Date June 2009
Author Huijsmans, R.H.M., R. Bosland and J.M. Dijk Address
Deift University of Technology
Ship Hydromechanics Laboratory
Mekelweg 2, 2628 CD Delft
TU Deift
DeIft University of Technology
Numerical prediction of thruster-thruster
interaction
by
R.H.M. Huijsmans, R. Bosland and J,M. Dijk
Report No. 1621-P
2009
Proceedings of the ASME 2009 28th International Confe-rence on Ocean, Offshore and Arctic Engineering, OMAE
2009, May 31June 5, Honolulu, Hawaii, USA, ISBN:
978-0-7918-3844-0, OMAE2009-79744)
WELCOME FROM THE CONFERENCE CHAIRS
fi!e://E:\data\chair-welcome.htrnl
8-6-2009
OMAE2009: Welcome from the Conference Chairs
Page 1 of2
R. Cengiz Ertekin H, Ronald Riggs
Conference Co-Chair Conference Co-Chair OMAE 2009 OMAE 2009
Aloha!
On behalf of the OMAE 2009 Organizing Committee, it is a pleasure to welcome you to Honolulu,
Hawaii for OMAE 2009, the 28th International Conference on Ocean, Offshore and Arctic
Engineering. This is the first conference with the new name, which reflects the expanded focus of the
OOAE Division and the conference.
OMAE 2009 is dedicated to the memory of Prof. Subrata Chakrabarti, an internationally known offshore
engineer, who passed away suddenly in January. Subrata was the Offshore Technology Symposium
coordinator, and he was also the Technical Program Chair for OMAE 2009. He was involved in the
development of the OMAE series of conferences from the beginning, and his absence will be sorely felt.
OMAE 2009 has set a new record for the number of submitted papers (725), despite an extremely
challenging economic environment. The conference showcases the exciting and challenging
developments occurring in the industry. Program highlights include a special symposium honoring the
important accomplishments of Professor Chiang C. Mei in the fields of wave mechanics and
hydrodynamics and a joint forum of Offshore Technology', Structures, Safety and Reliability' and
'Ocean Engineering' Symposia on Shallow Water Waves and Hydrodynamics. We believe the OMAE
2009 program will be one of the best ever. Coupled with our normal Symposia, we will also have
special symposia on:
Ocean Renewable Energy
Offshore Measurement and Data Interpretation
Offshore Geotechnics
Petroleum Technology
We want to acknowledge and thank our distinguished keynote speakers: Robert Ryan, Vice President
-Global Exploration for Chevron; Hawaii Rep. Cynthia Thielen, an environmental attorney who has a
special passion for ocean renewable energy; and John Murray, Director of Technology Development
with FIoaTEC, LLC.
A conference such as this cannot happen without a group of dedicated individuals giving their time and
talents to the conference. In addition to the regular symposia coordinators, the coordinators of the
special symposia deserve many thanks for their efforts to organize new areas for OMAE. We also want
to express our appreciation to Dan Valentine, who stepped into the Technical Program Chair position
OMAE2009: Welcome from the Conference Chairs
Page 2 of 2
on very short notice, following Subrata's passing. We also want to thank Ian Holliday and Carolina
Lopez of Sea to Sky Meeting Management, who have done a great job with the organization. Thanks
also go to Angeline Mendez from ASME for the tremendous job she has done handling the on-line
paper submission and review process.
Honolulu is one of the top destinations in the world. We hope that you and your family will be able to
spend some time pre or post conference enjoying the island of Oahu. Whether you're learning to surf in
legendary Waikiki, hiking through the rich rainforests of Waimea Valley, or watching the brilliant pastels
of dusk fade off of Sunset Beach, you'll find variety at every turn on Oahu.
Mahalo nui ba,
R. Cengiz Ertekin and H. Ronald Riggs, University of Hawaii
OMAE 2009 Conference Co-Chairmen
MESSAGE FROM THE TECHNICAL PROGRAM CHAIR
____________
Welcome to the 28th International Conference on Ocean, Offshore and Arctic
-
Engineering (OMAE 2009). This is the 28th conference in the OMAE series
guided by and influenced significantly by our friend and colleague, Subrata K.
