Lar ge Ta nke rs

Fixed and controllable-pitch impellers are the most common prime movingdevices located in tra nsv ers e thr us te rs. In recent liter atur e the use of car go-oil pumps of tanke rs as bow th rust er devices has been suggested: Such a sys

tem has the drawback that the efficiency is low due to the large kinetic energylosses in the sli pst ream. These los ses can be diminished by using an ejector inwhich a high-velocity jet with a low mass flow rate can be converted into alower velocity je t with a higher mass flow. There for e, the author s suggest touse bow and st ern thr us te rs working according to this ejector princi ple. Figur e3 shows schematically an ejector bow thruster.

Fig. 3 - Ar ra nge men t of ejec tor bow thr us te r

RESISTANCE AND PROPULSION

The trend in the design of tankers has been toward large dimensions atnearly constant speed and the refo re high-powered ships . The high requ iredshaft horsepower and the nonuniformity of the flow at the propeller disk maylead to vibration troubles and cavitation-erosion damage on the propeller blades.

In order to minimize the chance of propeller-induced vibration and cavitation damage, the inequality of the flow at the propeller disk must be reduced.Investigations performed by van Manen and Kamps (1), Nitzki (2) and Hadler andCheng (3) have shown that extreme shapes of afterbody for obtaining a uniformcircumferential velocity field in way of the propeller offer favourable prospects.

As a consequence of the decreasing propulsive efficiency with increasingshaft horsepower of tankers with conventional screws in addition to propellercavitation and vibration problems, the application of ducted propellers (1,4)tandem propellers, contrarotating propellers (5) and twin-screw devices (1,5,6)has been suggested. From these propulsion devices, the ducted pro pel ler andthe tandem propellers will be less expensive and complicated than contrarotating

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and twin-screw install ations. In thi s case of high-powered tank er s, the ductedpropeller is, according to the experience of the N.S.M.B., superior in propulsive efficiency compared to all other prope ll er s.

First, the considerations will be given which have led to the choice of theshape of afterbody and propul sor of the tanker discussed in thi s paper. Then,the results of resistance and self-propulsion model tests will be presented foreach of the conventional and the suggested versions of the tanker.

Effect of Afterbody on Propulsion

Investigations into the effect of the afterbody shape on propulsion of a39,000-ton deadweight tanker model carried out at the N.S.M.B. were reportedin Refs . 1 and 4. The following var iat ions in afterbody were studied (Fig. 4).

AFTERBODIES I

0 APP 20 FPP

AFTERBODY W

0 APP 30 FPP

Fig . 4 - Body plan and pr op el le r ar ra ng em en tof afterbodies I, II, III and IV

Model I, having an afterbody with moderately U-shaped sections and arudder shoe.

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Manoe uvrab ili ty and Pro pul sio n of Ver y Lar ge Ta nk er s

Model n , having an afterbody with extremely V-shaped sec tions and arudder shoe.

Model HI, having an afterbody with extr emely U-shaped sect ions and arudder shoe.

Model IV, having a cigar-shaped s te rn (Hogner-type of afterbody) and aMariner rudder arrangement (clear-water stern).

Fro m the experimental investigat ions, it was concluded that model n , hav ing an afterbody with ext remely V-shaped sections is to be recommended froma viewpoint of re si st ance . The improvement in res ist anc e qualitie s, however ,did not include a reduct ion of the requ ire d shaft horsepow er. The reduction inresistance is counterbalanced by a change in interaction effects between hulland propel le r. Invest igations performed by Nichols et al. (6) have led to the

same resu lt . Model I which is optimum according to the stat ist ical data of theN.S.M.B. is still optimum with respect to the delivered shaft horsepower at thedesign speed of 16 knots.

The cigar-shaped stern of model IV is bad from a viewpoint of resistance.However, the interaction effects between hull and propeller are such that in theloaded condition as well as in the light condition, model IV requires 2-3 percentmore power than the optimum hull form (model I).

Measurements of the longitudinal component of the velocity in the plane ofthe propel ler were made with the four models . Fro m these measurem ent s i twas concluded that the circumferential inequalities behind model n is worse

than that behind the models I and HI. Model IV shows the most homogeneouswake pat ter n of all afterbody varia tio ns .

Recently, the results of wake measurements made on a large number ofship models at the David Taylor Model Basin have been published by Hadler andCheng (3). Both the longitudinal and the tangential velocity components of thewake were measure d. Fro m experiments with a se ri es of tanke r models withsystematically varied stern shapes it was concluded that the models with themore U-shaped stern tended to generate a more uniform longitudinal velocityand a large downward velocity in the propeller plane near the propeller hub.The more V-shaped stern tended to give a large fluctuation in the longitudinalvelocity, while the tangential velocity is relatively smaller in magnitude as is

the downward flow near the hub. From an analysis of the effect of the nonuniform velocity field on pro pel ler cavitation and propeller -induced vibration, itwas concluded that the moderate U-shaped stern would provide best vibrationand cavitation characteristics.

Wake measurements on a ship model with a Hogner-type stern have alsobeen presented by Hadler and Cheng (3). From a comparison of the re su lt s ofthe Hogner-shape stern design and the conventional stern design from thestandpoint of minimizing vibration and cavitation problems in terms of uniformity in flow, the super ior ity of the Hogner-type st ern was evident. This conc lusion is in accordance with the results of observations of the cavitation patterns

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and mea surem ents of thrust and torque variations of the sc re ws , performed atthe N.S.M.B. with the Models I and IV.

Choice of Propulsion System

Among the special types of propulsors which have as their object an improvement in the propulsive quality of the ship, the ducted propeller takes animportant pl ace. The ducted propel ler is now extensively used in those caseswhere the ship screw is heavily loaded or where the screw is limited in diamete r. The possib ili ty of using the ducted prope lle r for the propu lsion of tugs ,towboats, and trawlers has been adequately demonstrated in practice in thecourse of the last thirty years.

In previous publications of the NlS.M.B. (7-10), the results of systematic

exper iment s with screw ser ie s in nozzles were given. These experiments su pplied design data for propell er- nozzl e sys tems with optimum efficiency. Thestandard nozzle profile applied by the N.S.M.B. from the viewpoint of efficiencyis nozzle No. 19a, having an angle between nose-tail line of nozzle profile andpropeller shaft axis ai = 9.2, a camber rat io f/e = 0.069, a maximum thickness rat io s/-t = 0.15 and a nozzle length-diameter rat io 't/D = 0.5. The or di -nates of this profile are given in Fig. 5.

