marhy 2014 conference proceedings

ADVANCES IN COMPUTATIONAL AND EXPERIMENTAL MARINE HYDRODYNAMICS

VOL. 2 CONFERENCE PROCEEDINGS

(ISBN: 978-93-80689-22-7)

OF

International Conference On Computational and Experimental Marine Hydrodynamics (MARHY 2014)

DECEMBER 3 - 4, 2014 AT IIT MADRAS, CHENNAI, INDIA

ORGANISED BY

Department of Ocean Engineering

INDIAN INSTITUTE OF TECHNOLOGY MADRAS

&

The Royal Institution of Naval Architects

PKKText Box Editors P. Krishnankutty, Rajiv Sharma, V. Anantha Subramanian and S. K. Bhattacharyya

i

ORGANISING COMMITTEE

PATRONS

Prof. Bhaskar Ramamurthi, Director, IIT Madras Mr. Trevor Blakeley, Cheif Executive, Royal Institution of Naval Architects Prof. V. Anantha Subramanian, Head, Department of Ocean Engineering, IIT Madras (Chairman) Prof. S.K. Bhattacharyya, Department of Ocean Engineering, IIT Madras (Secretary) Prof. P. Krishnankutty, Department of Ocean Engineering, IIT Madras (Secretary) Prof. S.A. Sannasiraj, , Department of Ocean Engineering, IIT Madras Dr. R. Panneer Selvam, , Department of Ocean Engineering, IIT Madras Dr. Rajiv Sharma, , Department of Ocean Engineering, IIT Madras Dr. V. Sriram, , Department of Ocean Engineering, IIT Madras

TECHNICAL COMMITTEE

Prof. Manhar Dhanak, Florida Atlantic University, USA Prof. Raju Datla, Stevens Institute of Technology, USA Prof. Pierre Ferrant, Ecole Centrale de Nantes, France Prof. Kostas Belibassakis, National University of Athens (NTUA), Greece Prof. P. Ananthakrishnan, Florida Atalntic University, USA Prof. D. Sen, IIT Kharagpur, India Prof. C.P. Vendhan, IIT Madras, India Prof. Shekar Majumdar, Nitte Meenakshi Institute of Technology, Bangalore, India

ii

About the Conference

Marine hydrodynamics deals with flow around marine vehicles, such as surface ships, submarines, AUVs

and ROVs, and offshore structures, both fixed and floating ones. Some of the important topics are

marine vehicle resistance and propulsion, controllability, wave loads, wave induced motions, and energy

and ecology considerations. Correct understanding and application of hydrodynamics on marine vehicles and structures are vital in their design and operation.

Computational methods in marine hydrodynamic problems are applied to solve a wide range of

maritime applications. Significant progress has been made over the recent past towards the

development of the 'numerical towing tank' and 'virtual basin or cavitation tunnel'. Research and

development work is still ongoing to enhance their stability, accuracy, computational speed and its

integration into the overall design process. While the computational hydrodynamics can provide

important insights into physical flow characteristics and offers an economic way to investigate a range of

design options, it may still lack the accuracy to match results obtained in real-life experiments. This

obviously points to the fact that the computational methods do not replace the experiments completely.

The development of non-invasive flow measurement and visualization techniques such as particle image

velocimetry (PIV) has resulted in better understanding and quantifying the complex hydrodynamic

behavior such as wake in ship propeller region, flow around appendages and vortex shedding from

risers.

The aim of the conference and the pre-conference workshop is to provide a venue for disseminating

advances made in computational and experimental marine hydrodynamics and explore outstanding and

frontier problems in marine hydrodynamics for further research and applications.

iii

INTERNATIONAL CONFERENCE ON COMPUTATIONAL AND EXPERIMENTAL MARINE HYDRODYNAMICS

(MARHY 2014)

PAPER INDEX

iv

Paper No. Title/Authors 1 Effect of Structural Deformation on Performance of Different Marine Propellers

HN Das, ChSuryanarayana , B TejoNagalakshmi, P VeerabhadraRao 2 CFD Simulation Of Ship Maneuvering

K RavindraBabu, VF Saji, HN Das 3 Spatial-Spectral Hamiltonian Boussinesq Wave Simulations

E. van Groesen, R Kurnia 4 Validation Studies for the Scaling of Ducted Propeller Open Water Characteristics

A. Bhattacharyya, V. Krasilnikov 5 Hydrodynamic Analysis Of Podded Propeller Using CFD

NishantVerma , Om PrakashSha 6 Predicting the Impact of Hull Roughness on the Frictional Resistance of Ships

PA Stenson, B Kidd, HL Chen, AA Finnie, R Ramsden 7 Numerical Wave Tank Studies for Floating Wind Turbines

ShivajiGanesan, DebabrataSen 8 Sea Trials of a Water Jet Propelled High Speed Craft

K.O.S.R. Ravisekhar Radhakrishna, R. Panneer Selvam 9 Biomimetically Inspired Autonomous Ocean Observation System AquaBot

Prasad Punna,JagadeeshKadiyam, D.Gowthaman, R.Venkatesan 10 Numerical Study of Self-Propulsion andManeuvering Characteristics of 90t AHTS

Vessel, Praveen Kachhawaha,P Krishnankutty 11 Investigation on Effect of Skew on Natural Frequency for a MarinePropeller Blade in

Water Using F.E.M;Md. Ayaz J. Khan, Sanjay D. Pohekar, Ravindra B. Ingle 12 Effect of Environmental Loads on the Maneuverability of a Tanker

Deepti B. Poojari,Saj A.V, Sheeja Janardhanan, A R Kar 13 Heave Damping Characteristics of a Buoy Form Spar by CFD Simulation and

Experimental Studies; N. senthilkumar, S. Nallayarasu 14 CFD simulation and experimental studies on frequency andamplitude dependency of

heave damping of Spar hull with andwithout heave plate; J. Mahesh, S. Nallayarasu, S. K. Bhattacharyya

15 Reduction in Ship's Resistance by Dimples on the Hull? A Complementary CFD Investigation, S. C. Sindagi, Md. A. J. Khan, A.S. Shinde

16 Hydrodynamic Analysis Of Flapping Foils For Near Surface Vehicles P.Ananthakrishnan

17 Application of Direct Hydrodynamic Loads in Structural Analysis YogendraParihar, S. K. Satsangi, A. R. Kar

18 Ship scale CFD self-propulsion simulation and its direct comparison with sea trials results, Dmitriy Ponkratov, ConstantinosZegos

19 Wake Estimation: A Comparative Study Between Different Solvers Jai Ram Saripilli, Prasada Naidu Dabbi, Ram Kumar , Sharad S Dhavalikar, ApurbaRKar

20 Experimental and CFD Simulation of Roll Motion of Ship with Bilge Keel IrkalMohsin A.R. , S. Nallayarasu , S.K. Bhattacharya

21 Pitch and Heave Control of Swath using Passive Fins AzaruddinMomin, V. Anantha Subramanian.

22 Numerical Evaluation of Sloshing Pressure in a Rectangular Tank Fitted in a Barge Subjected to Regular Wave Excitation;Jermie J Stephen, S.A Sannasiraj, V Sundar

v

23 Numerical Investigation of Ship Airwake over Helodeck for Different Configurations of Hangar Shapes of a Generic Frigate;B Praveen, RVijayakumar, SN Singh,VSeshadri

24 CFD Analysis for the Configuration of the Hydrodynamic Depressor SenthilPrakash M N , Jithin P N

25 Numerical & Experimental Investigation on Semi-submersible Platform for Offshore Desalination Plant; AshwaniVishwanath, PurnimaJalihal

26 Behaviour Of Ship Under Sloshing, AbhijeetSajjan,A.P.Shashikala 27 Estimation of Submarine Hydrodynamic Coefficients from Sea Trials Data using EKF

Amit Ray, DebabrataSen 28 Assessment of Slamming Dynamics on High Speed Vessel

Deepak Bansal, V. Anantha Subramanian 29 Estimation of Hull - Propeller System Performance for Variation in Pitch- Diameter

(P/D) RatiosMd.Kareem Khan, Amit Kumar, PC Praveen, Manu Korulla , PK Panigrahi. 30 The Effect of Moonpool and Damping Plate on Damping Characteristics of Spar Hulls

Using CFD Simulation;Tom P.M. , S. Nallayarasu 31 Hydrodynamic Analysis of Self Installing Mono Column Wind Float During Transition

Phase, UtkarshRamayan, R. PanneerSelvam, NaganSrinivasan 32 Experimental and Computational Study of Lift - Based Flapping Foil Propulsion System

for Ships;Naga Praveen BabuMannam, Krishnankutty P 33 Investigation on the Effect of Fineness Ratio on the Hydrodynamic Forces on an

Axisymmetric Underwater Body at Inclined Flow;Praveen PC , Krishnankutty P, Panigrahi PK

34 Flapping Flexible Foil Propulsion Sachin Y. Shinde,Jaywant H. Arakeri

35 Estimating Manoeuvring Coefficients of a Container Ship in Shallow Water Using CFD AnkushKulshrestha , P Krishnankutty

36 Analysis and Design of Geotube Saline Embankment S. SherlinPremNisholdR. Sundaravadivelu , NilanjanSaha

37 Numerical and Experimental Determination of Velocity Dependent Hydrodynamic Derivatives of an Underwater Towed Body; Roni Francis , K sudarshan , P Krishnankutty & V. Anantha Subramanian

International Conference on Computational and Experimental Marine Hydrodynamics MARHY 2014

3-4 December 2014, Chennai, India.

