Propellers with restricted diameter

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MARINTEKNISKA INSTITUTET SSPA

SWEDISH MARITIME RESEARCH CENTRE SSPA

GÖTEBORG

PUBLICATION NO 99 1984

DESIGN PRINCIPLES, EFFICIENCY AND

CAVITATION PROPERTIES

by

ERIC BJÄRNE

Paper partly presented at the

Symposium on Advances in Propeller Research and Design Gdansk. February 1981

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Distributed by:

MARINTEKNISKA INSTITUTET SSPA P.O. Box 24001

S-400 22 Göteborg. Sweden ISBN 91-86532-02-2

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i INTRODUCTION

Usually the pitch and diameter of a propeller is estimated with the aid of propeller-charts based on systematic tests with pro-pellers with identical blade shape and section profiles. If the propeller dimensions diverge from the optimal ones the properties

will be inferior to those possible due to inclined flow to the

blade profiles.

If the propeller is calculated according to the circulation theory assuming profiles with shock-free entrance, significant profile mean line cambers, pitches and blade widths are obtained

for each combination propeller loading-diameter i e the pro-peller is fully adapted to the conditions.

It is thus possible to reduce the propeller diameter within cer-tain limits keeping the machinery (i e power and rates of revol-utions) unchanged without too much increase in propeller power.

The influence of such propeller diameter reductions have been investigated for a 60 000 TDW tanker [U and a large

Ro-Ro-vessel.

Model tests, including open water and self propulsion tests, have been carried out in the towing tank. Cavitation tests, in-cluding hull pressure fluctuation measurements, have also been

performed.

2 THEORETICAL BACKGROUND

The lift of a propeller profile can be produced by an angle of attack and/or mean line camber. For propellers with different pitch ratios, but with the same section profiles and blade con-tour, the profile characteristics are identical and the open water performance at any pitch ratio can be calculated by the Lerbs method for equivalent profile [2].

The results of such calculations are presented in the propeller

charts, Fig 1. These figures are valid for propellers with

sym-metrical (non-cambered) sections, which means that the lift co-efficient at zero angle of attack, CLO1 is zero, and also for

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propellers with more cambered profiles, CLO = 0.2 and 0.4

respectively.

According to Reference [3] the drag coefficient is not expected to be significantly influenced by the camber within the limits

described here.

From the propeller charts, Fig 1, it can easily be found that for a certain propeller loading the optimal diameter for a pro-peller with cambered profiles is smaller than that for a propel-ler with symmetrical profiles.

If the propeller diameter is reduced, keeping the profiles

un-changed, the loss in efficiency will be greater than if the

camber is increased correspondingly, Fig 2.

3 PROPELLER DESIGN

Two propellers were designed for each type of vessel according to the circulation theory under the condition that the profile

angles of attack were zero. The mean value (K) method for

lifting-line calculations described in Reference [4] was applied. Corrections of the calculated pitch and camber were calculated by the lifting-surface method described in Reference [5]. The design data are given in Table i. The wake and circulation dis-tributions were the same for both the propellers, Fig 3. The main data for the propellers appear from Fig 4.

4 TESTS IN TOWING TANK

The open water tests were carried out according to the SSPA standard procedure. The rate of revolutions was thus kept con-stant, whilst the speed of advance was varied so that the range of advance coefficients of current interest was covered by measuring points. Propeller thrust and torque were measured and transformed to the non-dimensional coefficients KT and KQI Fig 5.

The self-propulsion tests were carried out using ship models of a 60 000 TDW tanker and a 30 000 TDW Ro-Ro-vessel respectively. The main data of the models are given in Table 2. The continental method using a friction allowance calculated with the ITTC 57

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friction line and with a ship roughness allowance, ACF, of 0.0004 was applied at the tests.

The full scale predìction has been extrapolated from the model

test results by the use of the method recontxnended by ITTC 78

[61. The predicted values are given as Tables 3 and 4 and

dia-grams , Figs 6 and 7.