Chakrabarti. It was a shock for me to learn that he had passed away so suddenly;
all involved with this conference express sincere condolence to his family, friends
and colleagues (the sentiments echoed by all of us are eloquently expressed in
the dedication included in this program). It is a great honor for me to have been
asked to continue his work on this conference. I and our community will miss his
leadership and friendship greatly. Although this series of conferences was
formally organized by ASME and the OOAE Division of the International
Petroleum Technology Institute (IPTI), it was Subrata's skill and dedication to this
division of ASME that made this series of conferences the success that it has
Daniel T. ValentineTechnical Program Chair
OMAE 2009
been and is today.
The papers published in this CD were presented at OMAE2009 in thirteen
symposia. They are:
SYMP-1: Offshore Technology
SYMP-2: Structures, Safety and Reliability
SYMP-3: Materials Technology
SYMP-4: Pipeline and Riser Technology
SYMP-5: Ocean Space Utilization
SYMP-6: Ocean Engineering
SYMP-7: Polar and Arctic Sciences and Technology
SYMP-8: CFD and VIV
SYMP-9: CC. Mei Symposium on Wave Mechanics and Hydrodynamics
SYMP-lO: Ocean Renewable Energy
SYMP-1 1: Offshore Measurement and Data Interpretation
SYMP-12: Offshore Geotechnics
SYMP-13: Petroleum Technology
The first eight symposia are the traditional symposia organized by the eight
technical committees of the OOAE Division. The other symposia are specialty
symposia organized and encouraged by members of the technical committees to
focus on topics of current interest. The 9th symposium was organized to
recognize the contributions of Professor C. C. Mei. Symposia 10, 11, 12 and 13
offer papers in the areas of renewable energy, measurements and data
interpretation, geotechnical and petroleum technologies as they relate to ocean,
offshore and polar operations of industry, government and academia.
The first symposium, Symposium 1: Offshore Technology was always Subrata
Chakrabarti's project. It was typically the largest of the symposia at OMAE. His
exemplary work on this symposium provided the experience and guidance for
others to continue to develop the other symposia. Symposium 1 in conjunction
with the OMAE series of conferences is Subrata's legacy. The Executive
Committee has a most difficult yet honorable task of finding a successor to carry
on this important annual symposium in offshore engineering. We are all grateful
file ://E:\data\chair-message.htnil
8-6-2009
OMAE2009: Message from the Technical Program Chair
Page 2 of 2
for the inspiration and encouragement provided to all of us by Subrata.
Please enjoy the papers and presentations of OMAE2009.
Daniel 1. Valentine, Clarkson University, Potsdam, New York
OMAE2009 Technical Program Chair
OMAE2009: International Advisory Committee
Page 1 of 1
INTERNATIONAL ADVISORY COMMITTEE
R.V. Ahilan, Noble Denton, UK
R. Basu, ABS Americas, USA
R. (Bob) F. Beck, University of Michigan, USA
Pierre Besse, Bureau Veritas, France
Richard J. Brown, Consultant, Houston, USA
Gang Chen, Shanghai Jiao Tong University, China
Jen-hwa Chen, Chevron Energy Technology Company, USA
Yoo Sang Choo, National University of Singapore, Singapore
Weicheng C. Cui, CSSRC, Wuxi, China
Jan Inge Dalane, Statoil, Norway
R.G. Dean, University of Florida, USA
Mario Dogliani, Registro Italiano Navale, Italy
R. Eatock-Taylor, Oxford University, UK
George Z. Forristall, Shell Global Solutions, USA
Peter K. Gorf, BP, UK
Boo Cheong (B.C.) Khoo, National University of Singapore, Singapore
Yoshiaki Kodama, National Maritime Research Institute, Japan
Chun Fai (Collin) Leung, National University of Singapore, Singapore
Sehyuk Lee, SamsLlng Heavy Industries, Japan
Eike Lehmann, TU Hamburg-Harburg, Germany
Henrik 0. Madsen, Det Norske Veritas, Norway
Adi Maimun Technology University of Malaysia, Malaysia
T. Miyazaki, Japan Marine Sci. & Tech Centre, Japan
T. Moan, Norwegian University of Science and Technology, Norway
G. Moe, Norwegian University of Science and Technology, Norway
A.D. Papanikolaou, National Technical University of Athens, Greece
Hans Georg Payer, Germanischer Lloyd, Germany