2014: The Royal Institution of Naval Architects and IIT Madras

EFFECT OF STRUCTURAL DEFORMATION ON PERFORMANCE OF DIFFERENT

MARINE PROPELLERS HN Das, NSTL, Defence Research and Development Organisation, India

Ch Suryanarayana, NSTL, Defence Research and Development Organisation, India

B Tejo Nagalakshmi, NSTL, Defence Research and Development Organisation, India

P Veerabhadra Rao, NSTL, Defence Research and Development Organisation, India

ABSTRACT

Propeller geometry is very crucial for its performance and a little deviation in shape can cause changes in its

hydrodynamic performance. Hydrodynamic loading causes deformation to the propeller blades, which leads to

change in shape. The change in shape is particularly of concern when new designs use different composite materials

instead of conventional metals. Effect of this change of shape on hydrodynamic performance of a propeller is being

studied in the present paper. A five bladed bronze propeller from an existing ship is analysed to examine effects in

conventional propeller. Its open water efficiency was estimated for original and deformed shape. Pressure based

RANS equation was solved for steady, incompressible, turbulent flow through the propeller. Numerical solution was

obtained using Finite Volume Method within ANSYS Fluent software. FEM based solver of ANSYS Mechanical

APDL was used to make the structural calculations. Fluid-structure interaction was incorporated in an iterative

manner.

Additionally a five bladed composite propeller was analysed for hydrodynamic performance. Its deformation was

estimated under hydrodynamic loading for different fibre orientations. Hydrodynamic performance of the deformed

propeller was compared with that of the original one.

NOMENCLATURE

All Dimensions are in SI Units

1. INTRODUCTION

Geometry of propeller is very crucial for its

performance. A little deviation in its geometry may

largely influence the performance of a propeller. A

previous study reveals that some deviation in geometry

of a propeller during fitting into a ship caused variation

in its performance from its original design [9]. This

raised curiosity about performance of any propeller

when it is deformed under hydrodynamic loading.

Composite materials being more flexible, deformation

of composite propeller becomes more crucial and

hence its performance will be more interesting. The

present study concentrates on open water performance

of a metallic propeller vis--vis a composite one. At

first stage a five bladed metallic propeller was

analysed. CFD analysis was carried out for pre-

deformed geometry of the propeller to obtain

hydrodynamic pressure. This pressure was then applied

to the propeller to estimate its deformations. A FEM

code ANSYS Mechanical APDL was used for this. A

further CFD analysis was carried out with this

deformed shape to get the hydrodynamic performance

of the deformed propeller. This process was repeated

for few times to arrive at hydrodynamic load and a

compatible deformed shape of the propeller. At second

stage analysis of a five bladed composite

propeller is carried out in similar way.

2. LITERATURE REVIEW

Computation of viscous flow through propeller was

demonstrated in 22nd

ITTC conference in Grenoble,

France in 1998[10, 11 etc.]. In the last decade, Das et.

al. has carried out CFD analysis of contra-rotating

C D

E1,E2,E3

G12, G31, G23

J

Kt Kq k

n

p

Q

S

T

U Xt

Yt

Xc

Yc

12, 13, 23

Coefficient in k- turbulence model Diameter of Propeller

Youngs Modulus

Modulus of Rigidity

Advance Ratio

Coefficients of thrust

Coefficients of torque

Kinetic Energy of Turbulence

Revolution per second for

propeller Pitch

Torque of Propeller

Shear Strength

Thrust of Propeller

Free-stream Velocity

Tensile Strength in direction of fibre

Tensile Strength in direction normal

to the fibre

Compressive Strength in direction of

fibre

Compressive Strength in direction

normal to fibre

Dissipation rate of Turbulence

Kinetic Energy

Efficiency

Coefficient of Viscosity

Poissons Ratio

Density of Water.

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PKKText BoxAdvances in Computational and Experimental Marine Hydrodynamics (ACEMH 2014)Proc. of Conf. MARHY-2014 held on 3&4 Dec. , 2014 at IIT Madras, India - Vol.2 (ISBN: 978-93-80689-22-7)Editors: P. Krishnankutty, R. Sharma, V. Anantha Subramanian and S. K. Bhattacharyya

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propeller [3], hull-propeller interaction [4] and study of

propeller noise [5]. Many studies on static analysis of

propeller blades are available in literature. Stress

analysis for isotropic material by Sudhakar M [7] and

study for composite propeller by Y. Seetharama Rao

et. al. [8] is few examples.

Towards the end of last century, analysis of

composite propeller was started to exploit the

advantage of its flexibility [13]. In recent times,

Blasques et al. [14], Mulcahya et al. [15], Motley et al.

[16] and Liu and Young [17] reported study of

composite marine propeller for static deformation,

dynamic analysis and hydrodynamic performance.

3. GEOMETRY OF THE PROPELLER

3.1 METAL PROPELLER

A five bladed propeller is considered for the present

study (Fig. 1). Considering its diameter to be as D,

other geometrical parameters are expressed. The hub

diameter is 0.313D. Pitch ratio (p/D) of its blades at

radial section of 0.7D is 1.547. The propeller was

modelled using Catia V5 software.

3.2 COMPOSITE PROPELLER

Geometry of marine propeller is very

complex. An actual ship propeller, for which

experimental results are already available, is already

described in para 3.1. However, to ascertain the effect

of fibre orientation in a composite propeller, study is

done for a propeller with simple geometry which is

specially designed for this purpose.

A wing of uniform aerofoil section is chosen

to be propeller blade. This wing is placed over a hub of

1.314 m diameter. The length of wing is taken as

1.443m, which makes the diameter of the propeller as

4.2m. A constant pitch is maintained throughout the

blade. Pitch ratio (p/D) becomes 1.547 which was the

pitch ratio at a radial section of 0.7R of the actual

metallic propeller-blade. Blade thickness, thus, varies

in only one direction, from leading edge to trailing

edge and does not vary from root to tip. The maximum

thickness of blade is so decided that stress remains

within the allowable limit. This simple blade becomes

a wing with uniform cross-section. For analysis, the

blade is modelled in XY plane and its thickness run in

Z direction.

The geometrical description of the simple

propeller is given in Fig. 2. Surface model of the

propellers were made in CATIA V5, R9 software.

4. GRID GENERATION

4.1 GRID FOR FLUID STUDY

The surface model of propeller was imported from

Catia to ANSYS ICEM CFD 12.0. A suitable domain

size was considered around the propeller to simulate

ambient condition. A sector of a circular cylindrical

domain of diameter ~4D and length of ~7D was used

for flow solution. The sector of 72 was so chosen that

only one blade is modelled in the domain. Periodic

repetition of this sector simulates the whole problem. A

multi-block structured grid was generated for the full

domain using ICEM CFD Hexa module. The grid thus

generated was exported from ICEMCFD to ANSYS

Fluent 12.0 solver. Extent of domain and grid over the

blade is shown in Figs. 3 and 4. A grid with total 0.268

million cells were employed to descritise the flow

field.

4.2 GRID FOR STRUCTURAL ANALYSIS

The grid from only the blade surface was imported to

ANSYS mechanical APDL software. A view of

imported mesh is shown in Fig 5. Total 361 elements

(around 400 Nodes) were used over the blade. Fig. 6

shows grid and boundary condition for composite

propeller.

5. SETTINGS UP THE PROBLEM

5.1 FLOW SOLUTION

The problem was solved using the segregated solver of

ANSYS Fluent 12.0. In brief the code uses a finite

volume method for discretization of the flow domain.