NOZZLE No. 19"

Fig. 5 - Profile and ordinatesof nozzle No. 19A

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Manoeuvrability and Propulsion of Very large Tanker s

A method used frequently for expressing the characteristics of a propellertype is the relationship between Bp, S, and r;p for optimum propeller diameter.The design coefficient Bp and the speed ra tio 8 ar e defined as

NP1/2

y5/2

and

ND

V

where N is the number of revolutions per minute , the power P is in ho rsepower,the speed of advance ve is in knots and the screw diamete r D is in feet. Theoptimum relationship between B , S, and 7?p is given in Fig. 6 for the Ka 3-65,

Ka 4-70, and Ka 5-75 screw series in nozzle No. 19a and for the B 4-70 screwseries . At the top of the figure the ranges of Bp-values typica l for differentship types ar e indicated. The light ly-loaded scr ews of fast ships are at theleft-hand side, while the heavily loaded propellers of towing vessels are at theright.

TANKERSSINGLE SCREW C'OASTERS ^TRAWLERS ^TUGSCARGO SHIPS

S-65 SCREW SERIES IN NOZZLE-70 i i. NO 19KoS-65 SCREW SERIES IN NOZZLE-70 i i. NO 19

0.70

Vp*0 .60

0.50

0.40

400

350

S300

250

200

150

Ka 5 - 7 5 " " '