Preben T. Pedersen, Technical University of Demark, Denmark
George Rodenbusch, Shell IntI, USA
Joachim Schwarz, JS Consulting, Germany
Dennis Seidlitz, ConocoPhillips, USA
Kirsi Tikka, ABS Americas, USA
Chien Ming (CM) Wang, National University of Singapore, Singapore
Jaap-Harm Westhuis, Gusto/SBM Offshore, Netherlands
Ronald W. Yeung, University of California at Berkeley, USA
OMAE2009: Copyright Information
Page 1 of 1
COPYRIGHT INFORMATION
Proceedings of the
ASME 2009 28th International Conference on Ocean, Offshore and Arctic
Engineering (OMAE2009)
May 31
- June 5, 2009' Honolulu, Hawaii, USA
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Proceedings of the ASME 2009 28th International Conference on Ocean, Offshore and Arctic Engineering
OMAE2009
May 31 - June 5, 2009, Honolulu, Hawaii, USA
ABSTRACT
Many vessels deploying offshore activities nowadays are dynamically positioned by multiple azimuth thrusters instead
of anchors. The multiple
propulsor set up, gives a considerable flexibility to work fast and accurate. Due to the fact that the thrusters are positioned relative close to oneanother their performance is influenced. Normally to quantify this influence and take into account in the DP control algorithm, elaborate experiments have to be performed. To optimize the results a robust numerical flow solver is developed to predict the interaction effects. The
program is used to optimize the effort put into these
experiments.
The developed propeller interaction model is a first order potential based panel method, which uses zero order doublets and sources panel elements. This method is selected to prove the main objective of this research that; Although the slipstream of a thruster has a very turbulent character the interaction can be modeled without taking the viscosity into account as long as an accurate distorted flow field behind a propeller can be predicted.
At the 2nd thruster the distorted flow field due to the 1 thruster is modeled by means of two wake field models; a linear potential wake model and an empirical turbulent jet model. Due to the intersection of wake and body panels at the 2" thruster, numerical instabilities occur at the collocation points. These instabilities are removed by applying a realistic vortex model instead of the analytic vortex model which has infinite velocities in the core. The second problem is to capture the divergent and subsiding character of a propeller wake field by means of a linear potential wake model. This problem is resolved by validating the region for which the results are still accurate.
From the results it is concluded that the thruster
interaction propeller model coupled to the turbulent Jet wake field yield accurate thruster interaction results. For the
linear potential wake field results are promising but R Bosland, Aliseas Engineering by Poortweg 12 2612 PA Delft The Netherlands J.M Dijk Allseas Engineering by Poortweg 12 2612 PA DeIft The Netherlands
OMAE2009-79744
Numerical prediction of thruster-thruster interaction
R.H.M. Huijsmans TU DeIft
Mekelweg 2 2628 CD Delft The Netherlands
adaptations are needed to improve the prediction of the divergent and subsiding character of the physical wake field.
Keywords: Thruster interaction, wake field development, panel methods.
INTRODUCTION
To develop a control algorithm for the DP system a thorough prior knowledge of the involved forces and the interaction of the different thrusters is essential for a good design and steady operation of a vessel. Thrust reduction due to interaction effects between thrusters except from an economical point of view is not directly an issue if it can be predicted. Nowadays prediction of DP interaction effects is based upon extensive experimental research which has been performed over the years. Resulting in empirical
formulas for the pre-design phase, forbidden zones and nozzles tilted from the ships bottom surface. Forbidden zones in the DP algorithm take care of the slipstream interaction effect of one thruster upon one another and tilted nozzles take care of the Coanda effect. These adaptations work quite well however because of its importance model tests are almost always performed to
validate.