The Reynolds Time Averaged Navier-Stokes (RANS)

Equations were framed for each control volume in the

discretised form. For the present solution,

STANDARD scheme is used for pressure and a

SIMPLE (Strongly Implicit Pressure Link Equations)

procedure is used for linking pressure field to the

continuity equation. The detailed formulation of

numerical process is given in Ref [6]. The

computations were carried out on an Eight Core Dell

Precision T7500 Workstation (64bit Xeon E5640

Processor @2.67 GHz, 4GB RAM, 64 Bit Windows

XP OS). The flow is treated as incompressible and

fully turbulent. Standard K- model has been used for modelling turbulence. The near wall turbulence was

modelled using standard wall functions and the free

stream turbulence has been prescribed as follows

K = 10

-4* U

2

The continuum was chosen as fluid and the properties

of water were assigned to it. A moving reference frame

is assigned to fluid with different rotational velocities

to simulate appropriate advance ratio. The wall

forming the propeller blade and hub were assigned a

relative rotational velocity of zero with respect to

adjacent cell zone. A constant uniform velocity was

prescribed at inlet. At outlet outflow boundary

condition was set. The farfield boundary was taken as

inviscid wall.

The following boundary conditions are used in this

analysis [Fig. 2]:

5

2

KC

2

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(i) Velocity Inlet, (ii) Outflow, (iii) Moving Wall

(iv) Inviscid Wall, (v) Periodic

5.2 DEFORMATION STUDY

Metal Propeller

The deformation of the metal propeller blade was

estimated using ANSYS Mechanical APDL 12.0

software. The solver used Finite Element Method

(FEM) for descritisation. For structural analysis, only

one surface of the blade was modelled. The pressure,

estimated from flow solution, was applied to this blade

surface. Fluents output of pressure distribution over two surfaces of blade, face and back, was written to a

file. A program picked up the pressure values from this

file and put to the nearest node points over the single

surface of the blade, to be used in Mechanical APDL

software. An four nodded shell elements i.e., SHELL

181, available with ANSYS solver were chosen for the

analysis. Propeller blade was considered as cantilever.

The root of the blade was considered as fixed,

restraining all degrees of freedoms there.

The blade was made of Aluminium Nickel Bronze,

which has Youngs Modulus 1011 N/m2 and Poissons Ratio of 0.34. A constant thickness of 0.1 m was

applied for the blade. This makes the volume of the

blade approximately same to the actual blade. Mesh

and boundary condition for FE solver is shown in Fig

5.

Composite Propeller

The deformation of the propeller blade was estimated

using ANSYS Mechanical APDL 12.0 software. One

surface of the blade was considered for analysis.

Geometry with mesh was imported from ANSYS

Fluent software (where CFD study was done). The

pressure over both face and back was written to a .cdb

file from ANSYS Fluent. The same .cdb file was read

in ANSYS Mechanical APDL 12.0 software to get the

loading over the blade. Properties of Graphite Epoxy

Composite Lamina with volume fraction 0.3 were

obtained from Jones [18]. Properties are given below.

Stiffness:

E1=207 GPa E2=E3=5.0 GPa G12=G31=2.6 GPa G23=2.87 GPa 12= 13=0.25 23=0.33

Strength:

Xt= 1035 GPa Yt= 41 GPa Xc= 689GPa Yc= 117 GPa S = 69GPa

Mesh and boundary condition for FE solver is shown

in Fig 6.

5.3 FLUID-STRUCTURE INTERACTION

The deformed shape of the propeller blade under each

operating condition was transferred to ICEM-CFD

software. After developing the actual blade around this

deformed surface, mesh was again generated. This

mesh was exported to Fluent and corresponding

operational conditions in terms of propeller rpm and

linear velocity was assigned in the solver. The

hydrodynamic results obtained from flow solution

represent the behaviour of the deformed propeller. A

new pressure distribution now develops over the blade

due to the change in geometry. The new load is again

exported to ANSYS APDL software for deformation

analysis. The original blade geometry is considered for

this. The process is repeated iteratively till the time

when pressure distribution does not change any further

between two successive iterations.

6. RESULTS

6.1 METAL PROPELLER

Analysis is carried out for the hydrodynamic

performance of the propeller. Open water

characteristics i.e., thrust (Kt) and torque coefficients

(Kq) as well as efficiency () were computed at different advance ratios (J), defined as

KT = , KQ =

= , nD

UJ (1)

According to the convention, thrust and torque are

expressed as non-dimensional quantities which remain

same under similar operating condition.

The propeller was analysed under a constant linear

velocity of inflow (U). Its rpm was varied to obtain

different values of the advance ratio. Analysis was

done for five advance ratios, ranging between 0.6 and

1.3. Pressure distribution over the propeller blade for J

= 0.6 is plotted in Fig 7.

Von-Mises stress over propeller blade for operating

condition J=0.6 is shown in Fig 8. The deformed shape

of the blade is shown in Fig. 9 and 10. The maximum

deformation is observed as 0.006428D. This

deformation is corresponding to an Advance Ratio of

0.6.

The open water characteristics for original and

deformed propeller geometry are shown in Fig 11.

Experimental results were available for a scaled down

propeller model [12]; so CFD results could be

compared with observations from experiment. From

Table 1, it is observed that change in hydrodynamic

efficiency due to deformation is very small (around

0.01).

6.2 COMPOSITE PROPELLER

Analysis was made for composite propeller to

get the deformed shape of the propeller blade with

different Laminates. Strength was checked from Tsai-

42 Dn

T

52 Dn

Q

2

J

Q

T

K

K

3

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Hill Criteria. Deformation and stress levels of the

propeller blade for different composite materials are

given in Table 4 to 6.

Propeller was run at different rpm and

advance velocities to produce advance ratio in the

range of 0.6 to 1.3. Amongst all the conditions, case

corresponding to 200 rpm and 30 Knots velocity of

advance gave the maximum structural loading. So all

the structural results corresponding to this case are

only reported here. Different thicknesses were tried out

and table 2 shows the deformation and stresses

developed within the propeller-blade. The minimum

required thickness is found out to be 80mm, where

stresses are within allowable limits and maximum twist

angle is 0.448. Table 2 shows that 80mm thick

laminate with 90/0/0/90/90 stacking satisfies the failure

criteria.

Final stacking sequence of composite layers is

arrived to 90/0/0/90/90 to avoid failure. Results for

other stacking sequences are given in Table 3. It is

observed that fibres need to be oriented at 90 at least

at the outermost layers to get the stress within

allowable limits. Other orientations of fibre lead to

higher deformation and stress which causes failure.

Table 4 shows that a suitably designed graphite- epoxy

composite laminate (90/0/0/90/90) could withstand all the load cases with 80mm thickness. Maximum

deformation of such propeller blade is observed to be

32mm with twist in blade as 0.4 (Table 2).

To keep the volume of the paper short,

hydrodynamic performance for only 80mm thick

propeller blade with graphite epoxy is reported. Open

water characteristics of deformed and pre-deformed

propeller with blade thickness 80mm are shown in Fig.

12. It is observed that its hydrodynamic performance

remains almost unchanged before and after

deformation. However, a meagre 0.85% improvement

is obtained after deformation for operation at J= 0.6.

The change in pressure distribution due to deformation

of blade marginally alters the stress level.

7. CONCLUSIONS

The present study indicates that capability of

computational methods to solve complex engineering

problem like fluid-structure interaction for a propeller-

flow.

CFD results agreed well with experimental

observations (Fig. 11) giving good validation of this

study.

Study shows that a bronze propeller is rigid

enough to hold its shape under operational conditions,

so that its hydrodynamic performance is not affected

due to structural deformations.

Shape change in composite propeller alters its

hydrodynamic performance. Further studies may be

carried out to examine if this can be used for

improvement of design.

8. REFERENCES

1. Edward V. Lewis, Principles of Naval Architecture Volume II, Published by The

Society of Naval Architects and Marine

Engineers, Jersey City, NJ, 1988

2. JP Ghosh and RP Gokarn, Basic Ship Propulsion, Allied Publishers Pvt Ltd., 2004

3. H.N.Das and Lt.Cdr.P.Jayakumar, Computational Prediction and Experimental Validation of the Characteristics of a Contra-

Rotating Propeller", NRB seminar on Marine

Hydrodynamics, Feb 2002

4. Commodore N Banerjee, HN Das and B Srisudha Computational Analysis And Experimental Validation of Hull Propulsor

Interaction For An Autonomous Underwater

Vehicle (AUV) Seventh Asian CFD Conference 2007, Bangalore, India, November

26-30, 2007

5. GV Krishna Kumar, VF Saji, HN Das and PK Panigrahi Acoustic Characterization of a Benchmark Marine Propeller Using CFD National Symposium on Acoustics (NSA-

2008), NSTL, Visakhapatnam, 22 - 24 Dec

2008.

6. ANSYS FLUENT 12.0 Documentation 7. Sudhakar M, Static & Dynamic Analysis of

Propeller Blade M Tech Thesis submitted to Andhra University, 2010.

