Ship research activities in the Netherlands

May 12, 1971 REPORT No. M288083/259839

SHIP RESEARCH ACTIVITIES IN THE NETHERLANDS

JOINTLY SPONSORED BY THE ROYAL NETHERLANDS NAVY,

THE MINISTRY OF ECONOMIC AFFAIRS AND THE NETHERLANDS INDUSTRY.

by

OOSTERVELD and P. van OOSSANEN

MINISTERIE VAN DEFENSIE (MARINE) HOOFDAFDELING MATERIEEL

BUREAU SCHEEPSBOUW

TORENSTRAAT 172 's-GRAVENHAGE

NEDERLAND

This paper presents in digested form some important results of research activities carried out by the Netherlands Ship Model Basin during the period 1963 to 1969. These activities mainly refer to investigations of various high-speed ship propulsion problems. They were sponsored by the shipyards and engineer-ing companies, participatengineer-ing in the Netherlands United Ship-building Bureaux Ltd. (now Rijn-Schelde Group), Lips Propeller Works and by the Royal Netherlands Navy.

The co-operation of industry and government in their sponsor-ship of this extensive research program became a fact when it was realized that probably in the future the conventional ship propeller would no longer suffice due to the trend of increasing

ship speed. Even though the demands on their performance _are

different, both merchant ships and naval vessels are subject to this trend.

Due to the importance placed on this research _{program by} the Royal Netherlands Navy, the investigations into the propul-sion problems of naval ships were somewhat accentuated. As such, the Royal Netherlands Navy, on behalf of the sponsors, became the chief accomplice to the Netherlands Ship Model Basin in performing these investigations.

Finally, the many investigations carried out have not only led to a vast amount of knowledge applicable to present and future propulsion problems, but also to basic knowledge, inherently present in various other branches of naval architecture.

Rear Admiral Mr. Jr. P. P. van de Vijver, Royal Netherlands Navy.

SHIP RESEARCH ACTIVITIES IN THE NETHERLANDS

JOINTLY SPONSORED BY THE ROYAL NETHERLANDS NAVY,

THE MINISTRY OF ECONOMIC AFFAIRS AND THE NETHERLANDS INDUSTRY.

by M.W.0 Oosterveld.) and P. van Oossanen.)

Summary.

This paper presents results of an extensive ship research program carried out at the Netherlands

Ship Model Basin during the period from 1963 to 1969. Various high-speed propulsion problems_are

dealt with. Results are given of fundamental cavitation research into the mechanism of _cavitation inception and related scale effects. A computer program for the unsteady lifting surface theory for ship screws was developed and some results obtained herewith are given. Various unconventional propulsion devices were theoretically and experimentally investigated, and in particular _some re-sults of these investigations on ducted propellers, contra-rotating propellers, low-noise propellers and the water-air ramjet are presented.

Research into various interaction effects was carried out and in this regard the results of com-parative tests carried out with a model of a tanker and a model of_{a cargo-liner, each equipped} with successively contra-rotating propellers and conventional propellers _{are given and discussed.} Finally, various ship motion studies were performed, and the results of theoretical and

ex-perimental work on the motions of a ship in a seaway, slamming phenomena. _{and on the added}

resistance in a seaway are presented. I. Introduction.

During the past years the trend of most ship designs has been towards higher speed and. therefore, high powered ships. As a result the problems of propeller cavitation, propeller in-duced vibration, and especially in the case of naval ships the radiation of propeller noise, be-came matters of great concern.

The main requirements for a ship propeller are:

high efficiency.

minimum danger of cavitation erosion. minimum radiation of propeller noise.

minimum propeller induced vibratory forces. good stopping abilities.

favourable interaction with rudder to im-prove manoeuvrability.

dependability with minimum vulnerability. low initial and maintenance costs.

The choice of the propeller for a certain ship type must be a compromise between the above requirements.

*) Netherlands Ships Model Basin, Wageningen, The Netherlands.

Up till 1963 the investigations promoted by the Netherlands Navy were primarily focussed on the determination of the characteristics of conventional screw propellers with respect to propeller induced vibrations, cavitation, and

propeller noise.

Due to the limitations of the conventional

screw propeller at extreme conditions (hydrofoil boats. hovercraft etc.) as regards cavitation, noise and vibration characteristics, the Nether-lands Navy decided to promote the investigation of more or less unconventional propulsion

de-vices. It was expected that these propulsion

devices would favourably meet the mentioned requirements for various ranges of operating

conditions.

Most of the results of these investigations would also be of use for merchant ships. Sub-sequently. the Netherland,s Shipbuilding Industry was also interested in the results of the invest-igations on unconventional propulsion devices. As a result, a project concerning investigations on unconventional propulsion devices came into existence, sponsored jointly by Government and

3

.1. 2..

the shipyards and engineering companies parti-cipating in the Netherlands United Shipbuilding Bureaux Ltd. (now Rijn-Schelde Group), and

Lips N.V. Propeller Works. The Netherlands United Shipbuilding Bureaux Ltd., recognized the fact, furthermore, that to obtain full profit of the investigations on unconventional propul-sion devices, more fundamental research had to be performed on the occurrence not only of cavitation etc., but also on the behaviour of a ship in a seaway (the problem of maintaining the

ship's speed at severe sea conditions, slam-ming etc.).

The investigations were theoretical as well as experimental, and were performed at the Ne-therlands Ship Model Basin. A summary of the results of the most important non-confidential investigations which were carried out in the period from 1963 to 1969 are given in this re-port. These investigations included the

fol-lowing:

Fundamental cavitation research.

Development of a computer program for the unsteady lifting surface theory of screw pro-pellers.

Determination of the characteristics of ducted propellers.

Determination of the characteristics of contra-rotating propellers.

Determination of the characteristics of low-noise propellers.

Determination of the characteristics of the water-air ramjet.

Comparative tests with a tanker and a cargo-liner equipped with successively contra-rotat-ing propellers and conventional screw propel-lers.

The study of ship motions in a seaway with particular reference to naval vessels.

Slamming phenomena.

Added resistance of ships in waves.

2. Fundamental cavitation research.

Cavitation in a liquid is a phenomenon which

occurs when the pressure decreases below the

vapour pressure of the liquid. This causes the

formation of cavities and subsequently, when the pressure increases, the collapse of these

cavities. Cavitation on ship propellers, sonar

domes, torpedoes and other devices often means a definite limitation to their performance. A better understanding of the factors that promote or prevent cavitation was needed. A theoretical approach of this problem has been accomplished by Van der Walle [1]*). He first distinguished gaseous and vaporous cavitation. Gaseous cavi-tation occurs when bubbles adhere to the wall due to a positive pressure gradient and then grow by diffusion of air. Gaseous cavitation can

occur when the pressure is higher than the

vapour pressure of the liquid. Vaporous cavita-tion occurs quite suddenly when the pressure decreases below a critical value which is always smaller than the vapour pressure. Besides, the concept of wall nuclei and stream nuclei was

in-troduced. Wall nuclei are held by the wall of a solid

submerged body while stream nuclei exist in the free stream. Their action of introducing cavita-tion is schematically shown in Figure 1. Ac-cording to Van der Walle, cavitation is mainly introduced by wall nuclei. His analysis of the forces that act on a bubble either in the free stream or originating from a wall has already become a basis for many other investigators

in the field of cavitation.