In 1975 Wise & English [11] were the first to discuss the nature of interaction between thrusters based on experimental results. Van der Made & Bussemaker [4] in 1976 continued the research which was followed by extensive measurements of interaction effects presented by Lehri [5] in 1980. The first extensive numerical research was performed by Nienhuis [7] in 1992, in which it is assumed that the thruster slipstream behaves as a turbulent jet. His work is still the fundament of many research performed nowadays. The rapid increase in computational power and the development of sophisticated numerical software over the past years creates the opportunity to continue the research on these complex phenomena
numerically.
OBJECTIVE
Although the conditions under which the phenomena of thruster interactions occur are highly turbulent and viscous, the question is how important it is to take into account the viscosity of water to model these interaction effects. The main objective of this research is formulated as:
"A/though the slipstream of a thruster has a very turbulent character the interaction can be modeled without taking the viscosity into account as long as an accurate distorted flow field behind a propeller can be predicted."
To reach the research objective a thruster inviscid
model capable of predicting thruster interaction is
developed. The flow is modelled as being inviscid and irrotational with the focus on the development of the distorted wakefield
at the 2nd thruster due to the
1thruster.
NOMENCLATURE
Strength of doublet singularity element.
a
Strength of source singularity element.cl) Potential of a fluid.
Distance from the core of a vortex element to a point of interest (m).
N Number of body panels.
N Number of wake panels.
x/D Distance downstream of a propeller made dimensionless by propeller diameter. (-)
Advance ratio of a propeller(-) S Panel surface. (m2)
v Kinematic viscosity (kg m r,, Core radius (m)
Q Free stream velocity vector (mis)
F
Vortex strength.UI,,d Induced velocity by a vortex element with strength, F at distance r. (mis)
V Kinematic viscosity (kgm1s') THEORETICAL BACKGROUND
The general solution for potential flows over bodies submerged in a fluid are based upon the simplifications of inviscid, irrotational and incompressible flow. Potential flows can be solved in terms of integrals taken over the boundary
surfaces
of the flow
field, after selecting the correct fundamental solutions. Each of these fundamental solutionssatisfies the Laplace equation.
V295=O
(1) Due to the linear nature of the potential flow problem a superposition of fundamental solutions is possible to yield the overall solution. The solution of this superposition ofsolutions can be resolved integral so individually solutions
are not necessary.
To physically represent a submerged body in a
fluid using potential flow, a correct distribution of fundamental solutions needs to be determined and correct boundary conditions have to be set to solve the Laplace equation. The basic boundary condition to represent
submerged bodies is also known as the "no-leakage"
condition.
(2)
In case the water free surface is included in the fluid domain this adds a static and dynamic boundary condition. In the thruster interaction model these free surface effects are neglected because the thrusters are assumed to be underneath the ship and assumed to be far away enough
from the free surface so the influence is negligible.
Panel methods
Panel methods are a numerical implementation of the general solution for the Laplace equation. In the
development of a thruster interaction performance model a
first order panel method is selected. An elaborate description and discussion of how to develop a panel method can be found in Joseph Katz et al [2]. Although there exist higher order panel methods for flow around propellers as e.g. presented by Kinnas [9] and Vaz [10] it
was felt that for thruster interaction problems this
sophistication was not needed and therefore a first order
panel method was developed.
THRUSTER MODEL
The Kaplan 4-70 propeller with nozzle 19A from the Wageningen propeller series is selected as a benchmark geometry mainly because extended experimental open water and thruster interaction results are available to
validate the interaction model. The geometry is deduced from Kuijpers [3] and an initial grid is produced accordingly,
see figure 1.
Figure 1: Grid for the Kaplan 4-70 propeller
To introduce the lift on elements like the duct and the propeller blades a wakefield model is used to model the shedding vorticity from the trailing edge. Although this vorticity is entrained by the local velocity of the fluid, modeling it requires an update of the wake geometry every time step, which is computationally very expensive. For this reason a linear wake field is adapted, which leaves the propeller blade trailing edge with a constant pitch angle equal to the average propeller blade pitch angle, see Figure 2. The linear wake field is constant in time and space. From Joseph Katz et al. [2] it is concluded that this only introduces a small error on the total lift and drag of the profile.