8. Y.seetharama Rao, K. Mallikarjuna Rao, B. Sridhar Reddy, Stress Analysis of Composite Propeller by Using Finite Element Analysis, International Journal of Engineering Science

and Technology (IJEST), Vol. 4 No.08 August

2012

9. HN Das CFD Analysis for Cavitation of a Marine Propeller 8th Symposium on High Speed Marine Vehicles, HSMV 2008, Naples,

Italy, 22-23 May 2008

10. KN Chung, Fedric Stern and KS Min, Steady Viscous Flow Field Around Propeller P4119,

Propeller RANS/ Panel Method Workshop,

22nd

ITTC Conference in Grenoble, France,

1998

11. A Sanchez Caja, P 4119 RANS Calculations at VTT, 22nd ITTC Conference in Grenoble,

France, 1998

12. NSTL Internal Report on Hydrodynamic Model Tests For New Design Frigate (Open

Water, Self Propulsion & 3d Wake Survey

Tests); Report Number NSTL/HR/HSTT/221/2 November 2010

13. Lin G. Comparative Stress-Deflection Analyses of a Thick-Shell Composite Propeller

4

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Blade. Technical Report, David Taylor

Research Center, DTRC/SHD-1373-01, 1991

14. Blasques JP, Christian B, Andersen P. Hydro-elastic analysis and optimization of a

composite marine propeller, Marine Structures

2010; 23: 22-38.

15. Mulcahya NL, Prustyb BG, Gardinerc CP. Hydroelastic Tailoring of Flexible Composite

Propellers. Ships and Offshore Structures

2010; 5/4: 359-370.

16. Motley MR, Liu Z, Young YL. Utilizing Fluid-Structure Interactions to Improve Energy

Efficiency of Composite Marine Propellers in

Spatially Varying Wake. Composite Structures

2009; 90: 304-313.

17. Liu Z, Young YL., Utilization of Bend-Twist Coupling for Performance Enhancement of

Composite Marine Propellers, Journal of Fluids

and Structures 2009; 25: 1102-1116.

18. Jones R M, Mechanics of Composite Materials, Scripta Book Company, 1975

Fig 2 Geometry & Solid Model of Composite Propeller

Fig 1 Solid Model of Metal Propeller

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Fig 3. Extent of Domain and Boundary Conditions

for Flow Analysis

Fig 4. Surface Grid over Metallic and

composite Propeller

Fig 5. Grid, Boundary Conditions with Applied

Pressure for Structural Analysis (Metal propeller)

Fig. 6 Mesh and Boundary Conditions with Loading for

Structural Analysis over Composite Propeller

(SHELL 181 Element)

6

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(a) Face

(b) Back

Fig. 7 Pressure Distribution over Face & Back J=0.6

Fig. 9 Deformed Shape at J=0.6

Fig. 10 Deformed Shape at J=1.2

Fig. 8 Von Mises Stress(N/m2) over

Propeller Blade, J=0.6

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0

0.1

0.2

0.3

0.4

0.5

0.6

0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4

Co

-eff

icie

nt

of

Thru

st ,

To

rque

and

Eff

icie

ncy

Advance Ratio, J

Fig. 12Open Water Charateristics : Before & After Deformation

Kt After Deformation

Kq After Deformation

efficinecy After Deformation

Kt

Kq

efficiency

Table 1 Open Water Characteristics for Metal Propeller : Before and after Deformation

Before Deformation Deformed Difference

J kt kq

kt kq

Kt

(%)

Kq

(%)

0.6 0.471 0.097 0.4649 0.463 0.097 0.4562 -1.71 0.16 -0.00867

0.8 0.368 0.079 0.5905 0.366 0.081 0.5800 -0.54 1.24 -0.01039

1 0.268 0.063 0.6803 0.269 0.063 0.6835 0.66 0.19 0.003251

1.2 0.162 0.043 0.7139 0.164 0.044 0.7004 0.90 2.86 -0.0135

1.3 0.105 0.033 0.6692

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Table 2: Stress Level & Deformation for Composite Propeller Blade with different thicknesses

Propeller rpm=200; Material: GRAPHITE EPOXY, Laminate (90,0,0,90)

Thickness

(mm)

Maximum

Deformati

on

(mm)

Twist

Angle

() Extreme Normal Stress (MPa)

Extreme

Shear

Stress

(MPa)

Failure Condition

t x (min)

x (max)

y (min)

y (max)

z (min)

z (max)

xy Layer Tsai-Hill Index (Max)

100 16.30 0.240 -32.3 30.1 -328 316 -0.47 0.861 83.3

90 0.066

0 0.014

0 0.0844

90 0.554

90

21.78

0.315

-39.8

37.2

-405

386

-0.638

0.945

103

90 0.098

0 0.025

0 0.141

90 0.851

88

23.17

0.33

-41.6

39

-424

403

-0.679

0.966

107

90 0.107

0 0.028

0 0.157

90 0.932

85

25.51

0.3628

-44.6

41.8

-454

430

-0.748

0.1

115

90 0.123

0 0.034

0 0.187

90 1.073

80 31.96 0.448 -54.6 40.4 -613 454 -1.07 0.845 146

90 0.3465

0 0.0399

0 0.365

90 0.277

90 1.00

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Table 3: Stress &Deformation for Composite Propeller Blade with different fibre orientation

Plate Thickness=80mm; Propeller rpm=200

Maximum

Deformation

(mm)

Twist

Angle


Extreme

Shear

Stress

(MPa)

Failure Criteria

Laminate x (min)

x (max)

y (min)

y (max)

z (min)

z (max)

xy Layer Tsai-Hill (Max)

(90/0/0/90) 30.19 0.4126 -50.3 47.2 -513 482 -

0.883 1.07 129

90 0.157

0 0.0485

0 0.251

90 1.369

(45/-

45/45/-45) 54.41 0.355 -99.2 159 -422 332 -1.26 4.15 -182

45 1.2918

-45 0.159

45 33.0611

-45 6.076

(120/30/75

/30/

-15/30)

100.55 0.709 -174 105 -776 1110 -43.8 20.9 288

120 1.69144

30 5.985

75 97.741

30 62.224

-15 16.597

30 237.402

(302/902/30

2/902/302)

67.28 0.6322 -106 71.2 -302 803 -6.4 9.02 233

30 1.6964

30 1.0175

90 0.708

90 0.2454

30 12.581

30 24.437

90 2.663

90 5.9468

30 83.732

30 111.250

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Table 4: Stress & Deformation for Composite Propeller Blade , rpm 200

Laminate Thickness

(mm)

Maximum

Deformation

(mm)

Twist

Angle


Max

Shear

Stress

(MPa)

Failure Criteria

t

x

(min)

x (max)

y (min)

y (max)

z (min)

z (max)

xy Layer Tsai-Hill

(Max)

(90/0/0/90/90) 50 119.68 1.58

(+) -139 104 -1550 1090 -3.27 1.78 370

90 2.2733

0 0.613

0 3.574

90 1.9814

90 6.678

(90/0/0/90/90) 40 221.582 2.904

(+) -215 162 -2380 1620 -5.53 3.11 571

90 5.524

0 2.209

0 11.274

90 5.092

90 16.093

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International Conference on Computational Experimental Marine Hydrodynamics

MARHY 2014



CFD SIMULATION OF SHIP MANEUVERING

K Ravindra Babu, NSTL, Defence Research and Development Organisation, India

VF Saji, NSTL, Defence Research and Development Organisation, India

HN Das, NSTL, Defence Research and Development Organisation, India

ABSTRACT

International Maritime Organization (IMO) sets the standard for ship maneuverability. Naval ships needs even better

maneuverability. Accurate prediction of ships maneuverability is very important even at the early stage of design. Basic step towards finding the maneuvering characteristic of any vessel is to find the hydrodynamic derivatives. There

are many methods available for hydrodynamic derivatives prediction such as free running model test, captive model test

etc. However these methods are expensive and time consuming. Predictions based on semi-empirical or empirical

methods are not accurate. Whereas, accurate estimation of hydrodynamic derivatives is essential for evaluation of

maneuverability and directional stability.

RANS based CFD code are becoming popular as an alternative method to determine hydrodynamic derivatives. This

paper presents prediction of hydrodynamic derivative for static maneuvers using SHIPFLOW software. CFD results in

terms of hydrodynamic forces, moments and derivatives are compared with experimental results for a naval vessel and

showed good agreement.

1. INTRODUCTION Predictions of ship-maneuvering performance have

been one of the most challenging topics in ship

hydrodynamics. Due to the lack of analytical methods

for predicting ship maneuverability, maneuvering

predictions have traditionally relied on either empirical

method or experimental model tests.

Recently, computational fluid dynamics (CFD) based

methods have shown promise in computing complex

hydrodynamic forces for steady and unsteady

maneuvers. Significant progress has been made

towards this goal by applying Reynolds-averaged

Navier-Stokes (RANS) based CFD codes to static

maneuvers and dynamic maneuvers with generally

good agreements with experimental data.