B 4 - 7 0 '0.70

Vp*0 .60

0.50

0.40

400

350

S300

250

200

150

Vp

0.70

Vp*0 .60

0.50

0.40

400

350

S300

250

200

150

Vp

^

0.70

Vp*0 .60

0.50

0.40

400

350

S300

250

200

150

^

0.70

Vp*0 .60

0.50

0.40

400

350

S300

250

200

150

~~~7/

0.70

Vp*0 .60

0.50

0.40

400

350

S300

250

200

150

y

/

^^

0.70

Vp*0 .60

0.50

0.40

400

350

S300

250

200

150

^^s*

^ f

0.70

Vp*0 .60

0.50

0.40

400

350

S300

250

200

150

.^ ^-

. * '

0.70

Vp*0 .60

0.50

0.40

400

350

S300

250

200

150

8' ^

^^">^

0.70

Vp*0 .60

0.50

0.40

400

350

S300

250

200

150

^J> -"^ ^ " '

20 30 40Bp

50 60 70 80 90 100 120 140 160

Fig. 6 - Comparison of the Ka-screw seriesin nozzle No. 19A and B4-70 screw se ri es

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Figure 6 shows clearly the reasons why ducted propellers are used at heavypropeller -loadings occurrin g in tugs, tra wle rs, and large tanker s. As tanker

si zes continue to grow, the advantage of ducted pr op el le rs , higher efficiency,and reduced optimum diameter, will become greater.

Analysis of the reduction in SHP due to the use of a nozzle has been madefor tank er s with an ins tal led power of 30,000 SHP and different deadweight. Therotative speed of the impell er has been fixed at 100 RPM. The resu lt i s givenin Fig. 7 which clearly shows the great reduction in SHP which can be obtained.For a tanker of 100,000 tons deadweight, a reduct ion of 8 percen t in SHP isattainable.

15r

c

u>

z

zot(J=> 5atua.

00 50000 TOW 100.000 150000 200.000

383 484 Bp 73.4 834

Fig. 7 - Reduction in SHP, due to theapplication of a nozzle for large tankers

Applications of Hogner-type Afterbody with Ducted Propeller

From the results of the investigations discussed so far it can be concluded

that for large tankers, the Hogner-type afterbody provided with a complete ringnozzle propeller system offers favourable prospects with respect to minimizingvibrat ion and cavitation probl ems in addition to a high propulsive efficiency. Inorder to obtain data on the differences in resistance and propulsive efficiency oftankers with a conventional stern arrangement and a Hogner-type stern withnozzle fitted, te st s were car ri ed out at the N.S.M.B. with models rep resent ing32,500-TDW and 48,500-TDW ta nk er s. Both tanker s had a conventionally shapedbow. The various stern arr ang ements considered were as follows:

32,500-TDW tanker : conventional st ern (moderately U-shaped sec tions andrudder shoe)

Hogner-type Clearwater stern fitted with a nozzle

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Manoeu vrabi li ty and Pro pul sio n of Ve ry La rg e T an ke rs

48,500-TDW tan ker : conventional st er n (cl earwater st er n with moderatelyU-shaped sections)

Hogner-type stern with rudder shoe and nozzle

Details of the st ern ar rangeme nts a re pres ente d in Figs. 8 and 9. Resi stanceand self-propulsion tes ts were run for the light and loaded condition. Table 1compares the results of the EHP and SHP tests of the 32,500-TDW tanker model.The results for the 48,500-TDW tanker are shown in Table 2.

Table 1EHP and SHP Tests of 32,500-TDW Tanker

(design speed 17 knots at 16,000 SHP)

Speedin

Knots

Hogner-Type Stern Bet ter (+) or Worse (-)

Than Conventional Stern ArrangementSpeedin

KnotsEHP SHP

Speedin

KnotsLoaded

ConditionLight

ConditionLoaded

ConditionLight

Condition

14

15

16

17

-6.2%

-6.3%

-6.0%

-5.1%

-5.1%

-6.5%

-8.0%

+13.1%

+11.9%

+ 8.6%

+ 4.8%

+0.0% g

+0.3% u 3

-0.3% * g

Fig . 8 - Detai ls of ste rn arr an ge me nt s for 32,500-TDW tan ker model

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Fig. 9 - Details of stern arrangements for 48,500-TDW tanker model

Table 2EHP and SHP Tests with 48,500-TDW Tanker

(design speed 16 knots at 16,000 SHP)

Speedin

Knots

Hogner-Type Sterji Bet ter (+) or Worse (-)Than Conventional Stern Arrangement

Speedin

KnotsEHP SHP

Speedin

KnotsLoaded

ConditionLight

ConditionLoaded

ConditionLight

Condition

13 -5.3% - +4.4% -

14 -7.7% - +4.1% -

15 -7.2% - +3.8% +3.6%

16 -6.7% - +2.7% +3.7%

17 -9.6% - +2.8% +4.5%18 -9.6% - +2.9% +5.1%

From the tabl es , it may be seen again that the Hogner-type afterbody isreally bad from a viewpoint of resistance; however, in the loaded as well as inthe light condition, the tankers with the Hogner-type stern and nozzle requirele ss power than the conventionally shaped tank er s. The reduction in SHP is 4.8percent for the 32,500-TDW tanker in the loaded condition at the design speedof 17 knots. A reduct ion in SHP of 2.7 pe rcen t has been found for the 48,500-TDW tanker in the loaded condition at the design speed of 16 knots.

248


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Manoeuvr abili ty and Pro pul sio n of Ver y La rg e Tan ker s

Advanced Propulsion for a High-Powered Tanker with Bowand Stern Thruster Control

In the introduction the elimination of the rudder was suggested by providinglarge tank ers with bow and st ern th ru st er s for tr an sv er se control. The el im ination of the rudder opens the poss ibili ty of extr emely cigar-shaped s te rn s. Acomplete nozzle ring can be fitted then, attached with four brackets to thisafterbody.

The presence of a rudder on a single-screw ship tends to improve the propulsive efficiency. In addition, the tailp iece may be at tractive from a viewpointof course stabilit y. The refore , a fixed tailpiece must be fitted to the nozzlering. In Fig. 1, such an afterbody ar rangement is shown.

Model tests have been carried out to obtain a comparison between the propulsive quality of tankers with a conventional stern arrangement and with thest er n arr angemen t indicated in Fig. 1. The different s te rn designs have beenmade for a 90,000-ton deadweight tanker, having an engine power of 18,000 SHPand a speed of 15.75 knots. The principal part ic ul ar s of the tanker a re given inTable 3.

Table 3Principal Particulars of Tanker

Loaded Condition

Length between perpendicularsBreadthDraftWetted surfaceMoulded volumeHalf angle of entranceBlock coefficientMidship ar ea coefficientPrismatic coefficient

Light Condition

Draft 8.13 mMoulded volume 65,664 m3

The body plans and the screw arrangements of the afterbody variations arepresented in Fig. 10. The scale of the model is 38.

The cha racte ris tic s of the various prope lle rs a re given in Fig. 11. Thescrew for the conventional model was designed according to the circulationtheory for wake-adapted propellers.

253.0 m40.0 m14.30 m

14,980 m2

120,171 m3

15 48'0.8310.9960.834

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Fig. 10 - Body plan s and pr op el le r ar ra ng em en tsfor the 90,000-TDW tanker model

The model with Hogner-type st er n has been tes ted with thr ee different no zzl es . Nozzle No. 1 is der ived from nozzle No. 19a by taking into account the effect of the hub shape of the flow. The taper of the hub is accounted for by lo ca ting the nozzle profile with reference to the direction of the undisturbed flow.The direction can be determined if one assumes frictional effects and staticpr es su re va ria tions to be negligible over the axial distance of the nozzle. It i sthen a ma tt er of satisfying the law of continui ty. This method is only permi tt edif the shape and the dia meter of the hub are not too ex tr em e. The design ca lculations for the ducted screw propeller were based on the method given in Refs.9 and 10. The nozzles No. 2 and No. 3 have in compariso n with nozzle No. 1, a