Figure 2: Kaplan 4-70 propeller with linear wake field. The grid as shown in Figure 2 consists of 2700 body panels and 1800 wake field panels, which was determined to be a good panel distribution after grid dependence study. VERIFICATION OF THRUSTER MODEL
The thruster model is validated by means of comparison of the numerical results for the thrust coefficients with the open water model tests as performed by MARIN and published by Kuijper [3]. No comparison between measured and calculated torque coefficients is performed because the emphasis is on thrust prediction.
Openwater diagram thrustcoefflcients, Kt. K, panel method
0.8 K)experimer1talsoution
0.6 8 0.4 0.2
From the results it is concluded that the accuracy of the thruster model is sufficient to be used in a thruster interaction model.
THRUSTER INTERACTION MODEL
The interaction effects are included
in the model by a
disturbed inflow field at the position of the 2' thruster as a
result of the 1 thruster. While the 15t thruster is subjected to an uniform inflow field. The interaction is assumed to be solely from the l thruster upon the 2 thruster and not vice versa, from experiments [4] this assumption can be validated if the 2 thruster is at least more than 2 times the
propeller diameter downstream.
The principle of the thruster interaction model can be described as follows:
Determine geometry and relative position of the
2nd thruster comparedto the l thruster.
Determine influence coefficients 1st thruster upon
the 2 thruster.
Determine total disturbed inflow field at collocation
points 2 thruster.
Determine thrust as results of disturbed flow field. In which the 4 step is again performed by the thruster model.
To determine the total disturbed inflow field at the
collocation points of the 2 thruster two wake field models are developed and validated. The development of both wake field models is discussed in the subsequent paragraph.
Validation of each model is done by comparison with two
systematic series of thruster interaction experiments
performed by Lehn [5]. WAKEFIELD MODELS Potential flow linear wake field
The potential linear wake field is constructed with the wake panels shed form the duct and the propeller blades. The panels are entrained by the flow and because the singularity strength is already known from the solution of the thruster model the induced velocities at the collocation points of the 2 thruster can be determined. To develop this model some
numerical and practical problems are to be considered.
1. Number of grid panels
To calculate correctly the influence of the wake upon the 2' thruster depending on its relative position at least double the amount of panels as used for the thruster are used to describe the wake. The influence coefficients diminish with
increasing distance but at least a considerable part
Copyright © 2009 by ASME
0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 J(-)
Figure 3: Results verification thrust coefficients open water diagram vs. numerical model.
upstream and downstream of the 2id thruster needs to be modeled for con-ed results. In the thruster interaction model computational effort is reduced by defining two additional approximations:
The 2 thruster has no significant influence upon the 1 thruster if its relative position is at least 2
propeller diameters away.
Area of interest for which results of the interaction model are considered to be accurate is assumed to be between 2-3 times the propeller radius. Main
reason is that because a linear wake model is implemented the divergent and subsiding character
of a turbulent jet
exiting the thruster is notmodeled correctly for the far field. 2. Intersection of wake and body panels
When the wake panels of the 15t thruster intersect the body panels of the 2 thruster it is inevitable and uncontrollable that certain collocation points from the 2' thruster will be close to or on the edge of the wake panels of the 1 thruster. This introduces a problem which is directly coupled to the choice of singularity element.
Figure 4: Intersection of wake field panels 1 thruster with
body panels 2nd thruster.
The velocity induced by a doublet element placed on the wake panels is comparable to a vortex ring placed over the edges of the panel. The induced velocity of a vortex element is reciprocal proportional to the distance to the vortex core. If any collocations points of the 2 thruster are
close to or on the edge of the 15t thruster's wake panels,
induced velocities will peak.
Induced velocity by vortex a vortex ring:
F
UIfld =
2,r r
To get a first indication of the extend of this numerical instability at first the induced velocities at a random transverse plane (x/D=2.0) in the wake field are examined.
The results for the induced axial induced velocities of a thruster with a propeller radius of 0.5 m are shown in Figure 5.