The CFD simulations provide more insight into the

entire flow structure around the hull, and the

simulation results can be used to compute the forces

and moment acting on the hull and also to determine

hydrodynamic derivatives of the ship hull. Although

RANS methods are considered promising, many

difficulties associated with time accurate schemes, 6

DOF ship motions, implementations of complex hull

appendages, propulsors and environmental effects such

as wind, waves, and shallow water remain challenges. Captive model test and free running test require large

set up and are time consuming, whereas in practices, both time and cost are limited. Thus the execution of

extensive model tests for every ship is practically

beyond possibility. Results of semi-empirical or

empirical methods are not very accurate. RANS based

CFD are hence becoming popular for calculation of

derivatives. Present work employs a RANS based CFD

tool (SHIPFLOW 5.1) for the calculation of

hydrodynamic derivatives.

2. SIMULATION OF SHIP MANEUVERS Two simulations corresponding to straight line test and

rotating arm test have been performed using the

SHIPFLOW software for finding derivatives. An actual

ship has been considered for this purpose. Fig 1 shows

the model of the ship. Total length of the ship is 151.5m

with beam 17.71m. For this analysis 4.9m of draft was

used. Derivatives calculated using forces and moments

obtained by SHIPFLOW are compared with

experimental results.

Fig 1 Ship model

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For a bare model without propellers or rudders, the

Abkowitzs mathematical models for hydrodynamic forces and moment can be reduced to eqn (2.1) and

(2.2) by dropping the terms related to rudder angle ( ).

For the straight line test (static drift):

2

3

3

vv

v vvv

v vvv

X X X

Y Y Y

N N N

(2.1)

For the rotating arm test (steady pure yaw):

2

3

3

rr

r rrr

r rrr

X X X r

Y Y r Y rv

N N r N r

(2.2)

3. CFD MODELING To solve the flow around the hull two different

approaches, i.e. global and zonal approaches are

available in SHIPFLOW. A global approach means

that the Navier-Stokes equations are solved in the

whole flow domain. A zonal approach means that the

flow domain is divided into different zones based on

the flow characteristics inside. Global approach has

been used here. Experimental results are already

available for a model scale of 1:19.2 [5]. The present

simulations are also carried out for same model

scale, so that the results can be compared and

validated.

3.1 FLOW SOLUTION

The potential flow analysis was carried out under the

XPAN module of SHIPFLOW. This estimates the

wave resistance. However flow near the stern end is

completely viscous. Therefore a RANS solver

XCHAP is used to resolve viscous effects. XCHAP

has been used in the analysis. It is a finite volume

code that solves the Reynolds Averaged Navier

Stokes equations.

3.2 MESH GENERATION

The total number of elements generated was 858400.

The total number of panels generated was 2834 and

nodes generated were 3086. For potential flow

calculations, required mesh was generated by

XMESH module and for RANS calculations, grids

were created by XGRID module. The mesh was

generated automatically by giving XMAUTO in

XMESH. The type of the mesh used in XGRID was

medium. Figure 2 & 3 shows generated mesh on ship

hull body.

4 RESULTS

4.1 POST PROCESSING OF RESULTS

USING SHIPFLOW

Pressure distribution for Froude number of 0.23 is

shown in fig 4. The wave height variation along the

length of the ship is plotted. This is obtained from

the potential flow analysis done in SHIPFLOW.

The variation in the wave height at Froude number

Fig 3 Mesh

Fig 2 Grids of domain

around the ship hull

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MARHY 2014



(Fn) =0.23 can be clearly visualized from the fig 5

and 6 shown below.

Fig.5 Wave height along hull (from free surface) for a

velocity 1.646m/s

Fig.6 Free surface elevation for a velocity 1.646m/s

4.2 SIMULATION OF STRAIGHT LINE TEST

The velocity-dependent derivatives Yv and Nv of a

ship at any draft and trim can be determined from

measurements on a model of the ship, ballastard to a

geometrically similar draft and trim, towed in a

conventional towing tank at a constant velocity, V,

corresponding to a given ship Froude number, at

various angles of attack, to the model path shown in

fig 7

V = -V sin

Where the negative sign arises because of the sign

convention adopted.

A straight line test was carried out in a towing

tank to determine the sway velocity dependent

derivative. The test condition is simulated for a naval

ship model using SHIPFLOW software at different

drift angles. Hydrodynamic derivatives are

calculated using the forces and moments obtained by

SHIPFLOW.

Fig 7 Straight line test

Hydrodynamic Derivatives

Hydrodynamic derivatives are calculated using the

least square method using forces and moment

obtained by SHIPFLOW. These hydrodynamic

derivatives are compared with experimental results

and presented in Table 1.

Plots of Y vs. v and N vs. v are presented (Fig 8 and

Fig 9 respectively)

Table 1 Non-dimensionalised

sway force & yaw moment

y = 0.0030x - 0.0000

0.00015

0.00025

0.00035

0.00045

0.06 0.11 0.16 0.21

Y'

v

Yv'

Yv'

Derivative Computed

value

Experimental

value

-Yv 0.003 0.00285

-Nv 0.0092 0.017

Fig 4 Pressure Distribution

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Fig 8 Y vs. v plot

Fig 9 N vs. v plot

4.3 SIMULATION OF ROTATING ARM TEST

This is carried out to measure the rotary derivatives Yr

and Nr on a model, a special type of towing tank and

apparatus called a rotating-arm facility is occasionally

employed.

An angular velocity r given by u

rR

The only way to vary r at constant linear speed is to

vary R. The derivatives Yr and Nr are obtained by

evaluating the slopes at r = 0. Because of ship

symmetry, the values of Yr and Nr at the negative

values of r are a reflection of their values at positive r

but with opposite sign. This test condition is simulated

using SHIPFLOW software for different radius of

rotation. Hydrodynamic derivatives are calculated

using the forces and moments obtained by

SHIPFLOW.

Fig 10 Rotating arm test

Hydrodynamic Derivatives

Hydrodynamic derivatives are calculated using least

square method using forces and moment obtained by

SHIPFLOW. These hydrodynamic derivatives are

shown in Table 2.

Graph has been plotted between Y vs. r and N vs. r which shown in Fig 11 and Fig 12 respectively.

Table 2 Non-dimensionalised sway force

& yaw moment

Derivative Computed

value

Experimental

value

Yr 0.0206 0.026

Nr 0.065 0.069

Fig 11 Y vs. r plot

Fig 12 N vs. r plot

4.4 TURNING CIRCLE SIMULATION Introduction

Sea trial and free running model tests are

straightforward methods to obtain IMO

maneuverability criteria. However the free running

model test is not practical due to limitations of towing

tank and it is also expensive.

Computational simulations are advantageous than free

y = 0.0092x - 0.0001

0.0004

0.0006

0.0008

0.001

0.0012

0.0014

0 0.05 0.1 0.15 0.2

N

v

Nv'

Nv'

y = 0.0206x - 0.0015

0.0009

0.0029

0.0049

0.0069

0.0089

0.05 0.25 0.45

Y

r

Yr'

Yr'

y = 0.065x - 0.0049

0.0025

0.0075

0.0125

0.0175

0.0225

0.0275

0.0325

0 0.2 0.4 0.6

N

r

Nr'

Nr'

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MARHY 2014



running model tests for assessing vessel

controllability and maneuvering performance. Once

the hydrodynamic derivative are calculated using the

captive model test or theoretical method or using

RANS based CFD, almost any maneuver or ship

operation can be simulated without additional model

tests. The simulation model can be readily and

economically modified to determine the effect of

changes, such as increasing of rudder size.

The linear equations of motion have only limited use.

If a vessel is straight - line stable, they can be used,

in principle, for maneuvering prediction, if the

considered maneuvers are not too tight. If they are

tight, the result will not be accurate enough, as

contributions of nonlinear terms become significant

and they could no longer be ignored. If a vessel

is path-unstable, the linear system of equations

cannot be applied at all, as the solution will have a

tendency of unlimited increase and only nonlinear

terms could stop its growth.

A nonlinear system is derived from nonlinear terms

in the Taylor series expansion of usually it is

expanded up to the third power, as the terms of

higher order are small in most cases. In general,

which terms will be retained is determined by both

theoretical consideration and practical experience.

Numerical values of hydrodynamic derivatives come

from model tests with planar motion mechanism

(PMM), rotating arm, a free running model, empirical

formulas or RANS based CFD. There are numerous

formulations of the nonlinear equations, but the most

common are the cubic and quadratic nonlinearity.

The quadratic nonlinearity be used here because of

the availability of a complete set sample data.

However, cubic nonlinearity may also be used.