5 ANALYSIS OF THE TEST RESULTS

The open water test results were analysed according to the method with an equivalent profile proposed by Lerbs [2]. It was hereby found that for both the small propeller models extremenly high values of the minimum drag coefficient were obtained, Fig 8. As the investigation was based on existing "optimum" propellers, the diameter of the small propellers was restricted and thus the

Reynold's number was low. The cambered profiles are according

to Fig 9 (see also Reference [3]) however more sensitive to the

Reynold's number and to obtain over-critical conditions Reynold's

numbers up to iO7 are required, which is difficult even with fairly large propeller models. This sensitivity for Reynold's number is verified by tests with another propeller model, P1816, which also had highly cambered profiles.

On the contrary to the above, turbulent conditions seemed to be valid at the self-propulsion tests. This was indicated by the extremely high values of relative - rotative efficiency when using uncorrected open water characteristics for the small pro-pellers. If these open water characteristics were modified to the minimum drag coefficient of the large propeller both wake fraction and relative - rotative efficiency became reasonable. The final results have been based on the original open water characteristics for the large propeller models and character-istics modified to identical minimum drag coefficient for the small propeller models, Fig 5.

Open water charts for propellers with different profile camber (i e lift coefficient, CL0) have been calculated according to Reference [2], Fig 1. Based on these charts and the interaction factors obtained from the self-propulsion tests the power has been predicted for optimal conditions at a certain speed for the

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two projects concerned, Figs lo and il. These diagrams indicate optimal geometry of the propellers and also how much required power will increase at off-design conditions.

6 TESTS IN SSPA CAVITATION TUNNEL NO 1

Tests were carried out with the respective propeller model

mounted behind an existing dummy model. In order to simulate

the velocity distribution behind the ship model, wake producing nets were applied to the dummy model. Measurements of the local velocities with Prandtl tubes indicated good agreement between

the ship and the dummy model concerning the velocity distribu-tion in the propeller field.

The mean speed of advance at a certain free water speed in the tunnel was determined on the basis of KT-identity: open water tests-present tests at atmospheric pressure over a range of advance coefficients. The velocities thus obtained were used at the calculations of required pressure and rate of revolutions atthe cavitation tests.

The cavitation was studied at propeller loadings according to Table 5. The trial cases for the "optimal" propellers were based on the predictions obtained from self propulsion tests with the original ship model. At service the required power was assumed

to increase by 15 % and the wake fraction by 0.05 in relation

to trial. The RPM was then calculated with the aid of the open water characteristics. The relation between the propellers with regard to ship speed and wake fraction was determined from the self propulsion test results.

Sheet cavitation combined wïth tip vortex cavitation was at the tanker cases noticed outside radii 0.8 R in the low speed region behind the stern, Figs 11 and 12. The cavitation was somewhat more stable on the smaller propeller model. No bubble foaming of face cavitation was indicated, but the margin against incipi-ent face cavitation was fairly small, Fig 13.

In the Ro-Ro-vessel cases the back sheet cavitation was some-what more extensive (0.6 R to tip for the larger propeller and

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cavitation was however somewhat more durative for the smaller than for the larger propeller model. Some face cavitation was observed in the trial cases and on the smaller propeller also at service condition. Neither in the Ro-Ro-vessel cases was foaming or bubble cavitation observed. Incipient face cavitation conditions häve been determined, Fig 13.

The propeller induced pressure fluctuation was measured in two points on the centre line of the dummy model above and in front

of the propeller model, Fig 14. A detailed description of the measuring devices and methods is given in Reference [7]. The

signals were recorded on a loop oscillograph (blade frequency and resultant signals) and on a multi-channel tape recorder. The oscillograph recordings are presented in Fig 15, indicating a

reduction of 4 5-60 % of the blade frequency pressure pulses above

the propeller model with the smaller diameter. The tape record-ings for 100 consecutive revolutions have been digitalized and

analysed according to Fourier. The values for the revolutions

with the 5 % highest amplitudes are compared for the two

propel-ler couples in Fig 16, verifying the above relations.