mor e pronounced converging par t before the sc rew . This may be att rac tivefrom a viewpoint of a mo re homogeneous flow in the nozz le . The pa rt ic ul ar s ofthe nozzle profiles Nos. 1, 2, and 3 are given in Fig. 12.

Model resistance and self-propulsion tests were conducted in the deep-wate r basin in accordance with establish ed proc ed ures . All model data wer eextrapolated to the full-scale ship values using Sch oen her r' s friction coefficientswith an addition of 0.00035 for co rrelat ion allowan ces . A tr ip wire of 1 mmdia meter, was fitted around each model at station 5 percen t LBP for turbul encestimula tion. The re su lt s of the re si stan ce and self-p ropuls ion te st s in theloaded and in the light condition are given in Tables 4 and 5.

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Manoe uvra bili ty and Pro pu lsi on of Ver y Lar ge T an ke rs

PITCH DISTRIBUTIONIN PER CENT

D=6700mmP 0 7 R / D =

0 - 6 3 4

Z =4A D / A Q = 0.565

p / A o =0.526

0=6300P07R/D=-9

5 7

Z=4A D / A 0 =

0 - 5 2 8

A p / A o = 0.510

Fi g. 11 - Ge ne ral pla ns of s cr ew s

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NOZZLE No. I

No. X

Wo. 3

Fig. 12 - Profiles for nozzles No. 1, 2, and 3

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Manoeu vrabi lity and Pro pul sio n of Ver y Larg e Tan ke rs

Table 4Results of Resis tance and Self-Propulsion Te st s in Loaded Condition

Speedin

Knots

Resis tanceTests EHP

Metric

Se l f -Propuls ion TestsSpeed

inKnots

Resis tanceTests EHP

MetricSHP

Metric

RPM

Screw ^ t o t .

Conventional

model

1314151617

9,002

11,164

13,790

17,198

21,826

13,72117,07921,02326,41734,399

98.9106.7114.5123.8134.9

0.6560.6540.6560.6510.634

Hogner- typestern with

nozzle No. 1

1314151617

8,43111,03513,87717,07421,203

12,62615,78019,82425,13732,164

99.7107.3

115.6

125.0

135.5

0.6680.699

0.700

0.679

0.650

Hogner- type

stern with

nozzle No. 2

1314151617

8,43111,03513,87717,07421,203

13,46316,66920,75626,15733,324

100.3108.0116.4125.8136.3

0.6260.6620.6690.6480.636

Hogner- typestern with

nozzle No. 3

1314151617

8,43111,03513,87717,07421,203

14,36717,83922,24128,03434,862

104.8112.9121.5131.2141.6

0.5870.6190.6240.6090.608

Tabled compares the results of the EHP and SHP tests run at the loadedcondition of the conventional model and the model with the Hogner-type st ernwith nozzles Nos. 1, 2, or 3. Table-7-compares the re su lt s at the light condition.

P From the tab les , it may be concluded that the extremely cigar -shapedafterbody is favourable from a viewpoint of re si st an ce . The EHP of the model

{ with the extremely ci gar-shaped afterbody is about 1 percen t les s in the loadedLcondition and 5 percent less in the light condition, compared with the conven

tionally shaped model.

It is worth noting that this conclusion is cont rary to cigar-shaped sternsinvestigated up til l now. The tanker with the Hogner-type st ern and nozzle No.1 fitted, re qu ir es les s power than the Hogner-type ster n with nozzles No. 2 and oNo. 3 and the conventional st ern. The Hogner-type ster n with nozzle No. 1, "^ "Z^(1fitted requires about 6 percent less SHP in the loaded condition and 9 percentles s SHP in the light condition, "ERgfTthe conventionally shaped ste rn .

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Table 5Results of Resistance and Self-Propulsion Tests in Light Condition

Speedin

Knots

Resis tanceT e s t s EHP

Metr ic

Se l f -Propuls ion Tes tsSpeed

inKnots

Resis tanceT e s t s EHP

Metr icSHP

Metric

RPM

Screw ' ' t o t .

Conventional

model

14

15

16

17

18

9,014

11,348

14,154

17,806

22,576

11,851

14,862

18,673

23,855

31,131

94.2

101.2

109.0

118.2

128.9

0.761

0.764

0.758

0.746

0.725

Hogner- typestern withnozzle No. 1

1415161718

8,30210,54913,60517,05221,778

10,06913,80517,08521,75428,645

94.0101.2108.4117.6128.7

0.7500.7640.7960.7840.760

Hogner- typestern withnozzle No. 2

1415161718

8,30210,54913,60517,05221,778

11,47714,24617,64622,39429,409

94.6102.0109.4118.6129.8

0.7230.7400.7710.7610.741

Hogner- typestern withnozzleNo. 3

1415161718

8,30210,54913,60517,05221,778

12,53115,61419,33824,09630,803

99.1106.8114.9123.5134.1

0.6630.6760.7040.7080.707

Table 6Hogner-Type Stern Better (+) or Worse (-) Than Conventional

Stern Arrangement at Loaded Condition

Speed

in

Knots

Res i s t anceTests

EHP

Self-Propulsion Tests

SHPSpeed

in

Knots

Res i s t anceTests

EHP Nozzle No. 1 NozzleNo. 2 Nozzle No. 3

1313.5

1414.5

1515.5

1616.5

17

+6.3%+4.4%+ 1.2%-0.5%

-0.6%+0 %+0.7%

+ 1.9%+2.9%

+8.0%+8.2%+7.6%+6.4%+5.7%

+5.4%+4.8%

+5.3%+6.5%

- +1.9%+2.8%

+2.4%+1.9%+1.3%

+1.2%+ 1.0%+1.9%+3.1%

-4.7%

-3.7%-4.4%-5.2%-5.8%

-6 .1%

-6 .1%

-5.0%

-1.3%

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Manoeu vrabil ity and Pro pul sio n of Very La rg e Tan ker s

Table 7Hogner-Type Stern Better (+) or Worse (-) Than Conventional

Stern Arrangement at Light Condition

Speedin

Knots

Resis tanceTestsEHP

Self-Propulsion Tests

SHPSpeed

inKnots

Resis tanceTestsEHP

Nozzle No. 1 Noz zle No. 2 Nozz le No. 3

13.5

14

14.5

15

15.5

1616.5

17

17.5

18

18.5

+6.6%

+7.9%

+7.9%

+7.0%

+5.3%

+3.9%+3.4%

+4.2%

+4.0%

+3.5%

+3.4%

+6.5%

+6.6%

+6.9%

+7.1%

+7.6%

+8.5%+8.9%

+8.8%

+8.0%

+8.0%

+8.0%

+2.3%

+3.2%+3.3%+4.1%

+4.6%

+5.5%+6.1%+6.1%

+5.6%+5.5%

+5.9%

-5.7%

-5.7%-5.5%

-5 .1%

-4.4%

-3.6%-2.0%-1.0%-0.2%+1.1%

+2.3%

Finally, it may be concluded that an afterbody arrangement as suggested inthe introduction offers a means of improving the propulsive efficiency for largetanker s in addition to minimizing vibration and cavitation probl ems . Economical considerations with respect to the increase in building costs and the de

crease in costs as a consequence of the smaller required SHP and probablylower costs of maintenance will give the final answer to the shipowner on thequestion in how far application of this advanced propeller arrangement for high-powered tankers is not only a dream of the hydrodynamist.

3. A COMPARATIVE STUDY OF THE STEERING OFTANKERS WITH LATERAL JETS AND WITHA RUDDER

As stated in the introduction of this paper, the technical merits of steeringa tanker with lat eral thrust units instead of with a rudder a re twofold. Fi rst ly

by eliminating the rudder a more favourable afterbody shape-pr opulsor combination can be designed resul ting in a substantial dec rease in requi red shafthorsepower and improved vibration and cavitation characteristics; secondly abetter manoeuvrability at low speeds in confined areas, harbours, and riverestuaries can be expected resulting in partly or completely eliminating tug boatassistance, the economic benefit of which will be important to tanker owners.The main problem of the jet steering units ar e their horsepower req uir ement sat the higher ship speeds.

English (11) and Chis lett -Bjrheden (13) have measured the side force c r e ated by a lat er al th rust unit as a function of the rat io ship speed to jet speed v/u.English found a 50 percent decrease in side force for a typical screw-driven

la te ra l thrust unit when the ship speed was increased from 0 to 2 knots . This

255


19/38


effect was also measured by the authors of Ref. 13; they went to higher v/u va lues than English and establ ished that the total side force and effective moment

act ing on their model both had a min imum for v/u at about 0.4 to 0.5.