Simulation Program

The system of equations used here is given in ABS

Rule for Vessel maneuverability, which is a more

simplified form. The system of equation is integrated

with respect to time using MATLAB (2012 b)

software to get the trajectory for turning circle

maneuvers.

In the input block, the code will read the input data

such as rudder angle and hydrodynamic coefficients.

These input data will then be used in the process

block in order to calculate the hull, rudder and

propeller forces.

Hull modules are divided into three sub-blocks

called surge, sway and yaw sub-block. Surge, sway

and yaw acceleration are calculated using the

nonlinear equation.

The equation of motion was double integrated to

obtain the translation of motion in the x and y

direction. Fig 11 shows the predicted turning circle.

Fig 11 Turning circle plot

The steady turning diameter has been found to be

27.615m

Calculation of tactical diameter according to abs

guidelines

0.910 0.424 0.675SVTD STD

L L L

Eqn 4.1 shows the calculation of tactical diameter

Where,

TD = tactical diameter in m,

Vs = test speed in knots

L = length of the vessel in m, measured

between perpendiculars,

STD = standard tactical diameter in m

Tactical Diameter = 35.27 m < 5L. Hence IMO criteria

have been satisfied.

Table 3 gives the comparison between turning circles

calculated in different ways.

Table 3 Comparison of tactical diameter in ships length

Parameter ABS

guidelines

Present

result

Sea trial

result

Tactical

diameter in

ships length 5 4.47 3.8

(4.1)

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MARHY 2014



(m)

The difference between computational and sea trial

results may be attributed to the nonlinear terms of

hydrodynamic coefficients, which were neglected in

the present analysis. In spite of the inaccuracy of

present linear analysis, the predicted tactical diameter

qualifies the ABS criteria in a very similar way as the

actual sea trial result does .

5 CONCLUSIONS In view of the present state of art, successful analysis for computational estimate of Tactical

Diameter for ship, as reported in the present work

is very encouraging.

Velocity dependent variables were calculated using static maneuvers.

Stability condition was checked.

Turning circle maneuver has been simulated using ABS guideline for maneuverability. Results

agreed well with sea-trial observations.

As the results obtained are in good agreement with the sea-trial results, RANS based CFD tool

can be used for calculation of turning

circle/hydrodynamic derivative calculation at early

design stage to predict maneuvering characteristic

of vessel.

6 REFERENCES 1. American Bureau of Shipping, 2006, Guide for Vessel manoeuvrability, American Bureau of

Shipping.

2. Fossen, T. I., 1999, Guidance and Control of Ocean Vehicles, University Of Trondheim,

Norway.

3. Lewis, E. V., 1988, Principles of Naval Architecture, The Society of Naval Architects and

Marine Engineers, Jersey city, NJ.

4. SHIPFLOW 5.0 Users Manual, 2013, Flowtech International AB, Sweden.

5. NSTL Report Number NSTL/HR/HSTT/203 A Hydrodynamic Model Tests For P-15 Vessel-Mar 2008.

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3-4 December 2014, Chennai, India


1

SPATIAL-SPECTRAL HAMILTONIAN BOUSSINESQ WAVE SIMULATIONS

R. Kurnia, University of Twente, The Netherlands E. van Groesen, University of Twente, The Netherlands & Labmath-Indonesia, Email: [email protected], [email protected]

ABSTRACT This contribution concerns a specific simulation method for coastal wave engineering applications. As is common to reduce computational costs the flow is assumed to be irrotational so that a Boussinesq-type of model in horizontal variables only can be used. Here we advocate the use of such a model that respects the Hamiltonian structure of the wave equations. To avoid approximations of the dispersion relation by an algebraic relation that is needed for finite element/difference methods, we propose a spatial-spectral implementation which can model dispersion exactly for all wave lengths. Results with a relatively simple spatial-spectral implementation of the advanced theoretical model will be compared to experiments for harmonic waves and irregular waves over a submerged trapezoidal bar and bichromatic wave breaking above a flat bottom; calculation times are typically less than 25% of the physical time in environmental geometries. 1. INTRODUCTION The dynamic equations for incompressible, inviscid fluid flow have a well-known Hamiltonian structure in the surface potential and elevation as state variables [1, 2, 3, 4]. The dimension reduction is obtained by modelling instead of calculating the interior flow, as in Boussinesq equations. A spectral implementation makes it possible to treat the non-algebraic dispersion relation in an exact way above flat bottom; a quasi-homogeneous approximation makes it possible to deal with varying bathymetry. As a consequence, waves with a broad spectrum, such as short crested irregular waves in oceans and coastal areas, can be dealt with. By truncating the required Dirichlet-to-Neumann operator at the surface to a desired order of nonlinearity, nonlinear long and short wave interactions and generation can be calculated exactly in dispersion to the order of truncation. In our research over the past years, difficulties with spectral modelling when spatial inhomogeneities are present have been overcome by using Fourier Integral Operators leading to hybrid spatial-spectral implementations. Then waves above varying bottom, waves colliding to (partially) reflecting walls or run-up on coasts can be simulated. Using a kinematic initiation condition, a breaking algorithm (of eddy viscosity-type) has been implemented [5]. Waves can be initiated by a prescribed initial wave field or generated from given elevation at points or lines.

Comparing simulations with experimental data shows that the simulations are of high quality, typically the correlation with experiments is above 0.9, and are numerically efficient with calculation times typically less than 25% of the physical time in environmental geometries. In the present contribution examples of simulations for long crested waves will be shown: high frequency wave generation for harmonic and irregular waves running over a bar, and extensive frequency down-shift in bi-chromatic breaking waves above a flat bottom. With a good quality transfer function from wave elevation to wavemaker motion, the simulations can be used to design experiments in wave tanks in an efficient way [6]. 2. BASIC EQUATIONS Waves on a layer of incompressible, inviscid fluid can be described for irrotational internal fluid motion by variables depending on the horizontal variables only, namely the surface elevation and the fluid potential at the surface. The structure of the equation is special: it is a dynamical system as in classical mechanics, with a Hamiltonian structure. This was described by Zakharov [2] and Broer [3], and follows from Lukes variational principle [1] as was shown by Miles [4]. The equations are completely determined by the Hamiltonian (, ) and read (using partial variational derivatives denoted by ! , and ! )

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2

! = ! , ! = !(, ) The Hamiltonian is the sum of the kinetic energy (, ) and the potential energy (). Unfortunately cannot easily be expressed in the basic variables since it requires to solve the interior fluid potential (, , ) to determine the Dirichlet-to-Neumann operator ! = () at the surface: = 12 ! = 12 In [5] the operator is constructed up to 5th order in the surface elevation . Here we will only describe the 2nd order method since this case is especially simple. Introduce the tangential fluid velocity = ! for simplified notation. Then is a quadratic expression in , and it can be written as = 12 ()! where is some operator. In fact has a clear physical interpretation (when the gravitational acceleration is taken out of the integrand). In two limiting cases is easily determined to be (related to) the phase velocity. One limiting case is the shallow water equations, which are above bathymetry with depth () obtained for !" = () + . The other limiting case is the linear wave theory, for infinitesimal small waves above constant depth ! . Then the Laplace problem can be solved in the strip with Fourier expansion and becomes a pseudo-differential operator ! = (,!)()!"# /2 with = () !!"# the Fourier transform of and ,! = tanh ! . Note that ,! is the usual phase velocity that corresponds in linear theory with the dispersion relation ! = tanh ! . Above varying bottom () this generalizes in a quasi-homogeneous way to ,() = tanh ()

which is a Fourier integral operator. Even more so, by taking the total depth , = + (, ) , the expression ,(, ) = tanh (, ) leads to a second order correct approximation for nonlinear wave propagation above varying bottom. Observe that the limiting cases (shallow water and linear theory) are obtained in a consistent way. For higher order approximations the expression becomes a bit different but with a similar structure. For details we refer to [7]. These models are part of HaWaSSI software (Hamiltonian Wave Ship Structures Interactions) that has been developed over the past years. 3. SPATIAL-SPECTRAL IMPLEMENTATION Most important in the result above is that using the phase velocity operator provides the correct dispersive properties without any restriction on the wavelengths, a substantial improvement above other Boussinesq models. However, in order to retain this property in a numerical implementation, Fourier truncation has to be used; with finite elements or finite differences, the non-algebraic expression in has to be approximated by an algebraic expression, leading to restrictions on the wavelengths that are propagated with the correct speed. A technical problem arises in the use of (adjoints of) Fourier integral operators that appear in the explicit expressions of the right hand sides of the Hamilton equations. To facilitate the use of fast (inverse) Fourier transform, the spatial-spectral phase velocity (,(, )) has to be simplified. That can be done by a piecewise constant approximation, or by a interpolation method; see [5, 8] for more details. 4. TEST CASES In this section we will illustrate the simulation capacity of the HaWaSSI code for various different cases. 4.1 HARMONIC WAVE OVER A TRAPEZOIDAL BAR Beji and Batjess [9, 10] conducted a series of experiments to investigate wave propagation over a submerged trapezoidal bar. The experiments correspond to harmonic and irregular waves for either non-breaking, spilling breaking and plunging breaking cases. These test cases are very challenging since they involve a number of complex processes such as the amplification of the bound harmonics during shoaling process, wave breaking on the top of the bar and wave decomposition in the downslope part.