*) Numbers in brackets refer to the references at the end of this paper.

cavitation of a stream nucleus stream nucleus Zr cavitating body growth of a nucleus in a wall crevice Liquid wall

_,\\

\\\

phase 3

,M\\ \\\V\

0-- 0

\\\phase

Figure 1. The process of the generation of wall nuclei.

cavitation of a wall nucleus interface flow phase 1 phase 2 -iThdary Ayer

CENTRIFUGAL PUMP, DciMOTOR Ian -COOLING WATER EAT EXCHANGER ROTARY _PUMP .FLOW

In order to verify the theory of Van der Walle by experimental investigations and to study the

influence of certain parameters on cavitation, it was considered essential to build an appropriate cavitation tunnel. In this tunnel all parameters which were thought to influence cavitation should be made adjustable and controllable. So, in 1966 a small high-speed cavitation tunnel was built.. A schematic drawing of the tunnel is shown in Figure 2.

The tunnel has a plexi-glass test section with an internal diameter of 40 mm. The maximum liquid speed in the test section is about 65 m/sec and the maximum tunnel pressure can be as high as 35 atm. As shown in Figure 2, an additional circuit, which contains a de-aeration tank and

filter, provides an easy control

of the air

content and purity of the tunnel water. Further details of this facility are given by Witte 121.. extensive test program in this timnel has been

carried out.

The test models consisted of

cylindrical shafts with a hemispherical nose., Three 'diameters were chosen: D-= 3, 16 and 12 min. The dissolved air content was varied be-tween a/as = 0.02 - 2,. The flow velocity in the FILTER

SAFETY

POSITIVE DISPLACEMENT PUMP

and

Figure 2. Schematic diagram of the High Speed Cavitation Tunnel.

p p

od

v HYDRO PNEUMATIC ACCUMUIATO:

Ifl

dd '-1/, P Vo

Po, i and Po, d are the static pressures in the test section 'corresponding to the state of in-cipient cavitation(the first appearance of cavita-tion), and the state of desinent cavitation (the disappearance of cavitation), respectively. More

than 1800 04 and 1800 ad -values were

de-termined and compared on a Reynolds number basis.

A presentation and analysis of results was' given by Van der Meulen [31. The main

con-clusions can be summarized as follows: DEAE RATION

TANK

RD TAM

test 'section Vo was varied between 15 - 60 m/sec. Moreover, different test procedures were used. Each test resulted in a cyi -value and a _ad -value, where:

5 Po, - V cr. . -2 % p Vo An = -WATER R P UP

-- Cavitation inception is a random phenomenon

which depends upon the pressure history of the liquid (Figure 3).

On the other hand, the disappearance of

cavitation is a much more repeatable

pheno-menon and is independent of the pressure

history. From this it follows that the use of ad in cavitation tunnel tests is preferable..

Both csi and ad decrease with decreasing air

content.

- od increased with an increase of the Reynolds number and with an increase of the diameter

of the test model (see Figure 4).

- All tendencies point to stream nuclei as being responsible for desinent cavitation.

This last conclusion is more or less contra-dictory with one made by Van der Walle. How-ever, in Van der Walle's theory two

assump-tions were made which have been proved tobe

untenable.

First,

he assumed that stream nuclei are

smaller that a few microns. In the U. S. A. it

was found that stream nuclei may have

dimen-sion of about 100

Second, he neglected convective gas diffusion

effects. It has been demonstrated since,

how-ever, that convective diffusion is about 1000

times faster than normal diffusion. Stream

nuclei are often trapped in vortices ofturbulent boundary layers. This may cause vaporous cavitation even while the mean static pressure is still above vapour pressure.

A method to suppress the vorticity in a

bound-ary layer is

the injection of high-molecular weight dilute polymer solutions. These polymers cause drag reductions of over 60% in turbulent

boundary layers. At the NSMB it was expected

that the suppression of the vorticity might also lead to a reduction of cavitation. In order to be

able to investigate this matter, it became

ap-parent that first a device had to be

made to measure the drag-reducing ability of a dilute polymer solution. So, a turbulent rheometer was built in which a fluid sample is forced to flow

through a capillary tube.

The pressure difference across the tube and

the flow velocity in the tube are measured and from these the drag-reducing ability follows.

Four different polymers of variously concen-trated solutions have been tested. One of these

0.9

0.8

0.1-Figure 3. Effect of pressure history on incipient cavita-tion versus cavitacavita-tion number.

0.90

/

150-12,

//

12012

/././

\.N 75-12 -12 150-6 _V. Z. 085 0.80-am 070-m z 0.65 A )1( )=1 )11( A o=i2mm 0=6MM A A en CD z A M X 7

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3 4 5 6 8 10' 2 3 4 5 6 8 10'

RE, REYNOLDS NUMBER

120-6 90-6 75-6 /r;--10-3 50-6 10-6 25-6₆

3 6

030 3

456

8 10 2 3 4 5 6 8 106 RE,REYNOLDS NUMBER

Figure 4. Effect of air content and model diameter on de -sinent cavitation versus Reynolds Number.

A 50/3 0.7 z. 05 4

-polymers, Polyox WSR-301, appeared to have a very favourable drag-reducing ability. A

dis-advantage, however,

for use in a cavitation

tunnel is the fact that the polymer degrades due to shear forces. More investigations on these subjects are needed. It is expected that these investigations will lead to some applications in the field of naval activities in the near future.

3. Development of a computer program for the unsteady lifting surface theory for ship screws.

Due to the presence of the ship hull, the flow behind the ship in way of the propeller is non-uniform. In general this flow is time-independ-ent, but the screw blades operating in this wake experiences time-independent entrance velo-cities. The circumferential variation of both the axial and tangential velocities is the origin of the periodically fluctuating pressure distribu-tions along the blade chords and leads to the un-steady force-pattern at the screw and at the stern of the ship. These periodically fluctuating pressure distributions lead to unsteady cavita-tion phenomena which may be serious from a view-point of erosion and noise radiation. in addition, these unsteady pressure distributions lead to propeller induced vibrations.