7 CONCLUSIONS OF THE TEST RESULTS

According to the test results it is for the tanker case possible to reduce the propeller diameter by 20 % at constant RPM without

losing more than about 4 % power in efficiency.

In the Ro-Ro-vessel case the corresponding loss of efficiency is 10 % due to a smaller wake gradient.

The cavitation properties are not deteriorated by the diameter reduction. The back cavitatation is somewhat more stable for the

smaller propeller. Due to the increased tip clearance the small propeller gives a reduction of 45-60 % of the blade frequency pressure pulses above the propeller and thus also the vibration. This reduction is less significant in front of the propeller.

The present results are based on a restricted meterial but is expected to give some guidance at the choice of propeller

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In this work only the efficiency and cavitation properties for propellers with reduced diameter have been investigated. Also other properties as for instance those at manoeuvring must be

investigated.

The thoughts described here are not new. Already in 1955 Dr Burrill penetrated the possibilities to reduce the propeller diameter without significant efficiency loss [8]. So far no attempt to apply the ideas to full scale projects have been made, partly due to the difficulties to prove the advantages of reduced diameter propellers with model tests.

8 ACKNOWLEDGEMENT

The author expresses his sincere thanks to the Swedish Board for Technical Development for the financial support of this work and to those members of the SSPA staff involved in the

analysis of the material.

9 NOMENCLATURE

A = As index: for trailing edge

AD = Developed blade area measured to hub

A0 = Propeller disc area (= -i---)

C = Chord length

C075 = Chord length at x = 0.75

CL = Lift coefficient, here for x = 0.75 (=L/p/2

C075VQ275

CD = Drag coefficient, here for x = 0.75 (-D/p/2 C075V75)

D = Propeller diameter, drag force

e = Vapour pressure of water

F = Profile mean line camber, as index: for leading edge

g = Acceleration due to gravity (9.81 m/s2)

H = Position of propeller shaft centre line below water

surface

VAT resp VAQ)

J = Advance ratio _Dn

K = Pressure flucutation coefficient (

22)

p pD2n

K0 = Torque coefficient ( pi-i2)' behind conditions

K00 = Ditto, open water

T

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L = Lift force

m = As index: model

n = Rate of revolutions

p0

= Static pressure at the propeller shaft centre line

P = Propeller pitch

= Delivered shaft power

= Pressure fluctuation single amplitude measured on

the hull

Q = Propeller torque

R = Propeller radius (= ), resistance

co.75 R = Reynolds number (

O75

n V r = Radius s = As index: ship

t = Water temperature, thrust deduction factor (=

T = Propeller thrust, as index: trial or based on thrust

identity

V = Speed

VAT = Mean speed of advance determined on thrust identity

VAQ = Mean speed of advance determined on torque identity

V075

= Resultant velocity at X = 0.75

V_VAT

WT = Mean wake fraction = _V

V-V

wQ = Mean wake fraction

(-X = Radius ratio (r nR)

z = Number of blades

= Profile angle of attack

= Open water efficiency (=

r-y

Qo

1.-t

= Hull

efficiency =

= Relative rotative efficiency ₍₌

j2)

Q

n = Quasi-propulsive coefficient (= _no _nR _nH)

= Kinematic viscosity of water

p = Mass density of water

p -e o

o = Cavitation number

p72 VAT2)

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10 REFERENCES

Ejärne E

Propellers with Restricted Diameter - Design Principles, Efficiency and Cavitation Properties.

Symposium on Advances in Propeller Research and Design. Gdansk 1981, p35-53

Lerbs H W

On the Effect of Scale and Roughness on Free Running

Pro-pellers.