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3

The simulation for harmonic wave plunging breaking case has been shown in [5]. In this section we will show results for the non-breaking harmonic wave with frequency f = 0.5 Hz, wave height H = 2 cm. In Figure 1, the bathymetry is presented; the water depth varies from 0.4 m in the deeper region to 0.1 m above the top of the bar. In the experiment at seven position the wave height is measured: s1, s2, , s7 at positions x = 5.7, 10.5, 12.5, 13.5, 14.5, 15.7, 17.3 m. The measured wave surface elevation at s1 is used as influx signal for our simulation.

Figure 1: Lay out of the experiment of Beji and Battjes [10]. The locations of the wave gauges are indicated.

Figure 2: Shown are at the top elevation time traces and at the bottom, normalized amplitude spectra at positions s2 to s7 for the non-breaking harmonic wave case, the measurement (blue, solid) and the simulation with the HaWaSSI code (red, dashed-line). In Figure 2 we compare at all measurement points the elevation time traces in the time interval (60;95) s and the spectra of the measurements and simulations. It shows that the simulated surface elevation is in good agreement with the measurement: the wave shape is well reproduced and in phase during the shoaling process at up-slope, the wave amplification at the top and the wave decomposition at the down-slope. The corresponding normalized amplitude spectra describe the generation of bound harmonic at the upslope and

annihilation at the downslope. Good agreement between measurement and simulation is obtained, except for a slight underestimation of the amplitude spectra of third and fourth harmonics at s5, s6, s7. 4.2 IRREGULAR WAVES OVER A TRAPEZOIDAL BAR In this section we show results of propagation of non breaking irregular waves over the same trapezoidal bar.

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4

The input signal consist of irregular waves with JONSWAP type of spectrum with peak frequency f = 0.5 Hz, significant wave height Hs = 1.8 cm. For this test case the simulated surface elevation is also in good agreement with measurement, as shown in Figure 3 at the top. The wave shape is well reproduced and in phase, with a slight underestimation of the wave crests at s4 and s5. The generation of high frequency wave components due to nonlinear interaction occurs when the wave propagates over the bar in reasonable good agreement with measurement is shown in Figure 3 at the bottom; the generation of high frequency waves is observed as the appearance of a second peak frequency near f = 1 Hz. 4.3 BICHROMATIC WAVE BREAKING OVER A FLAT BOTTOM In this section we show simulation results for a bichromatic wave with initial steepness kp.a = 0.18, amplitude a = 0.09 m, periods T1 = 1.37 s, T2 = 1.43 s

over a flat bottom with depth D = 2.13 m. This test case is one of a series of wave breaking experiments that have been conducted in the wave tank at TU Delft and registered as TUD1403Bi6 [6]. In the experiment at six position the wave height is measured: W1, W2, , W6 at x = 10.31, 40.57, 60.83, 65.57, 70.31, and 100.57 m. The measured surface elevation at W1 is used as influx signal in our simulation. In this simulation we use a third order Hamiltonian model with extended wave breaking as described in [5]. In Figure 4 at the top we show the good agreement of the time traces of elevations of simulations and measurements at W2 to W6. The wave shape is well reproduced and the breaking position is well predicted; the breaking takes place at multiple positions starting at W3. In Figure 4 at the bottom we show the corresponding normalized amplitude spectra; high frequency wave generation and downshift in the spectra are observed.

Figure 3: Same as in Figure 2. Now for irregular waves with peak frequency f = 0.5 Hz and significant wave height Hs = 1.8cm.

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5

Figure 4: Same as in Figure 2. Now for bichromatic wave breaking over a flat bottom (TUD1403Bi6) .

Table 1: Correlation between simulations and measurements at measurement positions and the relative computation time (Crel) for the test cases.

No Case s2 (W2) s3 (W3) s4 (W4) s5 (W5) s6 (W6) s7 Crel 1 Harmonic waves over a bar 0.99 0.99 0.97 0.96 0.96 0.96 1.44 2 Irregular waves over a bar 0.97 0.96 0.93 0.89 0.88 0.89 0.78 3 Bichromatic wave breaking 0.98 0.94 0.92 0.90 0.86 - 1.89

In Table 1 we give quantitative information of the correlation and the computation time for the test cases that have been presented. The correlation between the measurement and the simulation is defined as the inner product between the normalized time signals. Deviations from the maximal value 1 of the correlation measures especially the error in phase, a time shift of the simulation. The relative computation time is defined as the cpu-time divided by the total time of simulation. Since the laboratory experiments are scaled with a geometric factor of approximately 50, the relative computation time for real scaled phenomena is a fraction of 7 of the test relative time; hence our simulations at geo-scale run in less than 25% of the physical time. All the calculations were performed on a desktop computer with CPU i7, 3.4 Ghz processor with 16 GB memory.

4. CONCLUSIONS The accuracy of the code as shown above makes it possible to use simulations in the design of experiments in wave tanks as was shown in [6] for a series of breaking waves of irregular, bi-chromatic and focussing type. Since in the present code waves are generated based on a time trace at an influx position, a high-quality transfer function is needed that transforms the influx signal to the corresponding wave maker motion. An extension to a fully coupled Hamiltonian-Boussinesq wave-ship model is presently being implemented as part of HaWaSSI.

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ACKNOWLEDGEMENTS We thank Prof. S. Beji for providing the experimental data over the bar. This work is funded by the Netherlands Organization for Scientific Research NWO, Technical Science Division STW, project 11642. REFERENCES 1. J. C. Luke. A variational principle for a fluid with

a free surface. J. Fluid Mech. 27, 395-397. 1967. 2. V. E. Zakharov. Stability of periodic waves of

finite amplitude on the surface of a deep fluid. J. Appl. Mech. Tech. Phys. 9, 190-194. 1968.

3. L. J. F. Broer. On the Hamiltonian theory of surface waves. Appl. Sci. Res. 29, 430-446.

4. J. W. Miles. On Hamiltons principle for surface waves. J. Fluid Mech. 83, 153-158. 1977.

5. R. Kurnia, E. van Groesen. High order Hamiltonian water waves models with wave breaking mechanism. Coast. Eng. 93, 55-70. 2014.

6. R. Kurnia, et al. Simulation for design and reconstruction of breaking waves in a wavetank. 2014. (to be published).

7. R. Kurnia, E. van Groesen. Accurate dispersive Hamiltonian wave Boussinesq modelling and

simulation for coastal wave applications. 2014. (to be published).

8. E. van Groesen, I. van der Kroon. Fully dispersive dynamic models for surface water waves above varying bottom, Part 2: Hybrid spatial spectral implementations. Wave Motion. 49, 198-211. 2012.

9. S. Beji, J. A. Battjes. Experimental investigation of wave propagation over a bar. Coast. Eng. 19, 151-162. 1993.

10. S. Beji, J. A. Battjes. Numerical simulation of nonlinear wave propagation over a bar. Coast. Eng. 23, 1-16. 1994.

AUTHORS BIOGRAPHY Ruddy Kurnia holds current position of Ph.D student at Department of Applied Mathematics, University of Twente, The Netherlands. His research focuses on modelling and simulation of accurate dispersive wave for coastal wave applications. E. van Groesen is professor of Applied Mathematics at the University of Twente, and scientific director of Labmath-Indonesia, Bandung, Indonesia. His main research area is the variationally consistent modeling and simulation of water waves, recently also including the interaction with ships.

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2014: The Royal Institution of Naval Architects and IIT Madras 1

Validation Studies for the Scaling of Ducted Propeller Open Water Characteristics

A. Bhattacharyya, Department of Marine Technology, NTNU, Trondheim, Norway; V. Krasilnikov, MARINTEK, Trondheim, Norway

ABSTRACT

This paper presents the results of validation studies for the open water characteristics of a four-bladed controllable pitch propeller operating inside two ducts of different designs. The results of numerical calculations by CFD are compared with model test results in terms of propeller and duct thrust, propeller torque and efficiency, and also in terms of velocity field downstream of propulsor. In order to quantify the scale effects on open water characteristics, CFD calculations are also carried out at Reynolds numbers corresponding to full scale conditions, and comparisons between the propulsor characteristics in model scale and full scale are presented for the range operating conditions from bollard to free sailing.