A large number of investigations have been carried out to describe the unsteady flow around a propeller blade. The two-dimensional ap-proach, using Sear's response function for the sinusoidal gust in a stripwise manner gives a very poor correlation with experimental results. Therefore a three-dimensional approach is ne-cessary. Using the linearized potential theory the action of the screw propeller can be des-cribed as a distribution of free vortices in the wake behind the screw. The geometrical repre-sentation of the screw is given in Figure 5. The

influence of blade thickness can be treated

separately in the linearized theory, and is only of importance if unsteady cavitation phenomena on the screw are to be studied.

The periodically fluctuating forces on the

screw blades are only caused by the lift distri-bution. The theoretical calculation of the influ-ence of the thickness of the propeller blades is relatively simple in comparison with the

de-termination of the lift. The principal problem that has to be solved, is the determination of the lift distribution over the propeller. In the lifting surface theory the screw action can be described by distributions of vortices over the screw blades and in the wake or by a distribu-tion of pressure doublets over the screw blades

only. Both formulations are given by Sparenberg [4] for the steady case. A slight modification gives the equations for the unsteady case as shown by Hanaoka [5]. In the computer program for the unsteady lifting surface theory for ship screws developed at the NSMB, the screw was represented by pressure doublets. This mathe-matical model of the screw is given in Figure 6. The potential of a pressure pole in ( p, e

t) which lies on the screw blade is:

y, t) -

(, p,

e, t)

4 Tr R

in which R is the distance between the pole in

the point ( p, e, t ) and the location of the

control point in ( x, r, y, t ). The strength of

the pole is defined by p ( p, 8, t). The

potent-ial of a pressure doublet in the same point is:

P ( , p, 0, t) a 1

(x, r, Y, t)

₄ ( )

-n _On R

in which n is the normal direction of the screw blade in the point ( , p, e,

t).

From the linearized equations of motion the velocity potential cl) of a pressure doublet in( ,

p, e, t) can be found: 1 (1) (x, r, y, t) = 4,TU

f P

ID, t X -X 1 a 1 U ) an (-137) dx1-4,p (x, r,

Figure 5. Helicoidal propeller surface.

7

yd

,

(4 )

cD (x. r, p, t)

(1)

41 (x,r,,t).

/

0) af(xl) + r sinta(x-r)-wt-0

U(1+er) (x1)'+r*.i.cs-2rs cos(a-o t-clek

(x,r,t) j-u-f ( x,r,,,t - )(t-j') dx,

Ji

m(wt -a(x-1))K(x,rA,c,)

r X,

im(a.r+ airr+ c os(a-c +1-2n) 3larr- rsin(ar+44.2n)lia ft, sin(ar4 271)1

Z e dr

k.0 Fe

Figure 6. Formulation of the lifting surface integralequation.

This velocity potential is necessary in order to be able to use the boundary condition that on the screw blade the normal velocity is zero. For a distribution of pressure doublets over the screw blade the velocity potential becomes:

1

=;fff

P(

X-X

131

u ) an ( R ) dxl

when A is the surface of the screwblade. The normal velocity at the blade inthe point ( x, r, t) can be found by differentiation of

the velocity potential in the normal direction at this point: w n 4-rrU acD 1 ail

a.1.xmc,

A -p, @, t x-x a (--) dx dA. U an R 1

From the condition Wn= 0 at each point (x, r,

t) of the screw blade, the pressuredoublet

strength, which is the local loading of the pro-peller blade, must be calculated. An extensive description of the formulation of this integral equation is given by Verbrugh [6], where the

number of blades of the screw propeller is also

6, t

F _tT-ax.ridt=0 m= harmonic

n number of blades U mean velocity

ue,wo= wake in x and direction

r.

pressure dipole inner and outer radius

xvxt. leading and trailing

edge.

taken into account. The inner integral in the

last equation is an integration over the wake. It can easily be seen that by integrating over A the

distance R becomes zero. Thus the integral

equation has a singularity and can not be solved directly by numerical methods. The separation of this singularity is very complicated and in

various investigations different assumptions

have beenusedto make the solution of this

prob-lem possible.

A very rough simplification is the reduction of the blade area to a line, (lifting line theory). Investigations based on this simplification were carried out by Joosen [7].

Normally the aspect ratio of a screw blade is too low to permit the negligence of the effect of the blade width. In addition, this effect may not

be neglected by considering the interaction ef-fects between the different blades. Tsakonas et al have been working on this problem over a number of years trying several simplifying as-sumption [8], [9], [10] and [11].

The most rigorous simplification in their

ap-proach is the alteration of the wake in the

normal direction of the flow. The wake is

chang-ed into a set of circular discs

normal to the

mean flow as indicated inFigure 7. The normal

directions n and fi are taken in axial direction. It

is clear that assumptions are

only valid for

2 ) '( 3 -1 = y, dA '9,) -, p, y, ar

u-wfd

Figure 7. Wake and staircase approximation.

propellers with low pitch ratios P/D. In addi-tion, a general distribution of the chordwise loading was used [11].

A comparison of the results of calculations with this theory and the experimental work of Wereldsma [12], [13] was made.

The exact solution of the integral equation is

also formulated by Tsakonas [14], though only for

a circular type of blade form and a flat plate chordwise loading (the first term of the Birn-baum series).

A complete and exact solution of the linearized integral equation has been developed by

Ver-brugh [6]. In this solution no simplification is made, while an arbitrary chordwise pressure distribution is possible. With this theory it is also possible to determine the effects of skew, rake and arbitrary blade form. A numerical

pro-gram has been made to compute the loading of the

blades. The convergence of the chordwise press-ure distribution is however not realized by using more terms of the Birnbaum series:

p = co cotg +

C sin p 9.

2 p = 1 p

Using this formulation the leading edge sin-gularity is not only present in the first term of the series but also in the other terms.

There-fore the series is

not convergent. This dif-ficulty has been solved by using the following formulation:

2 cosp 0 + cos (p + 1) 8

= 2 C

p = 0 pT

sine

In this case the coefficient C tends to zero for

the higher order terms.

This difficulty was later solved in another way by Tsakonas [14].

The numerical program for the exact solution developed at the NSMB is now nearly ready. For the steady case and for the first term Co of the Birnbaum series in the unsteady case (flat plate

200 15 100 SO 100 -150 200

Figure 8. Phase angle

torque). 20 15 100 .f, La, 50 cn

go

-50 100 150 200 2 te 11)

Figure 9. Phase angle of third harmonic (thrust and

torque).

pressure distribution) some results are avail-able. These results are given by Kuiper [15]. The phase angle between the angular position of the screw and the vibration forces and moments

9

0.6

a

10 _4E/A0

3.0.45 0 Measured.

Calculated (Flat Plal Distribution Pressure 0 4 06 08 -10 AE/A0 .r J-075 o Measured. Calculated(Flat Pl. Distributio Pressure

of third harmonic (thrust and

-were also computed and compared with the ex-perimental results of Wereldsma [13]. Some of the results are given in Figures 8 and 9. From these figures it can be seen that the theoretical and experimental results are in good agreement with each other.