This peculiar behaviour can be understood from the work of Keffer andBaines (14). They found that for v/u values higher than 0.2, a jet which wasinitially blown through a hole in a wall perpendicular to the mainst ream tu rns90 and tends to cling to the wall crea ting a region of low pr es su re downstreamof the nozzle. This low pr es su re region cre at es a suction force which tends todiminish the effect of the thrust unit. For ships with jet stee ring this is an impor tant effect since the center of action of the suction force moves aft for higherv/u values , hence the turning effect on the ship is no longer proport ional to theside force on the ship and must be assessed in terms of both side force andturning moment.

A qualitative view of the effect of the forces and moments involved is givenby Fig. 13.

///// LOW PRESSURE REGION

v = o

w

M s =Tx

T K = T

Fig. 13 - Forces and moments on amoving ship with a bow thruster

When v is increased, the cen ter of the suction force s moves downstreamdecreasing the effective turning moment Ms on the ship. This turning momentpasses a minimum and then in cr ea se s again. When s passes the center of theship, MShas the same value as when v = 0. In Fig . 14, taken from Ref. 13, thedimensionless moment Ms/pAu

2x is plotted against the velocity rat io v/u , and

the curve clearly disp lays the effect. The line Ts/pAV2

against v/u is also

256


20/38

Manoeu vrabil ity and Pro pul si on of Ve ry La rge Ta nk ers

Fig. 14 - The di men si onl ess turningmoment and side force as a functionof V/U ac co rding to Ref. 13

shown. Since T = pAU2 and T = Ts+ s signifies the constant thr ust of the l ateralth rus t unit, it appea rs that Ts increases, passes a maximum and then decreaseswith inc reas ing ship speed. Fro m these consid erat ions it is clear that bowthruster ship interaction is very important when studying the manoeuvrabilitycapacities of various bow thrusting systems.

Description of the Experiments

As mentioned in the preceding sections, two tanker models were built; onemodel with a conventional screw rudder arrangement and one model fitted witha Hogner afterbody, accelerating nozzle, bow and stern thrusters but without arud der . In order to compar e jet stee ring with conventional rudder ste eri ng itwas intended to carry out z manoeuvring as proposed by Kempf in the shallowwater basin of the N.S.M.B. For the z manoeuvring, the conventional modelwas equipped with a servo motor for the rudde r angle adjustment. The mode lwith jet steering was supplied with a centrifugal water pump and piping leadingto the four exhaust nozzles situated at the bow and the ster n of the model. Inthe water supply pipes, two venturi m as s flow meter s were mounted for me as uring the wate r mas s flow to the bow and ste rn th ru st er s. Servo motor -dr ive nvalves built in the water delivery pipes directed the water flow either to the port

257


21/38


or the starboard nozzles in such a manner that always a left- or right-handtorque acted on the ship . The time lag of the valve action was very smal l com

pared with the swinging period of the ship (Fig. 2).

During the z manoeuvring experiments the relevant quantities were measured using the following methods:

Water mass flow from two mercury filled U-tubes connected with the ven-turi nozzles.

Rudder angle 8 was adjusted on the servo motor moving the rudder .

Yaw angle 41and the drift angle /3 were measured with a gyroscope mounted

on the model.

Yaw angle rate 41 = d^/dt was derived by electronic differentiat ion of thegyroscope yaw angle signal with respect to the time t .

The nozzle exhaust velocity u was calculated with the aid of the continuityequation from the measured water mass flow rate and the known nozzle exhaustdiame te r. During the experiments the following quant ities wer e changed: thedraft H, the ship velocity V, the nozzle diameter Dn, the nozzle exhaust velocity U for the jet s teer ing model, and the rudde r angle S for the conventionalmodel.

The following experimental program was started:

Jet Steering

H = 8.13, 14.3 m

V = 5, 9, 13, 17 knots

Dn = 1.10, 0.357 m

U = 8.95, 12.1, 15, 17.9, 20.8 m/sec for Dn = 1.10 m

u = 18.6, 38.4, 58.5, 77, 97 m/sec for Dn = 0.357 m

Rudder Steering

H = 8.13 m, 14.3 m

V = 5, 9, 13, 17 knots

s = 5, 10, 15, 20, 25 degrees

Fo r the adjustment of the nozzle exhaust velocity at model scale , Froudescaling was used. Thus the velocity ra tio v/u for model scal e and tru e scaleremained the same.

The experiments were done in such a way that the model propelled itselfwith a certain desired model speed before rudder or jet action was initiated,after which the response of the model to alternating rudder movement or jet

258


22/38

Manoeuvr abili ty and Pro puls ion of Ver y Lar ge Tan ker s

action was meas ured as a function of time. A qualitative diagram of the different quantities such as \p,


23/38


RUDDER STEERING

\j/ z f -. o.

Fig. 16 - Definition sketch of some quantities given in Fig. 15

with the model with rudder steering, /max as rudder steering with maximum rudder deflection, in the range of

260


24/38

Manoeuvrabi l i ty and P ropuls ion of Ver y Large Tan ker s

o.o -

.!' M M .

1\

SET STEt-RiNc;: O ^ t R . ' i m

RU-SERS T t C R i N ^ : 0 ^s S -' w

^

o 6 (

w)--30 40

5ooo o.l - % 0.1 o.4i

IOOO 5oo too-a- Nj,(HP)

Fig . 17 - Yaw ra te \jj aga inst rud der angle S andveloci ty rat io V/U;V= 5knots , D = 0.357 m

o.6

t

HVi^x.

o.Z

STCt

H I M C , : O *

Q

BtNC :

- S l i m

Is K S n i

^* 8 . i3m

(= H . Sm

^Tr 0^Tr 0

20 - 6 ( ) -4 o

) 1IOOOO Sooo

o . a - V / a o.* o.4

s o W t(HP)-

Fig . 18 - Yaw ra te


25/38


I0OOO Sooo 500 Nj,tHP)

Fi g. 19 - Yaw ra te

V3CT ST

~RulttKltSTt

HINC

KIIXC :

O 1

a H

M N

= l U i n

e S.isrti

J r~^, ""i 1

20 6 c ;- 4o

-Vu- 0.4H-H

50OO Mf (Hf)-

Fig. 20 - Yaw rate \paga ins t r ud de r an gle S andve lo ci ty ra ti o V/U; V = 9 kn ot s, Dn = 1.10 m

262


26/38


27/38


Table 9Je t Steer ing (H = 8.13 m)

V = 5 Kn ot s, Dn == 1.10 m

U (m/sec) Nj (PK) v/u 08.95

12.11517.920.8

4501,1202,1503,6005,650

0.2&00.2060.1680.1400.120

0.160.180.320.390.46

V= 9 Knots, D = 1.10 m

U (m/sec)17.920.8

Nj (HP)3,6005,650

V/U

0.2470.214

0.330.41

V = 5 Kn ot s, Dn = 0.357 m

U (m/sec) Nj (HP) V/U 0

18.638.476.8

97

4203,750

44,00063,000

0.1340.0650.0330.026

0.060.190.540.70

V = 9 Kno ts, Dn = 0.357 m

U (m/sec) Nj (HP) V/U "max

58.276.897

13,20044,00063,000

0.0770.0590.047

0.320.460.62

V = 13 Kn ot s, Dn = 0.357 m


76.897

44,00063,000

0.0850.067

0.480.62

V = 0 Knot , Dn == 1.10 m

U (m/sec) Nj (HP) V/U 0rmax

20.8 5,650 0 0.445

Table 10Je t Steering (H = 14.3 m)

v = 5 Kn ot s, Dn = 1.10 m

u (m/sec) Nj (HP) V/U 0r max

8.9512.11517.920.8

4501,1202,1503,6005,650

0.2800.2060.1680.1400.120

0.150.160.180.200.26

V = 9 Knots, Dn= 1.10 m

u (m/sec)17.920.8

N: (HP)3,6005,650

V/U

0.2470.214

0.200.25

V = 5 Knot s, Dn = 0.357 m

U (m/sec ) NL (HP) V/U 0~max

38.458.577.097.0

3,75013,20044,00063,000

0.0650.0430.0330.026

0.150.240.320.48

V = 9 Kno ts , Dn = 0.357 m

U (m/sec) Nj (HP) V/U 0r ma x

38.458.577.097.0

3,75013,20044,00063,000

0.1180.0770.0590.047

0.190.250.290.43

V= 13 Kn ot s, Dn = 0.357 m


58.577.097.0

13,20044,00063,000

0.1110.0840.067

0.250.330.45

V = 0 Knot, Dn = 1.10 m

U (m/sec) Nj (HP) V/U 0rmax

20.817.915

5,6503,6002,150

000

0.7050.6000.547

264


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Manoeuvrability and Propulsion of Very Large Tankers

Table 11Rudder Steering

H= 8.13 m V = 5 Knots V = 9 Knots V= 13 Knots

s 0~max

'ma x 4>"max

57.5

101520253035

0.05

0.07

0.110.130.150.16

0.090.150.150.200.250.290.330.36

0.120.170.200.290.370.410.460.54

H = 14.3 m V= 5 Knots V= 9 Knots V= 13 Knots