Induced axial velocities In the slipstream, xID-2.O Vrnax = 2 2QQS ,sV 3 -1V,nx, = -29 ,ms ,e,adwsO:OOrn 2 - --0 1
Figure 5: Induced axial velocities in the wake field of the 1" thruster at transverse plane x/D=2.O
In Figure 5 it is clearly shown that instabilities occur at the evaluation points close to the vortices at tip and root of the propeller wake field. To evaluate these "instabilities" is very difficult because the physical flow does have a very
turbulent character especially at the tip vortices. Moderating
the excessive velocities by a vortex element can be modeled
by a
vortex element with a viscous core description.Applying a realist vortex model is in line with the research performed by Timme [6]. He introduced a vortex model not
only based on vortex strength but also on physical
characteristics as; size of the vortex core, the fluid viscosity and the elapsed time since the vortex was initiated.
-
i_
F
V.,,d(..XO )2ir r Where,
4vt
Factor that determines the maximum velocity induced by a vortex element.For practical implementation Timme [6] performed
numerical calculations to obtain the following relation for the size of the core, the core radius rm, as a function of viscosity and time,
r, = 1,25644
vt
(13)Using this mathematical description of the vortex "core", Timme [6] was able to predict the development of a vortex in time until it completely subsides. In the interaction model the vortex development is considered to be constant so only a con-ed core radius needs to be established. In figure 6-9 it is shown that this realistic vortex model is a very efficient way to remove these instabilities without changing the overall velocities too much. The proposed wake field model is considered to be suitable to determine thruster interaction
(12)
4 Copyright © 2009 by ASME
in the region between 2-3 times the propeller diameter downstream of the 1 propeller. Before the model can be applied the core radius needs to be determined. The core radius is considered to be best established by performing a simple tandem thruster interaction. By tuning the core radius in the model with experiments the physical vortex size can
be determined and further set ups can be
developed. The sensitivity to the grid spacing and the core radius should be taken into account.
Turbulent empirical wake field
The first extensive numerical and experimental research performed upon thruster interaction is done by Nienhuis [7]. He conducted numerous experiments to determine the axial velocities in the wake field of a thruster. From the results he developed a empirical model for the axial velocities in the thruster slipstream, based upon the assumption that the thruster slipstream behaves like a turbulent jet. The results are valid for low speeds only, ranging from J=[0-0.2]. No
rotational velocities are taken into account but the development of the axial velocities over the distance is. Making this wake model particular suitable to investigate thruster interaction over a whole range of 'D values.
The assumption that the slipstream of a propeller behaves like a turbulent jet makes it possible to describe it with 5 characteristic parameters by the theory of Schlichting [8]:
Maximum velocity Urn (mis)
Velocity at the slipstream centre line Ua. (m/s) The radial position of the maximum velocity rm () The radial position of the inner half velocity Rhi (-) The radial position of the outer half velocity Rh2 (-) The slipstream is divided into two zones; the initial zone and fully developed zone. For both zones the 5 characteristics can be coupled using a turbulent jet velocity profile
depending upon the zone, as shown in figure 10.
The initial developing zone is determined to be
there where the wake field
is still influenced by the presence of the thruster. Having a double maximum velocity peak and a decreased velocity hollow behind the propeller hub. The fully developed zone, is there where the velocity profile has a single peak at the centre line.A complete description of the coupling is given in Nienhuis [7]. The development of the turbulent wake field by Nienhuis in non dimensional distance x/D is shown in
Figure 11-14.
VALIDATION OF RESULTS
Validation of each model is as mentioned before done by
comparison with two systematic series of thruster interaction experiments performed by Lehn [5], for the Kaplan 4-70 thruster. The tests are the in-line tandem thruster setup as shown in figure 15 and the azimuth angle variation as shown in figure 16. Both tests are performed
underneath a flat plat so no free surface effects can interfere with the interaction results.
The in-line tandem test systematically varies the
dimensionless distance between the 1 and the 2 thruster. The dimensionless distance is determined by dividing the axial distance, x by the propeller diameter, D. Thrusters are tested for a dimensionless distance ranging from x/D= 0 to 30.
Figure 15: Set up thruster interaction in-line experiments. The variable azimuth angle thruster interaction test is
performed at constant dimensionless distance of x/D=3.0 and variable exit angle of the 15t thruster relative to the 2 thruster. The exit angle of the 1st thruster is systematically varied, a = 0 to 30 degrees.
x
Figure 16: Variable azimuth angle thruster interaction experiments.