Journ. Am. Soc. Nay. Eng. 63(1951):i, p58-94

Abbott I H, von Doenhoff A E Theory of Wing Sections.

Mc Graw Hill Book Co mc, New York 1949, 693p

Johnsson C-A

An Examination of Some Theoretical Propeller Design Methods. SSPA Publication No 50, 1962, 52p

Johnsson C-A

Applications and Experimental Verifications of a Theoretical Propeller Design Method.

(in Swedish). SSPA General Report Noii, 1965, 66p

Lindgren H, Dyne G

Ship Performance Prediction. SSPA Publication No 85, 1980, 22p

Lindgren H, Bjärne E

Ten Years of Research in the SSPA Large Cavitation Tunnel. Model Experiments as an Aid to Advanced Propulsion,

Stone Manganese Marine/Newcastle University Conference, 1979, Paper No 7, also SSPA Publication No 86, 1980, 86p

Burrill L C

The Optimum Diameter of Marine Propellers: A New Design

Approach.

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Table 1. Design and propeller data

DESIGN DATA Tanker Ro-Ro-Ship

P1801 P1813 P1758 P1890

Ship speed, V, knots 16.0 15.5 18.9 18.6

Rate of revolutions 107 107 127.3 127.3

Wake fraction, WT 0.41 0.44 0.25 0.32

Propeller thrust, T, KN 1240 1160 1079 1045

Estimated shaft power,

D' MW 11.6 11.6 12.6 12.6 PROPELLER DATA Propeller diameter, D, m 7.1 5.68 5.9 4.72 Hub diameter, DH m 1.14 1.14 1.17 1.17 Pitch at 0.7R, P07, m 4.67 5.65 5.192 6.070 Mean pitch, M' m 4.54 5.34 5.052 5.193

Blade area ratio, AD/AO 0.57 0.61 0.65 0.77

Rake, O, degrees O O O O

Estimated weight, G, tons 22 14 20 15.7

Estimated moment of inertia

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(+ forward - aft)

Table 2. Main data of ships corr to models tested

Tanker Ro- Ro- Ship

Ship model in 2079-B 2148-A

Draught ni in 12.2 9.15 T in stern 12.2 9.15 Length in L in pp 236.0 212.3

L1

in 241.0 208.0 Beam (WL), B in 41.6 32.6 Displacement, V, m3 98470 41752

Wetted surface, s, in2 13370 8362

CB = V/(L BT) 0.822 0.659 L /V1 wl 5.219 5.995 C = S/IL V 2.773 2.808 S _pp B/T 3.410 3.563 L /B wi 5.793 6.380 LCB from L /2, t/L , pp pp 2.69 -2.56

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Ship values Length : 236.0 m Length WL : 241.0 m Draught FWD 12.20 m Draught MT : 12.20 m Beam : 41.6 m Wetted surface: 13 370 m2 Displacement : 98 470 m3 Form factor : 0.240

Self-propulsion tests, propeller model P1801

Ship Thrust Wake Prop Hull Rel-rot QPC

speed deduct fract eff eff eff

factor

vs t _WT fo _H _R

knots - - -

-Table 3. Predicted full scale values of RPM, power, efficiency and interaction factors, tables, tanker

13 0.174 0.341 0.582 1.253 0.969 0.707 13.5 0.177 0.336 0.582 1.239 0.957 0.690 14 0.180 0.332 0.582 1.227 0.951 0.679 14.5 0.194 0.337 0.581 1.215 0.950 0.671 15.0 0.207 0.341 0.581 1.203 0.951 0.665 15.5 0.203 0.340 0.581 1.207 0.951 0.667 16 0.182 0.330 0.581 1.222 0.952 0.676

factor V t _WT fo H R knots - - - -13 0.175 0.369 0.521 1.307 0.957 0.652 13.5 0.170 0.357 0.529 1.291 0.962 0.657 14 0.173 0.357 0.527 1.286 0.964 0.653 14.5 0.181 0.359 0.523 1.278 0.960 0.642 15 0.190 0.364 0.518 1.273 0.957 0.631 15.5 0.177 0.358 0.518 1.282 0.951 0.632 16 0.145 0.340 0.526 1.295 0.953 0.649