s"max rmax 4>

^max

57.5

10152025

3035

0.05

0.06

0.120.100.130.14

0.070.100.150.150.220.25

0.280.30

0.150.160.220.300.330.37

0.410.47

We point out that Vmax was not corrected for the increasing effectivity ofthe rudder due to propel ler overload on model scale during the z manoeuvringte sts . This should decr ease the averag e 0 max value for rudder steering by 10percen t. So the above given re su lt for jet stee ring superiori ty is on the conserva tive s ide . After contemplating the re su lt s of the experim ent s, it was feltthat the nozz les test ed were chosen too sma ll. Whether the ship can be m a-noeuvered up till the design speed of 15.75 knots, with a larger nozzle, Dn = 2-4m say, rema in s to be seen. However nozzles with these smal l exhaust di ame

ters are attractive in connection with main cargo pump-driven steering systems.These installations will be described in the following section.

4. MAIN-CARGO-PUMP-DRIVEN JET-STEERING SYSTEMS

Simple Nozzle System

For a description of conventional lateral thrust units we refer to the publication of Jastram (12) and of Stuntz and Taylor (15).

265


29/38


Recently the idea of using the main cargo pumps of tanker s for steeri ngpurposes has been put forward. It is intended to deliver sea water with thesepumps to nozzles situated at stern and bow, the efflux of these nozzles creatinga lat eral force and a turning moment on the ship . The economic incentives forusing this system are that expensive machinery which is otherwise idle is put towork for manoeuvring the ship without any additional rotating machinery. Theposs ible elimination of tugboat as sist ance is al so an impor tant considerat ion.

A typical simple thrust generating syst em is illust rate d in Fig. 21 . Themain cargo pumps deliver a certa in volume flow rat e with a cert ain head. Thepressure increase across the pumps is needed to cancel the pressure drop inthe sys tem caused by friction in the pipel ines and appendages, the height dif ference between sea chest and nozzle exhaust, and the pressure drop in the nozzle.Generally the last pressure decrease is the largest, the friction losses in thepipeline can become prohibitive only when the pipe diameter is chosen too small.

The th ru st er s can be modulated with a thro ttle -val ve in the pipeline. The p r es sure drop in the various pipes, valves, bends, etc., can be computed from datagiven in engineering textbooks. In a smoothly rounded exhaust nozzle (Fig. 22),the friction losses are very low compared with the friction losses in the pipelinesyste m; these los ses can be neglected when designing these sys tems. In thatcase Bernoulli's law results in

1 2 1 2

n o z z l e ~ ~2 ^ n o z z l e ~~ 2 ^ pi pe '

or using the law of continuity,

" "l e

772 ID4 , D4 . /

^ n o z z l e p i p e '

Fi g. 21 - Ar ra ng em en t of sim ple nozz le sys tem

266


30/38


W////////'

' ' NOZZLE SHAPE

Fig.22 - Exhaust nozzle

Q being the volume flow ra te through the nozzle. The genera ted thrus t at ze rospeed amounts to

77 2 21 ~ "4 ^ n o z z l e

Un o z z l e "

For the manoeuvring experiments described in the preceding section, the capabilities of an installation supplied with a nozzle exit diameter of Dn = 0.357 mmust be seen as (representative.

, . , j ' ;

The plain nozzle system has the drawback that cargo pumps have an unfavorable head-capacity relationship (high head, low volume flow rate) which re

sul ts in a large kinetic energy l oss in the jet leaving the nozzle . This is thereason why the plain nozzle system has a low thrust per installed pump horsepower of the order of 2-3 kg/hp at zero ship speed compared to 10 kg/hp for atypical tugboat at zero speed.

The Ejector System

By placing the exhaust nozzle described above in a tunnel with open ends atboth sides of the ship, an ejector system is obtained (Fig. 3). The jet leavingthe nozzle mixes with the surrounding water and creates a flow in the tunnel.Since this installation increases the water mass flow rate and reduces the ex

haust velocity a better thrust-horsepower ratio compared with the plain nozzleinstallation is obtained. Fro m experimen ts ca rr ie d out on a special 100-hpejector test bench at the N.S.M.B., it was found that a thrust increase factor oftwo over the plain exhaust nozzle can be achieved corresponding to a thrust of4-6 kg/hp for a prac ti ca l installat ion at ze ro ship speed. The main advantageof this system is that a more favorable thrust-horsepower ratio is obtainedwithout sacrificing the basic simplicity and absence of moving parts of the plainnozzle installation. The re su lt s of the manoeuvring te st s with a nozzle di ame ter Dn = 1.10 m should be seen as typica l for the behaviour of such an ej ec to r-driven steering system.

267


31/38


A further increa se of thr us ter efficiency i s to be expected from a two-stageejector design (Fig. 23). By using the exhaust of a smal l pr imar y ejec tor a s thenozzle for a big secondary ejector, the ma ss flow is increased further. It is ex pected that a thrust of 6-8 kb/hp can be attained at zero ship speed with a cargo-pump-driven sys tem of this configuration. Since the exhaust velocity of the pr i mary nozzle is of the order of 30-40 m/ sec , cavitation in the prim ary ejectormay create problems.

THRUST VECTOR

Fig . 23 - Ar ra ng eme nt of two -st ag e ejec tor sys te m

5. CONCLUSIONS

The main conclusions of this feasibility study are:

The extremely cigar-shaped afterbody is favourable from a viewpoint ofres ist anc e. This is contrary to cigar-shaped st ern s investigated up till now.The EHP of the tanker with the extremely cigar-shaped afterbody is about 1percent less in the loaded condition and 5 percent less in the light condition,

compared to the conventionally shaped tanker.

The extremely cigar-shaped stern with shrouded propeller having a largehub-to-diameter ratio, as suggested in the introduction, offers a means of improving the propulsive efficiency for large tankers in addition to minimizingvibration and cavitation problems. The Hogner-type ste rn with nozzle No. 1re qu ir es about 6 percent less SHP in the loaded condition and 9 percent le ssSHP in the light condition than the conventionally shaped ste rn . This improve ment is of the same magnitude of an extreme bulbous bow, which leads to a 15percent SHP reduction in the light condition.

The experiments show a definite advantage of the lateral thrust arrange

ment over the conventional rudder for ship speeds up to 5 knots. Test s using

268


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jet nozzles with larger diameters than 1.10 m may lead to an extension of thespeed range where this type of ship contro l has a superior manoeuvrability,lowering at the same time the required pump power.

Comparing the two types of cargo-pump-driven lateral thrust units, it isfound that at zero forward speed the ejector type doubles the specific thrust of2-3 kg per ins tall ed pump horsepower of the simple nozzle sys tem. At 5 knotsship speed, the nozzle with an exit diamete r of 1.10 m, which is represen ta tivefor the ejector system, requires about half the power of the plain nozzle systemwith a manoeuvrability which, when expressed in maximum yaw angle rate, isthe same.

Since it was found that at speeds above 5 knots the jet steering system canonly assist the rudder but not make it superfluous, the tail piece just behind thepropelle r nozzle should be used as a rud der. Therefore it should be made mov

able.

REFERENCES

1. van Manen, J.D. and Kamps, J. , "The Effect of Shape of Afterbody on Pr o pulsion," Trans. SNAME, 1959