The program is to be able to calculate the unsteady pressure distribution on the screw blades. From these pressure distributions the vibratory forces and moments can be derived. The vibratory blade spindle torques can be com-puted which is of importance for controllable pitch propellers. Further results of the computa-tions will be given soon.

In addition, the program will be extended so that cavitation predictions can be performed (by taking the thickness of the blades into account) and also the interaction effects between screw, hull and rudder. It will then be possible to cal-culate the vibratory forces on rudder and after-body of the ship.

4. Determination of the characteristics of ducted propellers.

Insight into the working principle of a ducted propeller can be gained by the application of fundamental momentum relationships. Figure 10 shows the simplified system by which the ducted propeller can be replaced. Here the screw propeller is represented by an actuator disc rotating at infinite angular velocity. The tangential induced velocities and, consequently. the losses due to rotation of the fluid are then zero. The influence of friction is neglected.

With the momentum theory the following ex-pressions for the ideal efficiency ni and the ratio between the velocity at the impeller plane

and the undisturbed stream velocity _VP/VA can

be derived: 2 Ti 1 + V 1 + CT CT vp/vA - 2 [ -1 + \/ 1 + T where CT -, and 7 = --* T and 2 , 2 T P VA Tr/ . D 4

T denote the total thrust and the

impeller thrust respectively. D is the propeller diameter.

These formulae are graphically represented in Figure 11. From this figure it can be seen that due to the nozzle action the inflow velocity

of the impeller can be either less or greater

than the inflow velocity of an open propeller under equal conditions.

For a thrust ratio 7 equal to 1.0 no force acts

on the nozzle and the flow pattern is comparable with that of an open screw. With decreasing

values of 7, the nozzle produces a positive

thrust, the inflow velocity of the impeller is in-creased and an improvement in ideal efficiency is found. For thrust ratios 7 greater than 1.0, a

negative thrust or drag force acts on the nozzle, the inflow velocity of the impeller decreases

and the ideal efficiency will be lower.

Figure 10. Control volume used for momentum con-siderations.

VA U2

112

,.

oso

-

_{I 1 I I I}

ELM

PAM

I

- . . 64 1 'i11 080 ,, 0.70

Mil.

.,.dim rd

o.

Mill4 MIETIO

260

/I

-....

4

1.,24

worP-14

NuRci

150

mg

)4. ,,n Q25 QS CT 2 I. 16

Figure 11. Mean axial velocity and ideal efficiency of the ducted propeller. 20 1.6 04. 0

-.

Decelerating nozzle, VA. U2 'VP Accelerating nozzle VA H Propeller disk, -,r-,

Figure 12. 'General form of streamlines enforced by dif-ferent. nozzle types

Thus application of a flow-accelerating type of nozzle offers a means, of improving the

ef-ficiency; application of a flow-decelerating

nozzle leads to an increase of the static press-ure at the impeller (due to the reduction of the flow rate) which. may be attractive' for the re-tardation of screw cavitation phenomena. The particulars of flow-accelerating and decelerat-ing nozzles are shown in Figure 12.

During the past years extensive investigations on ducted propellers were performed at the NSMB. The investigations dealed with both ac-celerating nozzles (Van Manen [161), and de-celerating nozzles (Oosterveld '[171).

The investigations on decelerating ducted

propellers were sponsored by theOffice of Naval Research of the United States Navy and the Ne-therlands Navy.. The investigations performed for the US Navy were focussed on the

deter-mination of the characteristics of systematic series of flow-decelerating nozzles. The in-vestigations for the Netherlands Navy were fo-cusses on the determination of screw series

specially

designed for use

in 'decelerating

nozzles.

Some general data can be derived from these investigations. The choice of the shape of the nozzle profile and the nozzle length depends

mainly on requirements with respect to efficien-cy, danger of flow separation on nozzle and cri,-tical cavitation numbers on nozzle and impeller. General data for the design of ducted propellers can be obtained from vortex theory as described in [161 and by taking into. account the effect of friction. These data are given in Figures 13, 14

and 15.

The result of an analysis of the sectional lift coefficient CL of the nozzle profile is given in Figure 13. Here s/1 and L/D denote respectively the thickness ratio of the nozzle profile and the length-diameter ratio of the nozzle.. The value of the lift coefficient of the nozzle profile gives. an indication of the danger of flow separation on the nozzle. From two-dimensional airfoil flow follows that flow separation may occur when the lift coefficient CL exceeds a value of about 1. From Figure 13 it can be seen that the loading of the accelerating nozzle is strongly restricted by the risk of flow separation on the 'interior surface of the nozzle. The risk of flow

,separa-.0 .E w w Z N 3 0 - 0 0 04 U_{ix 4} 4( y A. IX -_ta Ii 0. 0. 0.4 0.2 0.2 10, If i I \ 1

\

11

I I ,

\

lallILW i i

11\1/4\

1 I

ik,

go . 0 . .0.6 gSA.. 0 r .S0

rMiViK

El

r

mem"

N g

I.1% . .0,3 II ,I L/0 . 0.6 S/i. .0.6 il = 0 ' . 0.09 ' 11 1 1

1141

1 J0.7 10 0.9 1.0 1.2 1.3

Figure 13. Sectional lift 'coefficient of nozzle profile. VA 1.0 04 0.6 0.8 1.0 I 1.4 \

\

0.6

09 07 05 04 0.2 Oh as CP, 0.6 1.6 1 64 25 10 08 12 -51. 06 0.8 10 T_. Ideal efficiency

--- Efficiency taking into account

the effect of frictional nozzle drag.

tion on the exterior surface of the decelerating nozzle is small, however, even in the case of large thrust ratios T.

An analysis of the optimum ideal efficiency and the additional efficiency losses due to the frictional nozzle drag has been made for ducted propellers with different length-diameter ratios and for zero thickness of the nozzle profile. The

result is given in Figure 14.

From this diagram the following conclusions can be drawn.

The use of a nozzle only leads to an increase of efficiency at higher screw loads (CT > 1), when the gain in ideal efficiency exceeds the loss in efficiency due to the frictional nozzle

drag.

A nozzle lengthdiameter ratio of L/D = 0. 5 -1. 0 is optimum from the view-point of effi-ciency at higher screw loads.

The application of the flow decelerating nozzle may be attractive if retardation of propeller cavitation phenomena is desired. The reduction of the flow rate inside the decelerating type of

nozzle results in an increment of the static

pressure at the impeller. This is attractive

from the point of view of retardation of screw cavitation.

However, the duct itself will produce a nega-tive thrust ( T ' 1 ) . In order to compensate for this thrust loss (induced nozzle drag), the

im-1 I 1 bvi-m..., I 2 Efficiency Efficiency of open screw of ducted .. proi

.1

45

11111

---

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_...