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Ship values:

factor V s knots t Length Length WL : 208 m Draught FWD : 9.15 in Draught AFT : 9.15 m Beam : 32.6 in

Wetted surface: 8 362 in2

Displacement : 41 750 m3

Form factor 0.206

WT rio

H

Table 4. Predicted full scale values of RPM, power, efficiencies and interaction factors, tables, Ro-Ro-vessel

n 16 0.175 0.288 0.610 1.159 0.999 0.706 17 0.169 0.282 0.609 1.156 0.992 0.698 18 0.171 0.274 0.613 1.141 0.982 0.687 18.5 0.170 0.272 0.611 1.141 0.988 0.689 19 0.170 0.274 0.609 1.142 0.986 0.686 19.5 0.169 0.276 0.606 1.147 0.986 0.685 20 0.170 0.278 0.602 1.150 0.986 0.683

factor V t no 11H R knots - - - -16 0.194

0.38

0.531 1.182 0.978 0.614 17 0.169 0.304 0.538 1.193 0.973 0.624 18 0.171 0.291 0.543 1.170 0.969 0.616 18.5 0.158 0.281 0.548 1.171 0.970 0.622 19 0.150 0.269 0.554 1.163 0.970 0.625 19.5 0.146 0.268 0.553 1.166 0.970 0.625 20 0.144 0.269 0.550 1.172 0.970 0.625

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Table 5. Loading cases at cavitation tests

P1801

Full

load Service 0.345 19.0 15.3 0.46 103.9

Ballast Service 0.344 15.2 15.5 0.47 103.6

P1813 Full load Service 0.408 21.1 15.1 0.47 106.7

Ballast Service 0.410 16.2 15.3 0.48 105.4

Ro-Ro-Ship

P1758

Full

load Trial 0.582 5.70 18.9 0.25 127.4

Full

load Service 0.535 7.06 18.2 0.30 124.6

P1890

Full

load Trial 0.695 6.08 18.3 0.25 129.2

Full

load Service 0.633 7.55 17.6 0.30 127.3

Loading case Rates

Prop Adv Cay Ship Wake of

model Load Cond coef f No speed fract revs

J

O V5 WT N

-

_-

knots - RPM

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1.3 1.2 2 1.1 o £1.0 Q- 0.9 14 08 0.7 0.6 O.5 1.3 .2 .2 1.1 o £10 °- 0.9 0.8 07 0.6 o 0.8 0.7 0.6 050 -'o o o O

I

va

NAIWDi

Wh4VIIDAY&VA

Ar41r&ir

05 1 0 Proper

253,0

1H14VAAYÀ VAV&V

'tIa!JII_ '

_{tWDiA4WA A}

05 10 15 20 25

Prope ter toading

IIIIfhW!L&i1PÀU

Y1!IIIIJILIiV&

iii

ri va vii

0.5 1.0 1.5 2. Propeller bingV3f Q1.3 1.2 .21.1 a 1.0 Q- 0.9 0.8 0.7 0.6 o.5o Ql.3 .2 a-01.1 o £1.0 Q- 0.9 0.8 07 06 0.50 1.3 1. 2 .9 1.1 o 1.0 0.9 0.8 0.7 0.6 0.5

I WIIrU1MVDNDAIIaV

\muuivirr

74 Wì. YÀIPÀUr

riiir

r

w#ixii a

CLO=GO I 0.5 1.0 15 Propett« -,_-oo O o o

ziwawa»

«4WiWj1AWA À

wJ, iiIra

t' ÁMirvd

0.5 1.0 1.5 20 2 PrOpe ter t0Cing,

tk\ I VIM WA AWIWAW

tJ/L&iIINIAVMa!]