2. Nitzki, L., "Einige Weitere Untersuchungen im Zusammenhang mit derAnwendung der A.G. "Weser" -Hi nte rschiffs form," Schiff und Hafen, 1962

3. Hadler, J .B. and Cheng, H.M., "Analysis of Experimental Wake Data in Way

of Propeller Plane of Single and Twin-Screw Models," Trans. SNAME, 1965

4. van Manen, J.D ., "Size , Type and Speed of Ships in the Fut ur e, " Th irdSymposium on Naval Hydrodynamics, 1960

5. Hadler, J.B ., Morgan, W.B., and Meye rs, K.A., "Advanced Propel le r P r o pulsion for High-Powered Single-Screw Ship," Trans. SNAME, 1964

6. Nichols, W.O., Rubin, M.L., and Danielson, R.V., "Some Aspect s of LargeTanker-Design," Trans. SNAME, 1960

7. van Manen, J.D. , "Open Water Test Ser ies with Pr opel le rs in Nozz les ,"International Shipbuilding Progress, Vol. 1, No. 2 (1954)

8. van Manen, J.D. , "Recent Researc h on Pr opel le rs in Nozzles," InternationalShipbuilding Progress, Vol. 4, No. 36 (1957)

9. van Manen, J.D. and Superina, A., "The Design of Screw Pr op el le rs inNozzles," International Shipbuilding Progress, Vol. 6, No. 55 (1959)

10. van Manen, J.D. , "Effect of Radial Load Distribution on the Per formance ofShrouded Propellers," Paper No. 7, The Royal Institute of Naval Architects,March 29, 1962

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11. English, J.W., "The Design and Perf ormance of La te ra l Thr ust Units forShips," Trans. R.I.N.A., July 1963, Vol. 105, No. 3

12. Ja st ra m, H., "Bu gstr ahlr ude r," Jahrbuch der Schiffbautechnischen Gesell -schaft 52, Band 1958

13. Chisle tt, M.S. and Bjrheden, O., "Influence on the Effec tiveness of aLa te ra l- Thrust Unit," 11th International Towing Tank Conference 1966 andHyA Report No. Hy-8

14. Keffer, J .F . and Baines, W.D., "The Round Turbulen t Jet in a C ros s -Wind," J. Fluid. Mech. 15, 1963, 481-496

15. Stuntz, G.R. and Taylor , R.J. , "Some Aspects of Bow Th rust er Des ign ,"Trans. SNAME, 1964

* * *

DISCUSSION

H. LackenbyBritish Ship Research Association

London, England

I was very interested to see the statement in the conclusions that the 9 percent improvement in SHP with the Hogner-type stern was of the same magnitudeas an extreme bulbous bow which leads to a 15 percent reduction in SHP in thelight condition:

In the first place, I would be glad to have some clarification on this matter,that is, whether the 15 percent is the figure for the bulbous bow alone orwhether it ref er s to the bulbous bow and Hogner -stern toge ther .

In this connection, I should like to mention an experience we had in carryingout bow and ster n vari ations to a tanker model. Fi rs t we did the bow varia tions

in conjunction with a conventional stern and obtained a small improvement ofthe order of 3 per cent ( this was not a ram bow, by the way). We then tri ed acigar-shaped stern in conjunction with a conventional bow and obtained an improvement, her e of about 3 percen t also. We then tr ied the bulbous bow andbulbous stern together, hoping of course that the two improvements would beaddit ive, but th is did not turn out to be the case . In fact the improvement wasbarely s ignificant. In other words, the effect of either the bow or st ern bulbalone was slightly better than both together.

I should like to underline here that the bulbs I am talking about were verymuch less pronounced than those dealt with by the authors, and to some extentwe were probably flapping around in the experimental scatter band.

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Manoe uvra bili ty and Pro pul sio n of Ver y La rg e Tan ke rs

In some analysis work we have been carrying out recently, however, involving more pronounced ram bows, there is a fairly clear indication that the bowmodification is affecting the propulsive efficiency, and one has to be very care

ful about considering the two ends of the ship separately and then assuming theeffects to be additive.

Perhaps the authors would comment on this question of bow and ste rn in te r

action.

* * *

DISCUSSION

J. Strom-TejsenDavid Taylor Model Basin

(after March 31, 1967 renamed the NavalShip Research and Development Center)

Washington, D.C.

It is certainly refreshing to read a paper like this in which the authors havestudied the application of an unconventional control system to very large tankers . It would have been more inte res ting had the proposed jet control systemshown definite m er it s as an alte rnat ive to the trad itional rudder configuration.Unfortunately the jet system alone according to the paper would require excessive power in order to produce forces comparable to those obtained by a conventional rudder . Also the au thor s' alt ernate proposa l, combining the jet systemand the small tail piece as a movable rudder, is questionable in its present configuration as it is unlikely that the small tail piece tested would suffice in areato produce the desi red maneuvering forces. It wouldibe of in te re st to know ifthe authors actually have explored this alte rnat ive further.

The extreme cigar-shaped afterbody having a shrouded propeller is apparently an attractive solution from a resistance and propulsion point of view sincecri tic al problems due to vibration and cavitation ar e reduced . But let us for amoment consider the merits of this form with respect to some of the problemsencountered in the field of controllability of very l arge tank er s. These ships ,for instance, are known to have very poor inherent course stability qualities.As deadwood has been reduced and a big conventional rudder eliminated in theproposed stern arrangement, it is possible that the dynamic stability problemwould become more critical in this case, and the proposal thus would be unrealist ic unless addit ional fin ar ea i s added. I think many of us would apprecia tethe authors commenting on this problem and possibly adding to the paper theresults from a course stability study.

The stopping and backing are other examples of maneuvers which by manyar e consider ed critica l for the very large tanker s. As shrouded pro pel ler s a reknown to be inferior to conventional propellers with respect to the retarding

and backing force which can be produced in a stopping maneuver, it is necessary

271


35/38


36/38


reduction in propeller horsepower compared with a conventional stern arrangement. Fro m the basic concepts it appears that a reduction in propelle r ho rs epower is possible either by reducing ship resistance or by improving the pror

pulsive efficiency. The present tes ts with a 90,000-TDW tanke r model showboth these improvements, the resi stance being reduced by 1 and 5 and the efficiency improved by 5 and 4 percent at the loaded and light condition, respectively. According to another well-known equation, the propulsive efficiency maybe exp res sed as the product of the propelle r, hull, and rel ative rota tive efficiencies . Unfortunately the paper does not give any information a s to the way inwhich these propulsive coefficients are affected by the proposed nozzle propellerarrangement, and I would like to ask the authors if any such information is available. It would al so have been in terest ing to know what efficiency one may obtainby using an unducted prope ller in combination with the Hogner ste rn . Fr om therelated work on a 39,000-TDW tanker model (Section 2 of the authors' paper),one gets the imp res sion that the Hogner s ter n per mi ts the propulsive efficiency

to be improved al so with a norm al free propel ler .