DISCUSSION OF THE RESULTS Potential flow linear wake field
Because the linear wake field is only valid for the region x/D ranging from 2 to 3. The line interaction test is only
performed to establish the core radius and the slip of the propeller. After which the variable azimuth angle thruster interaction test is performed with variable slip, the results are shown in figure 17. From the results it is seen that with increasing azimuth angle the results became less accurate. This inaccuracy can be explained by the absence of divergence in the thruster slipstream using the linear wake
model.
Further although the propeller slip factor for dynamic positioning is more realistic to be in the range of
40%80d/o the slip is only taken 00/c and 20%. The main reason for this is that with higher slip factors the number of wake panels and computations involved to obtain accurate result heavy burdens on the computational power. This can be explained by the fact that by increasing the slip factor the wake panels are compressed closer to each other and subsequently more panels are needed to cover the same
region.
In the thruster interaction model additional
parameters for slip and core radius are included to adapt the wake field to yield a more realistic wake field in the region of 2-3 times the average propeller diameter.
Turbulent empirical wake field
Results from the Nienhuis wake field description are
compared with the experimental results of Erik Lehn [5] as described. Although the rotational velocities are not taken into account the results between the experimental results of Lehn [5] and the turbulent wake field model of Nienhuis are remarkable accurate. Proving the capability of the propeller model to predict thrust under disturbed inflow conditions.
Results are shown in figure 18 & 19.
However to develop the model for thrusters with different P/D values then 1.0 and higher speeds then J=0.2
requires additional experiments to determine the
characteristics of the turbulent jet, which is dependent upon the thruster and the operational condition. Implementing
this empirical wakefield model leads to a substantial
decrease in computational effort. CONCLUSIONS
Main objective
The main conclusion of this research project is that the results from experimental and numerical calculation correspond well. The interaction effect as a result of the slipstream interference can numerically be modeled without taking into account the viscosity. The emphasis for good interaction results is upon the determination of the wake field velocities. From the two wake field models developed, the empirical turbulent jet by Nienhuis is very fast and gives very accurate results.
Thruster model
The interaction results obtained by the turbulent jet slip stream also confirm that with the developed propeller model it is possible to process a distorted inflow field into a new thruster performance. Only the flow velocities at all the collocation points needs to specified.
Potential flow linear wake model
For a thruster performing in open water conditions the linear wake is an adequate model to shed the body vorticity into the wake by means of constant wake panels giving accurate results. In thruster interaction conditions however the linear wake model does not correctly represent the physical properties of the wake. The constant induced velocity profile and the absence of divergence and subsiding
character of the wake field velocities limits the application. However in
the region of interest of 2 to 3 times the
propeller diameter downstream results using the realistic
vortex model are good.
For an accurate performance prediction a considerable number of panels is required and with it the
computational effort, especially when the slip factor
increases.
Realistic vortex model
Intersection of panels in potential based flow problems result in numerical instabilities. These instabilities increases in number and absolute value with decreasing grid size. To remove these instabilities a realistic vortex model needs to be applied to represent the physical core radius for the vortex under consideration. This core radius has considerable effect on the results, if a to large core radius is
selected thruster performance will be over predicted because induced velocities are suppressed. A simple thruster interaction test is needed to determine the correct
core radius.
REFERENCES
Hess, iL. and Smith, A.M.O. "Calculation of non-lifting potential flow about arbitrary three
dimensional bodies." Douglas Aircraft Co. Report
No. E.S. 40622, California 1962.
Katz, J. and Plotkin, A. "Low speed aerodynamics 2' edition." Cambridge University press, New York
2005.
Kuiper, G. "The wageningen propeller series." MARIN, may 1992.
Made, A van der, and Bussemaker, 0. "Thrusters for dynamic positioning." Proceedings, Offshore Craft Conference, London 1976,
Lehn, E. "Thruster Interaction effects." NFSI
Report R-102.80, 1980.
Timme, A. von. "Uber die
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Nienhuis, U. "Analysis of thruster effectivity for dynamic positioning and low speed manoeuvring."