101

ur

_N.,

\

n # 10

/---Jii

4 5 8 2 eller yo-on CL,.10 L-084 050 078 100 073 200 069 Q25 0 CT 2 4 8 16

Figure 14. Optimum ideal efficiency and efficiency losses due tofrictional nozzle drag of a ducted propeller system.

Figure 15. Minimum static pressure at impeller blades of a ducted propeller.

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peller loading must be increased. An improve--ment of cavitation properties of the impeller will therefore only be obtained if the gain in

static pressure at least compensates the

un-favourable effect of the increased screw loading. The result of an analysis of the minimum press-ures which may occur at the blades of a ducted propeller is given in Figure 15. This diagram shows that, for the particular screw considered (blade-area ratio AE/Ao = 1..0 and number of blades z = 5), only for low values of the thrust coefficient CT will the flow-decelerating type of nozzle favourably affect the cavitation

proper-ties of the screw.

If ducted propellers with

KT,KT5%,IK5 and Tio as ,tunclions of J

.Figure 17. Open-water test results of nozzle no. 33.

larger blade-area ratios of the impeller or with more rotor- (and eventually more stator-) rows are considered, the decelerating nozzle* may even for larger values of CT favourably affect the cavitation properties of the screw.

Application of the decelerating nozzle results in a reduction of the pressure at the 'exterior surface of the nozzle. From a comparison be-tween the minimum pressures which occur at the exterior surface of the nozzle and at the im-peller blades, it can be concluded that except in the case of very short nozzles or very low

load-ed systems, the impeller cavitation is more

critical.

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Systematic series of model tests with decele-rating ducted propellers have been performed. The design of the nozzles was based on the vortex theory as described in [16]. The

varia-tion of the design parameters considered is

shown in Table I.

Table I

Variation in design parameters for decele-rating nozzles.

with nozzle no. 33. The pitch distribution of the screws depend on the velocities induced by the nozzle at the impeller plane and on the radial load distribution of the screws. For the design, use was made of [16]. Particulars of the screw models are given in Table II and in Figure 16. The screws were located in the nozzle with a uniform tip clearance of 1 mm.

Table II

Particulars of screw models of the Kd 5-100 series.

Diameter

Number of blades

Pitch ratio (at 0.7 D) P/D

Blade area ratio AE/A.

Blade outline

Blade section

Propeller indicated by nos.

240 mm

5

1.0 - 1. 2 - 1. 4 - 1. 6 - 1. 8 1.00

Kaplan type

NASA 16-parabolic camber line

3920 - 3931 - 3932 -3933 and 3934

The tests were carried out with the tank ap-paratus for open-water tests of ducted propel-lers. The usual routine of open-water tests was

followed, the rpm of the screw was kept constant and by varying the speed of advance the desired value of the advance was obtained. The rpm was

chosen as high as possible to obtain

a high Reynolds number. Speeds above 3m/sec could not be investigated on account of the restricted speed of the towing carriage.

The open-water test results were faired by

computer and plotted in the conventional way

with the coefficients:

KT -p n2D4 TN Km

24

p n D KQ

₂₅

p n D no

2r

K

as functions of the advance coefficient J (J =

Va/nD). The open-water diagram of the Kd 5-100

screw series in combination with nozzle _{no. 33} if given in Figure 17.

The application of vanes located downstream of the impeller of the decelerating ducted

pro-15 Nozzle CT r

L/D a/L

S/L d/D 30

0.95 1.00 0.6

0.5 0.15 0.20 31 0.95 1.15 0.6 0.5 0.15 0.20 32 0.95 1.30 0.6 0.5 0.15 0.20 33

1.00 1.20 0.6

0.5 0.15 0.20 34 1. 00 1.20 0.6 0.5 0. 09 0.20 35

1.00 1.20 0.9 0.5

0.10 0.20 36

1.00 1.20 1.2

0.5 0.075 0.20

The experiments with these nozzles were all carried out with a series of five bladed Kaplan type screws (Kd 5-100 series). The Kd 5-100 series screws were designed in combination

Z

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,

I 0,K

KT

peller is of importance with respect to the pre-vention of strong vortices along the hub.

These vortices may result in hub vortex

cavita-tion. In addition the efficiency of the ducted

propeller system may be increased by the

ap-plication

of vanes or a stator if the latter

eliminates the rotational losses without incur-ring excessive frictional resistance.

Based on the theory as discussed in [16j a six-bladed stator was designed for application downstream of the Kd 5-100 screw with P/D = 1.4 in nozzle no. 33. Particulars of the stator are given in Figure 18. Open-water tests with

09 110

Figure 19'. Open-water test results of Kd 5-100 screw series in nozzle no. 33 fitted with stator.

1,3

j

the Kd 5-100 screw series in combination!with

nozzle no. 33 and the six-bladed stator were performed. The result is given in Figure 19, The stator was fixed to the nozzle ; consequent=:, ly, KTn denotes in the diagrath the thrust coef-ficient of nozzle and stator.

In the case of application of a screw or a

ducted propeller behind a ship, a part of the ro-tational losses of the screw or the impeller is already eliminated due to the presence of the rudder in the propeller slipstream. Consequent-ly, the gain in efficiency due to the stator will be lower than may be expected from comparing

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the Figures 17 and 19.

For comparison purposes, model tests have been carried out to, determine the effect of the presence of a rudder on the characteristics of a

'ducted propeller

system.. A more or less

common rudder configuration was therefore

fitted to nozzle no. 33 and open-water tests

were performed with this system in combination with the Kd 5-100, screw series., The result is given in Figure 20.

For design purposes, various practical re-sults can be. derived from Figures 17, 19 and

20. In the case where the power P, VA and n are given, the determination of the optimum diameter from a point of view of efficiency of the 'ducted propeller system can be solved by plotting 70 and 5 ,( 5 = 101. 27 ) as functions of

the coefficient :.

Bp (Bp = 33.08 K 1/2 ,J -5/2

In the case where the thrust T, VA and D are, given. the determination of the optimum rpm n

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Figure 22. Showing the effect of rudder and stator on the ducted propeller characteristics.

is of the accelerating flow type, see [18].

Screws of the B 4-70 screw series are usually applied behind single screw ships.

Typical Bp values for different ship types are as follows

torpedo's Bp < 10

twin-screw ships 10 - 15

fast warships (frigates) 10 - 25

single-screw cargo ships 15 - 35

tankers 35 - 70

towing vessels (tugs, pushboats) 80

The accelerating nozzle (nozzle no. 19A).

when compared with a conventional screw pro-peller (B 4-70 screw series), gives rise to an

400 360 320 28 240 6.6. 20 160 140 80 40

NozzLe 33 Screw series Kd S100 --stator . -rudder 02 03 04 05 07 2 3 4 5 7 10 CT 56 14 1.0 0.6 20 is

Figure 23. Optimum characteristics of an accelerating nozzle (nozzle no. 19A), a decelerating nozzle (nozzle no. 33) and the B 4-70 screw series in open-water.