The idea of using an ejector type of thruster is interesting as it seems tooffer a solution to the problem of combining a high-pressure cargo pump withan effective low-speed tran sv er se jet. However, I would like to draw your a t tention to the fact that it might occasionally be desirable to have the possibilityof using the thrusters and the cargo pumps simultaneously, for instance whenloading or unloading at mooring, for trimming purposes when getting out of harbors, et c. Obviously the re are sever al other aspect s, mainly of a pra cti cal nature, which should be regarded when considering a cargo-pump-d riven th rust er .With respect to the human factor, one might expect for instance an increasedri sk of accidental discharge of cargo oil. In my opinion, a cargo-pump thruster

should not be considered unless it shows marked economical benefits over conventional th ru st er s. The present paper does not give any cle ar indication ofthi s point. The power figures given re pr es en t only the jet power. I would liketo know if the authors have made any estimates of frictional and bend losses inthe rather extensive pipeline they are proposing, and if a normally sized pumpequipment is capable of producing a sufficient thr us t. It would al so be in te re st ing to see a compar ison between the cos ts of a cargo- pump th rust er includingvalves, nozzles, pipes, governors, etc., and a corresponding conventionalthruster.

Among the possible merits of a thruster steering system that the authorsmention ar e the reduced tug ass ist anc e cost s. To il lustr ate the ra te of thesesavings I would like to submit some statistical data given to us by a well-knownshipping company in Europe. The figures apply to two 12,000-TDW si st er car goliners which have been in service between Europe and the Far East for abouttwo years . One of them is equipped with a 600-hp bow th rust er plus acontrollable-pitch main propeller; the other has a conventional fixed-pitchpropel ler . It appears that for the ship with bow thruster , tug cos ts are reducedby about 80 per cen t compared with her s is te r ship without bow th ru st er . Forsome of the ports along the trade the following numbers of tugs are needed including both arrival and departure:

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Manen, Oosterveld and Witte

With WithoutBow Thruste r Bow Thr ust er

Aden 1 2HamburgRottendam

21

55

AntwerpGottenburgOslo

300

841

Aarh us, Denmark 1 4Copenhagen 0 2

The above figures have also been verified by several later installations.

An important matter which has not been treated by the authors is the stopping charact er is ti cs of these very large ves se ls . With a conventional ste rn a r rangement and a free fixed-pitch propeller, the stopping time for tankers between 50 and 100,000 TDW is normal ly between 8 and 10 minutes correspondingto a stopping dis tance of 7 to 9000 feet . One may expect these figures to be evenhigher when using a nozzle propell er as proposed by the authors. A conside rable reduction in stopping time and distance may be achieved by using a controllable pitch propeller . I would like to finish with a few figures obtained withsome of our own propellers.

For a 50,000-TDW twin-screw tanke r with an output of 8900 HP each shaftat 115 RPM and equipped with two 4-bladed 17-1/2-foo t cont roll able-pitch p ro

pe ll er s, a stopping time of 4.5 minu tes from full speed, 15.75 knots, was r e ceived. For a 72,000-TDW single-screw ore ca r r ie r with an output 17,600 HPat 115 RPM, fitted with a 4-bladed 21-1/2- foot CP propel ler , a stopping time of5 minutes was obtained from full speed, 16.75 knots.

Compared with similar ships with normal fixed-pitch propellers, this implies a reduction in the order of 50 percent.

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REPLY TO DISCUSSION

J. D. van Manen, M. W. C. Oosterveld, and J. H. Witte

We wish to thank Mess rs . Lackenby, Stron-Tejsen, Bindel, and Bj or hedenfor their interesting comments.

In regard to Mr. Lackenby's question about the improvement in SHP withthe Hogner-type stern and nozzle and the improvement in SHP with an extremebulbous bow, it must be remarked that our experiments were performed withships without bulbous bows. It was found that the Hogner s te rn with nozzle

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Manoeuvrabi l i ty and Prop uls ion of Very Large Tan ker s

requires about 6 percent less SHP in the loaded and 9 percent less SHP in thelight condition. The applica tion of an ex tr em e bulbous bow for thi s type of shipwill possibly lead to a 15-percent SHP reduction in the light and no reduction in

the loaded condition. The time this type of ship will sail loaded and in bal lastwill be about equal, so the improvements in SHP with the Hogner stern and nozzle fitted and with an ex treme bulbous bow are of the sam e magnitude. At th ismoment we do not have experience with respect to the resistance and the required shaft horsepower of ships with both an extreme bulbous bow and a Hognerstern with nozzle fitted.

We agree with Mr. Strom-Tejsen that the jet system according to the paperwould require excessive power in order to produce at service speed forcescomparable to those obtained by a conventional rudder, due to the relativelysmall nozzle di ameter s. The compar ison between the EHP of the model withthe Hogner-type and the conventional stern has been based on experiments with

out jet tunnels installed in the model. The increase in re sist ance due to thetunnel openings (for these small tunnel diameters less than 0.5-1.0 percent) wasnot taken into account.

We agree with Mr. Bindel that special attention must be paid to the problemof course stabilit y. Some re ma rk s can be made alr eady. Fi rs tly, we ar e awarethat the Hogner st ern will not affect favorably the course stabi lity. On the o therhand we expect that the nozzle itself will form a positive element with respectto course stabi lity. Finally the tai l piece will be an addition and forms a meansfor further affecting (by increasing the lateral area) the course stability of thetanker.

In regard to Mr. Bjorheden's questions, it must be remarked that for thedetermination of the propeller, hull, and relative rotative efficiencies, the thrustof both screw and nozzle must be meas ured . Only the thrust of the screw wasmeasured here.

The data given by Mr. Bjorheden with respect to the reduction in tug assistance costs due to the application of bow thrusters and his remarks with respectto the stopping cha racte ri st ics of large vess el s ar e very interesting. Recently,investigations were performed at the N.S.M.B. with respect to the stoppingabilities of a 100,000-TDW single-screw turbine tanker with an output of 28,000HP at 85 RPM and equipped with different propeller types. The stopping ab il ities have been derived from model test results by means of a quasi-stationarymethod. Stopping ti me s of about 32, 28, 23, and 18 min from full speed of 16knots were found for, respectively, a conventional screw, a screw and nozzlesystem, a contra-rotating propeller system, and a controllable-pitch propeller.The head reaches wer e, respect ive ly, 5100, 4800, 4450, and 2900 me te rs . Foran extensive discussion of this matter, we refer to the original paper of Hooftand van Manen, to be re ad at the Spring Meeting 1967 of the RINA. Fr om theseresults it can be seen that the stopping characteristics of the ducted propellerare better than those of the conventional screw.

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Documents

1966-Vanmanen-research on the Manoeuvrability and Propulsion of Very Large Tankers