TU DeIft, Delft, 1992.
Schlichting, H. "Boundary Layer Theory." Mechanical Engineering, McGraw-Hill, 1979.
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T'ecnico, Lisbon, Portugal. November 2005.
Kinnas, S. 'A potential based panel method for the
unsteady flow around open and ducted
propellers. p18th Sympoium on Naval
hydrodynamics, 1991.
Wise, D.A. and English. J.W. "Tank and wind tunnel for a drill-ship with dynamic positioning control." 7th
Offshore Tedinolgy Conference, Houston, texas, 1975.
Flaure 6: Induced axial velocity. wake=40X40 and core=
Induced axial velocity
Figure 9: Induced axial velocity wake=40X40 and core=O 0.5
-0.5
Line plot axial induced velocity at centre line, yO
-02 0 0.2 0.4 0.6 08 Induced axial velocity (mis)
0.0 & Line plot representation of induced axial velocities at centerline.
0.5 E 0 U) CU N -0.5
Line plot axial induced velocity at centre line, yO
Figure 7: Induced axial velocity, wake=40X40 and core=0.05 & Line plot representation of induced axial velocities at centerline.
0.5
-0.5
Line plot axial induced velocity at centre line, y0
Figure 8: Induced axial velocity, wake=40X40 and core=0.1 & Line plot representation of induced axial velocities at centerline.
0,5 E 0 cc cv N -0.5
Line plot axial induced velocity at centre line, yO
2 & Line olot reDresentation of induced axial velocities at centerline.
7 Copyright © 2009 by ASME
-02 0 0.2 0.4 0.6 08
Induced axial 'eIocity (m/s)
-02 0 0.2 0.4 0.6 08
Induced axial eIocity (mis)
08
-0 2 0 0.2 0.4 0.6
10-0 4nner devecpm.rg zon
D -Rhi Rm
8 - Rh2
- Rj
- 2Rj
-10 0 2 Slipstream width (.J=0)initiI zone My developed zone
4 8 10 12 14 16
XRDH
Figure 10: Nomenclature for development of an empirical wake field by Nienhuis [7].
E
N
0.5
-0.5
Line plot axial induced velocity at centre lIne, y0
0.5 1 1.5 2
Induced axial elocity (mis)
Fioure 11: Njenhujs wake field velocities at xJD '2.0
Fioure 13: Nienhuis wake field velocities at xID =6.0
Line plot axial Induced velocity at centre line, y0
0.5 1 1.5 2
Induced axial eIocily (m/s)
Figure 12:Nienhuis wake field velocities at x/D =4.0
Line plot axial induced velocity at centre line, yO
0 0.5 1 1.5 2
Induced axial eIocity (m/s)
Figure 14: Nienhuis wake field velocities at x/D =8.0
9 Copyright © 2009 by ASME 0.5 -0.5 1---0.5 E 0 SC Cs N -0.5
Line plot axial induced velocity at centre line, yO
0.5 E 0 Cs N -0.5 0 0.5 1 1.5 2
0,9 0,8 0,7 0,6 0,5 0,4 0,3 0,2 0,1 0
Interaction thrusters x/D=3.O, angle variable.
Eaperimentat total thrust +-- Numerical total thrust, potential
core=0.025 & slip 0% Numerical total thrust, potential core=0.025 & slip 20%
Figure 17: Variable azimuth angle interaction test, P/Drl.O, )mO.03, x/D=3O.
1,00 0,80 0,60 0,40 0,20 0,00
Thruster interaction, x/D=3.O
- U- Numerical total thrust, wake field by Nienhuis -s-- Expenmental total thrust
Figure 18: Interaction results for tandem configuration with variable azimuth angle.
1,00 0,80 0,60 0,40 0,20 0,00
Thruster interaction in-line.
-
Numerical total thrust, wake field by Nienhuis*
Experimental total thrustFigure 19: Interaction results for in-line tandem configuration, with variable axial distance.
10 Copyright © 2009 by ASME
0 5 10 15 20 25 30
Azimuth angle
0 5 10 15 20 25 30
Azimuth angle, deg.
0 2 4 6 8 10