LIV Villi!

iviîiiaw

TgVALWA

I7ilIffffi"

IIÍ

) 05 1.0 1.5 20 25 Propetlec tolIng

Fig 1. Propeller charts MS4-060, different profile cambers

000 O

O o O

o = O O

o

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Propeller efficiency, 0,55 0.50 0,45 0.40

\

0,4 0,5 Advance coefficient, J

Fig 2. Propeller efficiency at different diameter and profile camber

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16 1.5 L L t-Q 1.0 a 'C o

Prop mod P1801. optimal, w1

0.l

Prop mod P1758, optimat, w1 r 0.25

- - - Prop mod P1890. D = O'800pt . WI r 0.32

Fig 3. Wake and circulation distributions

C 0.6 .2 t-) o 0.5 a

04

0.3 0.2 0.1 0.2 0.3 0.1. 0.5 0.6 0.7 0.8 0.9 1M Radius ratio. Xr nR 'a -a C o a

0-

0 u 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 Radius ratio. X nR 0.6 0.5 u o 0.4 o 0.3 0.2 0.1 1.5 e L L t-. o C o 1.0 o a 'C o 0.5 a C o a -X U

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2

2 0.8

0.7

Mean pitch

Radius rollo, o.r/R

.0 0 0 OES 08 0,7 0.6 0.5 0.4 0,3 o 2.0 3.5 10 05 0 0.5 1.0

Blade width ratio rs _D

A

0.2 03 04 0.5 0,6 0.7 08 0.9 I .D

Rodios ratio, e nR

Prop nod P1801, optIonal

- -- Prop rood P1813, 0.0 IDopt

0.9 o- 150-E E Radius ratio, o t/R 1.6 10 05 0 0.5 resp C0

Blodn width talio _o

A0/Ao' 'y

/

I / / 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 Radiusrotto, o. nR

Prop rood P1758. optimal Prop rood P1890, 0=0.10001

/

0 1.5 ' 100o / 7 I

\

02 03 4 05 0:6 0:7 Ro ral:9n=r/R0

:

0I0,,

0.4 05 0.6 0.7 Rud

ratI.r/0

Fig 4. Geometric properties of the propellers

17 1.0- _{MOan pitch} 50 2 E loo 50 o

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0 K1 lOO K Ro-Ro-ship Adco-.ce oeftir.ent IO <i lOO K O 0.1 02 03 04 05 06 07 08 0.9 IO It Advonc, cotti J O O DI IO KT 00 K9 13 12 II lo 9 8 7 5 Fig 5.

Open water characteristics

05 06 0.7 0.8 09 10 Adoanot cotti J

Propeller model P1890 lOcOE800cpt}

K0 D 01 02 03 04 0.5 06 07 08 09 lO II 12 13 Advonc. cotti. J e -t 'opt A

Seit pnopotnoo poed

Ib

t.,

tL

lbohnd 0041010v

cocd,t,oc.l U4

opellet model P1813 lO 8nDoptl

Seit peopc( oe p0.10 Coen.cted f to coen -(behind tornei oo,dtiov KQ co voter tent. UCorrected Atoopoew.ton for CD tent, pt

u...

-, -tope et e 1=1 o' 02 03 IO K1 lOO K Tanker 4 3 7 6 s L 3 2 o

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1 .0 13 14 15 16 -______

Prop mod P1801. optimal

-

--Prop mod P1813, D r 0.8Dot

I-13

14

15

16

Ship speed. V0 . knots

Ship speed. V5 , knots

Fig 6.

Propulsive factors, predicted RPM and power, tanker

140 o-130 ° z 120 o e IA O1va 100 90 ne 0.9

Thrust deduct factor, t Wake fraction (thrust identity), w1 Propulsive efficiency, i

(QPC)

0.8

Open water efficiency. 1%

20

Relative - rotative efficiency.