. Kd 5-100 screw series without with stator and nozzleno33 der and

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Figure 21. Optimum characteristics of the Kd 5-100 screw series in successively nozzle no. 33 and nozzle

no. 33 with stator and rudder.

can be solved by plotting no and 8 as functions of the thrust coefficient :

8 -2

CT ( CT =

. KTJ) .

For comparison, the optimum efficiency

curves of

the Kd 5-100 screw series

in

combination with successively nozzle no. 33.

nozzle no. 33 with the six-bladed stator and

nozzle no. 33 with rudder, are given in the Fi-gures 21 and 22. From this diagram it can be seen that both the stator and the rudder give an increase of the efficiency. Further, it

is

inte-resting to note the effect of the stator on the

optimum diameter and on the optimum rpm of

the system.

For comparison, the optimum curves for ef-ficiency n0, diameter coefficient 5 and thrust ratio T of the Ka 4-70 screw series innozzle no.

19A, the B 4-70 screw series and the Kd 5-100

screw series in nozzle no. 33 are given in Fi-gure 23 on a base of Bp. Nozzle no. 19A is the

standard nozzle profile applied by the NSMB in

the case of heavily loaded screws. This nozzle

5 45 7 10 ap.5 30 40 50 70 1 260 20 10 2 L 0 2 14 Pio 57 14 06 '70 6 Si 4 10% - -1.8 4-70, -> . 1

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-improvement in open-water efficiency no in the case of heavy screw loads. The decelerating nozzle (nozzle no. 33) has a relatively low open-water efficiency. The application of the decele-rating nozzle becomes attractive, however,

when other than hydrodynamic factors influence the choice.

It can be seen from Figure 23 that at low Bp values the open-water efficiency of both the ac-celerating and the deac-celerating nozzle decreases

with respect to the efficiency of the B 4-70

screw series. This fact can be explained by the relative increase of the frictional and induced drag of the nozzle. The curves for the diameter coefficient 5 of the accelerating and the decele-rating nozzle almost coincide; the B 4-70 screw series has a larger optimum screw diameter. It is interesting to note that the curves for the dia-meter coefficient based on the maximum diame-ter of the system 5* of both the accelerating and the decelerating nozzle and of the B 4-70

screw series also almost coincide.

All these investigations have given more de-tailed data to determine suitable nozzle shapes for ducted propellers. Based on these data some special nozzle shapes were designed and tested. A propeller behind a ship operates in a non-uniform flow. The inflow velocity of the

propel-ler can be made more constant over the screw disc by surrounding the propeller with a non-axisymmetrical nozzle which is adapted to the wake distribution and flow direction behind the

ship.

In the case of a single-screw ship the intake velocity will be lower in the upper part of the screw disc than in the lower part. Consequently, the propeller is more heavily loaded in the up-per part of the screw disc. The inflow velocity of the propeller can be made more constant by surrounding the propeller by a non-axisymme-trical nozzle which accelerates the flow in the upper part of the screw disc and decelerates the

flow in the lower part of the disc.

In the case of twin-screw ships the propellers operate in a varying inflow, due to the shaft in-clination. The inclination is a consequence of the fact that the propeller shaft has a sizeable inclination to both the horizontal and the buttock lines in way of the propeller. The non-axisym-metrical nozzle must be designed in such a way that the actual effective incidence changes of the

blade sections of the impeller will be as low as possible during a revolution. From the stand-point of retardation of screw cavitation, the ap-plication of the non-axisymmetrical nozzle will be very attractive.

Investigations with respect to the latter type of non-axisymmetrical nozzle are underway at the NSMB. The first tests with a non-axisym-metrical ducted propeller system designed fora flow inclination of 9 degrees have given good

results. The shape of the nozzle profile is

shown in Figure 24.

hull buttock in way of propeller centerline

II'Al

bracket 1111,noz:laeer m propeller disc looking inboard WA vs ft Looking forward

Figure 24. Hull of twin-screw vessel fitted with non-axisymmetrical nozzles.

5. Determination of the eharacterististics of contra-rotating propellers.

The investigations on contra-rotating propel-ler systems consisted of the following.

A systematic series of contra-rotating pro-pellers were designed and manufactured. Each propeller system consisted of a four bladed for-ward screw and a five bladed aft screw. The forward and the aft screws were each designed for equal power absorption. Tests were carried out in the towing tank to determine the open-water characteristics of each system. The de-sign of the forward and aft propellers was car-ried out according to lifting line theory as des-cribed in [19]. The imminent danger of damage being caused to the aft propeller by the cavitat-ing tip vortices of the forward propeller, which may impinge on the aft propeller, was avoided

19 eso° 0.90° 0.180 8.270° 8.2700 8.90° direction, of rotation, foutboarca

- 'Table III

Principal characteristics of screw models of contrarotating propeller series.

by reducing the diameter of the aft propeller. This reduction was based on the expected slip-stream contraction at the design condition.

In addition, this reduction is attractive with regard to the fact that for equal screw loadings a five-bladed propeller has a smaller optimum diameter than a four-bladed propeller with equal blade area ratio.

Five sets of contra-rotating propeller were designed. Details of the resulting contra-rota-ting propeller series is given in Table III. With

this series, various application possibilities

may be studied.

The results of the open-water tests in the

towing tank were faired and plotted in the con-ventional way by using the KT, KQ and no co-efficients as functions of the advance .coefficient J.

These results are shown in Figure 25.. In this

diagram set of contra-rotating propellers is

considered as one propulsion unit. The thrust T and the torque Q are thus the sum of the forward and aft screw thrusts and torques. The repre-sentative diameter of the system was taken to be that of the forward propeller. In addition, the

aft propeller thrust-total thrust ratio Taft/T

and the aft propeller torque-total torque ratio Qaft/Q are important, and also given in this figure.

For design purposes, data can be derived from Figure 25. In the case where the power P. speed of advance Va and rpm n are given, the deter-mination of the optimum .diameter from an ef-ficiency point of view is achieved by plotting r-and the diameter coefficient 5 as functions of the coefficient Bp.