NOMENCLATUTRE

J : Advance Coefficient

KTD : Duct Thrust (N)

O : Open water efficiency D : Propeller Diameter (m)

KTP : Propeller thrust (N)

KQ : Propeller torque (Nm)

KT_Tot : Total thrust (N)

INTRODUCTION

The analysis of scale effects on open water characteristics of marine propellers is important to have accurate full scale power prognoses based on model test results. The flow around a rotating propeller is highly three-dimensional, and it involves high degree of swirl, adverse pressure gradients and, in some cases, flow separation and associated vortex shedding. For a ducted propeller, the propeller-duct interaction at different Reynolds numbers is of prime importance and has a strong influence on the corresponding thrust and torque characteristics and propulsor efficiency. The scale effects depend on the propeller and duct geometries as well as the loading conditions.

With advanced CFD techniques, robust flow solvers have been developed to resolve viscous turbulent flows, and they have become essential tools used in the marine industry to analyse complex flow around ship propellers. In this study, the scale effects on the open water characteristics of a four-bladed controllable pitch propeller operating with two different duct designs (a standard Wageningen 19A duct, and the Innoduct designed by Rolls Royce) have been investigated. The results of model tests performed at China Ship Scientific Research Center (CSSRC) and CFD simulations done with the commercial CAE software STAR-CCM+ are used for comparisons in model scale conditions, while full scale calculations are performed by CFD.

The strong duct-propeller interaction demands a separate scaling procedure for the open water characteristics of ducted propellers, where the simpler scaling methods developed for open propellers will not be applicable. In spite of the studies conducted earlier on scale effects on ducted propellers, the development of a universal procedure has not been possible due to complexity of interactions and geometry dependencies. In this study, it has been found that the trend of scale effects for the propeller working inside the two investigated ducts are similar. The detailed flow physics at different Reynolds numbers should be considered, in order to develop an efficient scaling procedure for the estimation of full scale open water characteristics of ducted propellers.

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BACKGROUND

The capability of performing efficient full scale simulations has made CFD a powerful tool for the investigation of scale effects of propellers. In most of the published CFD studies on scale effect on propeller characteristics, the RANS method is used with an isotropic turbulence model, the SST k- model (Menter, 1994) being the most common choice in the recent works. Most of the works are based on fully turbulent flow assumption (Stanier, 1998), Maksoud and Heinke (2002), (Krasilnikov et al, 2007), and only a few of them employ the recent extensions of the SST k- model to consider the laminar-turbulent transition flow regime (Mller et al, 2009).

Maksoud and Heinke (2002) performed systematic investigations into the scale effects on the open water characteristics of a Wageningen Ka 5-75 propeller fitted with a 19A duct at four values of propeller diameters and thrust loading coefficients.The increase of Reynolds number in full scale resulted in reduction of propeller thrust and increase of the duct thrust. Krasilnikov et al. (2007) presented a hybrid mesh generation technique for the steady RANS analysis of a series Ka propeller fitted with different duct designs using the SST k- model. This study shows that scale effects on the characteristics of ducted propellers depend on duct design, propeller design and loading conditions. Different ducts can produce different flow accelerations, which leads to variations in effective loading for the same propeller operating inside those ducts. This, along with different separation patterns on the duct in model scale and full scales, influences the magnitude of scale effect. The common conclusion from these studies is a larger reduction of propeller torque in full scale compared to that of an open propeller under equivalent operating conditions. This is due to the combined effect of the decrease of blade section drag and higher duct induced velocities on propeller. The Specialist Committee on Unconventional Propulsors of the 22nd ITTC (ITTC, 1999) have considered the three extrapolation methods proposed in (Stierman, 1984) for powering prognoses for the ships with ducted propellers. In this work, the most commonly used method 2 is followed, in the sense that the chosen approach implies that the resistance test is done for the naked

hull, while the open water tests are performed with the propeller operating in the duct.

TEST CASES

In this paper, flow analyses are performed for a 4-bladed controllable pitch propeller working within a standard 19A duct, using the RANSE flow solver implemented in STAR-CCM+. Comparisons of open water characteristics and induced velocities downstream of the propeller are made with model test results. The dependence of the propeller and duct forces on simulation methods and turbulence modelling is studied. Finally, the scale effects are investigated using CFD calculations of a full scale propeller, having the diameter 20 times of model scale and rate of revolution scaled according to the Froude number identity. The predicted changes with scale in propeller thrust and torque, duct thrust and propulsor efficiency for this propeller are compared with those obtained for the same propeller operating inside the Innoduct.

In Fig. 1 the profiles for the two ducts subject to investigation are shown with the mesh around the duct and blade tip.

CFD SIMULATION SET-UP

The propeller and duct are defined by their respective geometries which are used to generate

Fig. 1: Duct profiles including mesh

19A duct

Innoduct

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local grids in STAR-CCM+. Two different solution methods were used for the simulations. In the Moving Reference Frame (MRF) method the propeller is fixed while its rotation is taken into account by a local reference frame rotating at the desired speed. The stationary position of the duct is ensured by an appropriate setting of the zero rotation rate of the duct boundary. The additional acceleration terms from the rotating frame are incorporated into the modified equations of motions. This approach has been found to be suitable in the range of regular operation conditions (J= 0.2 to 0.6) where the interactions between moving and stationary parts can be approximated with sufficient accuracy by the quasi-steady solution. The Sliding Mesh (SM) model is used to resolve strictly the relative motion of stationary and rotating components and to account for all unsteady interactions. The bollard condition (J= 0) is a typical example where unsteady interactions are important, and where the MRF method is not sufficient to resolve the flow accurately. The simulation domains used with these two methods are shown in Figs. 2 and 3, and the details of mesh and solution settings are explained below.

One-block, One-blade passage:

(a) Only one fluid region whose rotational motion is considered in rotating reference frame. (b) Domain corresponds to one blade passage with periodic boundaries (c) Prismatic mesh in the boundary layers, and polyhedral mesh in the rest of the domain. (d) Mesh refinement by means of volumetric controls and local surface cell size near the leading and trailing edges of propeller and duct, in the region of tip clearance and in the propeller slipstream. (e) Methods used: MRF, steady. (f) Cell count is about 7.5 million per one blade passage.

Two-blocks, Whole domain:

(a) Whole domain divided into two fluid regions connected by the two internal interfaces. (b) Hexahedral trimmed cells in the outer fluid region, and polyhedral mesh in the propeller region. (c) Prismatic boundary layer mesh on the duct and propeller blade surfaces (d) Mesh refinement similar to one-block set-up. (e) Methods used: steady MRF to initialize the solution, and unsteady SM to iterate until convergence. (f) Cell count per blade passage is approximately the same as in the steady MRF method. (g) The time-accurate SM solution is done according to implicit unsteady algorithm, using the first-order temporal discretization scheme and time step corresponding to 2 degrees of propeller rotation.

For the model scale simulations with the 19A duct, solutions with the three different turbulence models have been compared, including k--SST model, k- realizable model, and Reynolds Stress model (linear pressure strain). For both two-equation models all y+ treatment has been used. The full scale simulations have been done using only the k--SST model.

The details of the near-wall mesh at the duct trailing edges are shown in Fig. 4.

Fig. 2: 1 block - 1 blade passage simulation set-up

Fig. 3: 2 blocks - whole domain simulation set-up

19A duct

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The boundary layer mesh of prismatic cells plays a very important role in providing adequate levels of the wall y+ function on simulated bodies, as well as in resolving accurately the velocity profiles in the boundary layer. In the present simulations, the values of wall y+ < 5 have been maintained on the blade and duct surfaces in both the model scale and full scale simulations (see Fig. 5 for the 19A duct). This has been achieved by reducing the total relative thickness for the prism mesh in full scale (0.002D, D being propeller diameter) compared to that in model scale (0.0025D) along with a higher stretching factor (1.4) for the prism layer mesh in full scale compared to that in model scale (1.2) The number of prism layers (20) has been same at both scales.

In the course of the studies it was also confirmed that a sufficiently smooth transition between the prism mesh and core mesh in terms of cell size change is essential for achieving physically correct flow picture, in particular, in the zones of larger velocity gradients, such as duct trailing edge and tip clearance.

The test calculations show that, in full scale simulations, one can employ the high Reynolds near-wall resolution (wall y+ >30) without reducing the accuracy of numerical predictions. However, for consistency of analyses, in the present study both the model scale and full scale simulations were performed with low Reynolds near-wall resolution (wall y+




Turbulence Model J KTP KTD KQ

SST

k-

m

odel

0.01 0.361 0.375 0.072 0.26 0.329 0.218 0.067 0.60 0.263 0.075 0.056 0.94 0.131 -0.024 0.035

Rea

lizab

le

k- m

marhy 2014 conference proceedings

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