R 0.7 11 o 0.6 15 o 'Vio 0.5 U) 0.4 w1 10 0.3 0.2

---t 5 0.1 O

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0-'

-Prop mod P1758, optimal Prop mod P1890, D0.8DOpt

Fig 7.

Propulsive factors, predicted RPM and power,

RoRoShip

o -16 17 18 19 20 16 17 18 19 20 Ship speed, V5 . knots

Ship speed. V, knots

1.0 0.9

Thrust deduct factor,

t

Wake fraction (thrust identity). w1 Propulsive efficiency, 'I) (QPC)

0.8

Open water efficiency, "lo

20 0.7 0.6 Relative - rotative efficiency, 'nR o _o- a' oo.)5 1) 'no o In 0.5 0.4 10 0.3

-- WI

0.2 5 0.1 150 140 z 130 o ai a 120 110 100

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Lift coefficient CL

Drag coefficient loco

Angle of attack, a, degrees

Fig 8. CharacterisheS of equivalent profiles (X = 0.75)

P181: o.:D01

Ò

PI'°U

3 01 813 C.rrect.

....

_a

18riu

758

u....

P1890

--N'.

r

3 -2 - 0 2 3 h 5 -3

-2

- O 2 3 4 5 6

Angle of attack, a, degrees Litt coefficient CL

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0.015 0.010 0.005 s P1890 S P1813 \ NACA 653-1.18 NACA OOI2

\

N

P1758

-I P1801 P1816 2CF (QCC to ITTC) 5.0 5.5 6.0 6.5 7.0 Log R

Fig 9. Minimum drag coefficients as function of Reynolds number

Cn

Min

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E _o E o o ea s

80

Ro-Ro-vessel

Trial oint

P1890

Ship speed. Vs 8.5 knots

lOO

20

140

160

180

Rates of revolutions. N. RPM Ship speed ,Vs t5 knots

2MW /125MW _{/ 13MW} I5MW 00 20 40 60 80 Rates of revolutions, N RPM

Diameter for given RPM

E b Eo ea 6

RPM for given diameter

Tanker 60 80 II MW 11.5 MW f175 MW 12MW Ship speed, V5 15.3 knots lOO 120 loO Rates of revolutions N, RPM Fig 10.

Optimal diameter for given RPM and optimal RPM for given diameter

to E o 149 Eo 148 a 6 5 4 IO-E o .89 E a 0 148 6 S 60 80 tOO 120 40 60 Rates of revolutions, N RPM

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PropeLler model P1801, J0.34S o=19

500 _30° 20° 10° 350°

Propeller modefl P1813, J=0.408 G=21.1 Tanker

Fully Loaded ship. service cond.

Bu (tasted ship. service cond.

Propeller modeL P1801, J=0.344 O15.2

500 300 _20° 100 _350°

Propeller modell P1813 J0.4ìO O16.8

Fig lia. Cavitation patterns. Tanker

500 _30° _20° 100 _350°

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RoRoship

Fully Loaded ship, trial cond.

Propeller model P1758. J =0.582 0=5.70

500 300 _20° 100 3500

Propeller model P1890, J=0.695 0=6.08 Fully loaded ship. service cand.

Propeller model P1758, J=0.535 07.06

500 300 _20° _10° _350°

Propeller model P1890, J=0.633 0=7.55

Fig lib. Cavitation patterns. Ro-Ro-Ship

50° 300 20° 100 _350°

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Tan k er

Propeller model P1801

O Angle, P. degrees

Fully Loaded ship , service cond

PropeLLer modeL P1813

O Angle, P, degrees

180 180

J=0 345, 0=19 J =0.408, c=21.i

Ballast ship, service cond.

180 180

J0.344, 0=15.2 J=0.410, 0=16.8

Fig 12a. The maximal radial extension of the cavitation

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Ro-Ro-ship

Fully loaded ship, trial cand.