For comparison purposes, the optimum curves for the efficiency and the diameter coefficient

of the contra-rotating screw series, and the

comparible B 4-70 screw series are given in

Figure 26.. At the top of this figure, the ranges.

of Bp-values, typical for different ships are in-dicated. The lightly loaded screws of fast ships are situated on the left hand side and the heavily loaded propellers of towing vessels etc., on the right. This diagram gives ready -at hand informa-tion as to which type of propeller system will, be best with regards to efficiency for various applications. For fast ships (cargoliners and navy vessels), contra-rotating propellers ap-pear to give a higher efficiency than conventional

4 5 T

SET _{Forward Aft Forward Aft Forward Aft Forward Aft Forward} Aft

Diameter (mm) 210. - 179.34 208. - 182.72 210. - 191:01 217:59 20a233 210. - 198, 90

Number of blades 4 5 H 4 5 4 5 4 5 4 5

Pitch ratio at 0.7 R 0.627 0. 957 0.779 1.034 0. 931 1. 110 1.083 1. 196i 1. 235 1.306

Expanded blade area ratio 0.432 0.507 H 0.432 0.515 0.432 0.523 0.432 0.531 0.432 0.539

D aft/D forward , 0. 854 0. 878 0. 910 0.935 0. 947 LI kiCtiltha 1 , 1 1 To __alt 1 4 _....-."( Torque to I I , I 1 Q2 06 08, 110 12 114

Figure 25. Openwater test results of the, c,ontrarokit -ing propeller series.

05 jet T tot 1.0 as 1 1 2 3 I 00 10 1.5 1.0 05

Figure 26. Comparison of contra-rotating propeller series and B 4-70 screw series.,

°

Figure 27_ Arrangement ,of contra-rotating propellers behind a cargo liner model.,

screws. It must be noted however, that the rud-der behind the conventional screw partly re-moves the rotational velocity from the propeller slipstream and thus .improves the efficiency of this propulsion system. It is obvious that this improvement inefficiency will not be found when contra-rotating propellers are applied.

Based on these open-water test

results,, contra-rotating propellers were designed for a

tanker and a cargo-liner. Comparative tests

were carried out with ship models equipped with a contra-rotating propeller system and a con-ventional single screw arrangement. In Figure 27, the contra-rotating propeller arrangement at the stern of the cargo-liner is shown. The propulsive efficiencies,, cavitation

character-istics, propeller induced vibratory forces and the stopping abilities of the tanker model and the cargo liner model equipped with both pro-pulsion systems were determined. The results tare given and discussed in a later section.,

6. Determination of the characteristics of low noise propellers.

Usually the variations of the flow field behind a ship at the location of the screw can be split

up into two components., viz.:.

The radial variation, especially of the axial velocities. This variation does not lead to un-steady phenomena. The screw can be adjusted to this radially non-uniform flow, and optimum, efficiency and cavitation properties May be expected.

The circumferential (at a given radius) varia-tion of both the axial and tangential velocities. This non-uniformity is the origin of the un-steady pressure distributions along the blade chords and leads to the periodically fluctuating force-patterns at the screw and at the ship. These unsteady pressure distributions also lead to unsteady cavitation phenomena which may be serious from a view-point of damage and noise radiation.

For naval applications, where twin-screw vessels are common practice, the unequal load-ing of the propeller is mainly caused by the oblique flow due to the shaft inclination. Recent-ly, comparative cavitation experiments have been performed at the NSMB with lightly loaded screw propellers (suitable for application on navy ships). with different radial load distribu-tions and contour shapes in oblique flow in cavi-tation tunnel no. 1. The following screw models were manufactured for these tests:

a conventionally designed screw with an op-timum radial load distribution from the view-i point of efficiency,

a low-noise propeller.

This screw has

strongly unloaded blade tips and a 'delta' contour,

a low-noise propeller with a conventional contour shape. This screw has the same ra-dial load distribution as screw B and has thus relatively lighter loaded blade tips.

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Particulars of the screw models are given in Table IV and in Figure 28,

In cavitation tunnel no. 1, these screws were

tested in axial- and in 6, 9 and 12 degrees

oblique flow. The cavitation patterns of these screws at a shaft inclination of 12 degrees are shown in Figure 29. The cavitation patterns of screw C at shaft inclinations of 0, 6, 9 and 12 degrees are given in Figure 30.

Figure 28. Particulars of propeller models A, B and C.

From these tests it follows that screw C has the best characteristics with respect to cavita-tion in oblique flow. In the near future these screws will be analyzed by means of unsteady lifting-surface theory. The results of the cavita-tion tests may then be correlated with the theo-retical results.

Further research on low-noise propellers has as yet to be performed, particularly as regards

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to the profile type to be used for blade section shaping. Also the effects of blade section thick-ening on the cavitation properties have to be de-termined. A new research program as regards these problems of blade section shaping of

low-noise propellers has been started.

7. Determination of the characteristics of the water-air ramjet.

An application of two phase flows for naval purposes is found in a special type of propul-sion: the water-air ramjet. The development of this propulsion system may be promising for

Figure 29. Cavitation properties of low-noise propellers, high speeds. A scheme of the system is given in

Figure 31.

Water enters a diffusor at an undisturbed

stream velocity. In the diffusor the kinetic energy of the flow is converted into pressure. The diffusor is followed by a mixing section in which the high pressure water stream is mixed with a large amount of air. The air normally enters through orifices situated in the wall of

the mixing section. The water-air mixture leaving the mixing section is accelerated in the exhaust nozzle where pressure is converted into kinetic energy. This system is completely ana-logous to the air ramjet. In both cases the thrust is 'developed by the reduction of the density of

Figure 30. 'Cavitation properties of low-noise propeller the working medium after the pressure has been, 'C at 0,6, 9 and 12 degrees oblique flow. raised in a ,diffusor. In the air ramjet this

dens-120 oblique flow.

'Table IV

Details of propeller models A, B and C.

23 PARTICULARS OF PROPELLERS Propeller A B

c

Diameter D' 4.200 mm 4.200 mm 1 4.2.00 mm Number of blades 5 5 5, Pitch at root

4.307 mm 3.918 mm 3.910 mm

'Pitch at blade tip 4. 055 mm 3.,851 mm 3.828 mm

'Pitch at 0.7 R P 0. 7 4.448 mm 4. 578 mm 4. 574 mm

'Disc area Ao 13.854 m2 13,854 m2 13.854 m2

Expanded blade area AE 12.629m2 11. 722 m2 12.629m2

P 0.7 D 1.059 1.090 ' 1.089

d/D 0.196 0.196 0.196

AE/Ao 0. 912 0. 846 0. 912

Figure 31. Showing principal parts of the water-ramjet.

ity reduction is realized by combustion, where-as in the water ramjet this reduction is caused by the introduction of gas bubbles into the mixing section. Due to this lower density the expansion in the exhaust nozzle to the ambient pressure results in an exhaust speed that is higher than the undisturbed stream speed.

The results of theoretical investigations

performed at the NSMB on the application of the water ramjet are given by Van der Walle [20],

Witte [21], [22] and Van Gent [23]. Van der Walle

[20] has derived expressions for the specific thrust and efficiency of the water ramjet based on the momentum theory. In his approach, the two-phase flow is described as a homogeneous compressible medium. To calculate the shape of the expansion nozzle, the assumption is made that the internal force in the mixture is purely viscous. This is true for a mixture of a very high quality (small bubbles uniformly distributed over the fluid). However, it is questionable whether this quality can be achieved in practical cases. Witte [21] extended the theoretical model as used by Van der Walle and gave a description of the mixing process. The assumption is made that formation of a homogeneous mixture is achieved by air injection through the porous wall of the mixing section.