Propeller model P1758 Propeller model P1890

0 Angle.P.degrees O Angle, 0, degrees

180 180

J=0.582, cJ7.55 J:0.695 C 6.08

Fully lauded ship, service cand.

180 180

J0535, G7.06

J:0.633, cï755 Fig 12b. The maximal radial extension of the cavitation

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Cavitotion number. o CavItation number, o

Full load

20-service

Prop mod P1801, optimal

Prop mod P1758, optimal

0.3 0.4 0.5 0.6 Advance cocU. J Ballast service s

Prop mod P1813. DO.8D0t

0.3 0.4 0.5 0.6 Advance coefl. J Cavitation number, o 9 8 7 6 5 0.5 0.6 0.7 0.8 Advance cactI J Prop model P1890. D0.8D0pt Cavitation number, O 0.5 0.6 0.7 0.8 Advance coeU. J Fig 13.

Incipient and decedent face cavitation

conditions

20 18 16 14 12

8

Full toad. Service

7

.

Foce cay. region

6

Full load, trial

5

(31)

I I _I / L) 76.0

N

- Dimmy model 1715 -B2A

Shuip model 2079-B

N

/

-'4

P1890 27.32

Dummy model 1715-ASB Shuip model 2148-A

N

\

/

_/

/

_z

75

N

Fig 14a.

Position of propeller models and

Fig 14b.

Position of propeller models and

pressure transducers, tanker

pressure transducers, Ro-Ro-Ship

Tanker

(32)

2p xX Pa (daible woplitude)

70 T 70

Full load service cand.

60 I 60

SO

20 I r 1 20

-c

15 20

Ship speed, V5 knots

10 15 20

Ship speed, Vs,knOts

Tanker

Wake Total Blade Wake Total 8le

tract ampI freg. tract mpl. freq.

0.46 -

Mv. coeff. J 0.35 Prop, moO P1801 0.47-- Ads. coeff, J Ø.34.4 Prop mod P1801

0.48 ---

JO.408 P1813 J0i.10 P1813

Ro-Ro- ship

Wake Total Blade Wake Total Blade

tract ampi. freq. fract amp) freq.

025 -

Mv. coeff. i0.582 Prop nmd P1758 0.30 -Ads. coeff. J=0.535 Prop mod P1758

0.25 ---

- J=O.595 * - P1890

0.30 ---J0.633

- - P1890

2p xX Pa (double ämplidude I

Ballast service cand.

15 20

Ship speed, V5, knots

10 15 20

Ship speedy5, knots

Fig 15. Estimated full scale pressure pulses. Transducer F

Full load trial cand. Full load service cand.

60

Thp xX Pa (double amplitude) 2pxX Pa(double amplitude)

(33)

Kp 0.05 O Kp 0.30 0.20 0.10 Propeller model P1801 Adu. coeff, J0.36S 0.05 Propellermodel P1813 Mv. coeff J 0.408 <p _<p e 0.05 0.05 íd Propeller modell P1758 Mv. coeff. J0.S82 Kp 0.30 Propeller modell P1890 Adv. coeff. JrO.695

<p 0.30 0.10 i 0.345

012345

Number of o der 5

012

J 0.410 Ro-Ro-vesset J r,535 J 0.633

45

Kp 030 0.20 0.20 0.10

JJAtmospheric pressure Covitating condition

Fig 16. Results of pressure fluctuation measurements, pressure pulse coefficient, Kp

Propeller modef P1801 Adv. coeff J0345 Propeller modell P 1813 #dv coeff. J0 608 Kp 0 05

012345 "Rl 2345

Number of orde Propeller modell P1758 Ad'.comff. .lO582 Kp 0.30 Propellermodell P1890 Ado. coeff. J0.b95 <p 0.30 0.20 0.10 J 0.410 J0.53S J0.633