The description of the two-phase flow by a homogeneous compressible medium, is abandon-ed in the model usabandon-ed by Witte [22], and the dif-ferences of the two phases (water and air) in

pressure, temperature and velocity are taken into account. The model is further extended with the calculation of the influence of the internal and external frictional losses. The mixing pro-cess, however, is idealized to a loss-free, in-stantaneous conversion of the co-axial air and water flow into a homogeneous mixture.

Ex-tensive calculations are performed with this model and in particular the effects of the length-diameter ratio of the ramjet and of the temper-ature of the injected gas are investigated.

To distinguish the ramjet from another appli-cation of two-phase flow in propulsion, the mist-jet, it is useful to consider the range of volu-metric mixing ratios which are suitable from a hydrodynamical point of view.

The lower limit of this range is given by the 'no air' condition, the upper limit follows from the condition that water must be the continuous phase. If air would be the continuous phase, the forces interacting between the two phases would be much smaller (these forces are roughly pro-portional with the density ratio), and a longer expansion nozzle is required. This is unaccept-able for high speed underwater propulsion due to the high frictional losses. It has been shown by Van Gent [23], that in the thus defined range of mixing ratios, the performance parameters of the ramjet van have practical values. Figure 32 shows the specific thrust-efficiency relations for various speeds, applying to a practical de-sign. When these theoretical results are suf-ficiently supported by experiments, a discus-sion of the feasibility of the ramjet principle is possible. The theoretical work described so far offers a means to determine the optimal prin-cipal dimensions of the water ramjet and the optimal mixing ratio.

The experimental investigations, primarily concern the internal performance of an air-water ramjet with circular cross-section under cir-cumstances corresponding to a forward speed of 10 - 35 m/sec. The development can be divided roughly in three stages. The first stage covered the period from 1963 to 1966. In this period a water ramjet model with a central air-injection plug was tested. The intention of this plug was to disturb the waterstream to improve the mix-ing of air and water. The results of these tests are given by Witte and Van den Brand [24], [25]. Different shapes of the central plug and different locations of the air-injection slit were tested. From these tests it was found that a central plug gives an unacceptable high resistance. Conse-quently the model has a low specific thrust and efficiency.

In the second stage, from 1966 to 1967, the idea of fitting the water ramjet with a central

no

-'0.64 0.60 0 56 0.52 0.48, 0.44 0.40 .0.36 D.32 0 28 024 02 036

plug was abandoned. Other methods ,of air

in-jection were considered during this period

these being the injection of air through a slit or holes along the interior surface of' the ramjet and through holes in radial vanes. The direction of the air injection was varied during the tests and different types of expansion nozzles were considered. The model was not fitted with a se-parate mixing section, so that the mixing of air and water had to take place in the expansion nozzle.. From the tests it was found that cir-cumferential air injection gives a better per-formance than injection through a central plug. In the case of air injection through vanes, the performance was found to be low. Apparently the additional resistance of the vanes exceeds the favourable effect of a better mixing

pro-cess. The results of these tests are

given by Van den Brand [261. In addition some experi-ments were also performed in this period to determine the effect of the injection of hot air on the performance. Although the duration time of the air-water mixture in the ramjet is very

short, the exchange of heat between air and water is too, large to find an increase in per-formance.

In the third stage, covering the period from, 1967 till 1969, water ramjet models with a cir-cumferential injection of air along the interior surface of the ramjet were tested.

The model had a mixing section. The cross-sectional area of the mixing section increased downstream in such a way that the water speed and pressure areeonstant in the mixing section. Tests were performed to determine the optimum length of the mixing section with regard to effi-ciency for different air-water mixing ratios. In addition, tests were carried out to determine the optimum shape of the expansion nozzle. The results of these investigations were given by Van Gent and Van den Brand [27]:. The results of the

investigations performed in this third period show a definite increase in efficiency and thrust of the water ramjet when compared with the configurations tested in the first and second pe-riods. 25 [[ 1 I

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Figure 32. n, relations for a practical. ramjet design.

014 034

The differences which were found between the results of the theoretical and experimental in-vestigations can probably be explained by the following facts:

110

- A small model had to be used in the tests, whereby the frictional losses in this model be-come relatively too large.

For the full scale water ramjet in a free

stream, diffusion can readily take place be-fore the water enters the diffusor. Thus the

diffusor can be shortened, thereby ensuring 20.0 00

lower frictional losses.

The external flow around the ramjet was not simulated in the tests carried out. Probably this fact has a worsening effect on the expans-sion process.

In the theory, the mixture is assumed to be homogeneous. During the model tests this was certainly not the case.

In the near future tests are planned with a model of the water ramjet in the High Speed Ba-sin of the NSMB. These tests will give the final answer to the question whether or not the appli-cation of a water ramjet for high speed ships forms an attractive propulsion system.

8. Compartive tests with a tanker and a cargo liner equipped with successively contra-rotating propellers and conventional screw

propellers.

As already mentioned in section 5, model

tests have been carried out to obtain a compar-ison of the propulsive quality of a tanker and a cargo liner both equipped with successively contra-rotating propellers and a conventional

screw propeller.

Resistance and self-propulsion tests were conducted in the deep-water basin of the NSMB, in accordance with established procedures. All model data were extrapolated to full-scale ship values using Schoenherr's friction coefficients with an addition of 0.00035 for correlation al-lowance. The tanker model was tested at the loaded and ballast condition;

the cargo liner

model at the loaded condition. Figures 33 and 34 show the performance predictions for the tanker and the cargo liner. Table V compares the results of the propulsion tests with the ship

11.5 15 155 15 16.5

Speed in knots

Table V

Contra-rotating propellers better (+) or worse (-) than conventional screw

pro-pellers. Propeller RPM Conventional screw Contra-rotating propellers

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Tanker Cargoliner Loaded condition Speed knots Loaded condition Ballast condition 14 + 5.1 % 14.5 + 5. 1 % + 6. 5 % 15 + 5.3 % + 8.2 % 15.5 + 4. 8 % + 8. 4 % 16 + 4. 2 % + 8. 2 % 16.5 + 3.4 % + 7.8 % 17 + 4. 2 % + 8. 2 % 17.5 + 9. 6 % 18 + 4. 6 % 18.5 + 5. 4 % 19 + 6.4 % 19.5 6. 8 % 20 + 6. 6 % 20.5 + 6. 2 % 21 + 5.1 % 11.5 15 155 16 165 17 17.5 18 Speed in knots

Figure 33. Power and rpm curves for tanker. 17 175 18 130 120 7.; ta. MOO 90 L-2 10.0 0 0

-+