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Prediction of High Reynolds Number Flow around Naval Vessels P. Bull (QinetiQ, United Kingdom) J. B. Verkuyl (MARIN, Netheriands) D. Ranocchia, A. Di Mascio, F. Di Felice (INSEAN, Italy) Cdr R. Dattola (Italian Navy, Italy) L. Merle, S. Cordier (Bassin d'essais des carenes, France) ABSTRACT This paper presents in multinational collaborative effort to develop numerical tools and acquire an experimental data base to enable the capability of predicting the flow around naval vessels (surface ship and submarines) at full scale. This effort concentrated on modelling complex geometries (appendages, shaft, struts...) and the mean action of the propeller. A large experimental data base with uncertainty estimates was obtained in a large tunnel facility with a significant difference in Reynolds number, in the towing tank and at full scale on mainly 3 hull forms. CFD development on two codes, one commercially available, the other in house, has concentrated on turbulence model evaluation and selection, as well as numerical uncertainty analysis. The validation effort based on the comparison of computed and measured velocity maps in the same conditions concludes that CFD is capable of modelling the flows considered with reasonable accuracy in the most complex case and with good accuracy for the bare hull case. Comparison with full scale data where free surface effects are present shows that CFD is able to capture the main flow features although more important differences appear. Further work needs to be performed in CFD development and analysis of the data base to make full ause of this data. INTRODUCTION Knowledge of the flow field (velocities, pressure, waves) around a ship is essential for the design of the hull, for example for resistance optimization. A detailed knowledge of the flow in different areas around the ship is also very important for the design of different appendages, fins, struts, sonar domes. Finally, the flow field into the propulsor is clearly one of the most important inputs into its design. Indeed, when hydroacoustic performance is important the design must be optimised in the correct flow field, that is a flow which is governed by the presence of the hull and its appendages, but also by the action of the propeller, and for surface ships by the presence of the free surface. The flow field around ships has been studied experimentally on model scale for a century now and engineers have found methods to make the results of experiment at model scale useful. However, hydrodynamicists are confronted with similarity problems between model and full scale due to different scale effects which involve mostly the Reynolds number discrepancy, but also hull surface condition (roughness). These effects lead to changes in the flow field around the hull, particularly close to the surface of the hull and in the propeller flow. Hence, empirical corrections are required when extrapolating experimental results from model to full scale. These corrections make possible the prediction of resistance, ship speed, shaft speed. However, no reliable method exist to extrapolate a measured model scale flow field, mostly due to lack of knowledge of the real full scale flow. New test facilities like large hydrodynamic tunnels are capable of generating flow field data at relatively large Reynolds number and represent a new tool to be used for flow studies. The most direct approach to gain knowledge of full scale flow is to measure it at full scale. This has been done very few times in the history of naval hydrodynamics because of the limitations in the instrumentation techniques, but also because of the cost associated with installing the equipment and running the experiments in a sufficiently controlled manner.
Flow computations are yet another way to gain knowledge on the flow around ships. Indeed, numerical methods which were limited to inviscid flows until recently are now progressing rapidly with the availability of computing power, and numerical methods are capable solving the Navier-Stokes equations which include viscous effects around complex geometries. The validity of these new tools still needs to be demonstrated, in particular for full scale high Reynolds number flows. Hence, we have three sources of information available, each with its own limitations: · model scale data can be obtained in quantity but is subject to scale effects, · full scale data is what we are interested in but it is not practical and costly to obtain in sufficient quantity and quality numerical data can be obtained in quantity but it requires development and validation The opportunity given by new test facilities and new computational methods to model in more details the flow around ships is the basis for the current project. Project goal This project was focused on two main objectives: 1. Computational Fluid Dynamics: Develop and validate a CFD package capable of modelling high Reynolds number flows around typical naval vessels. 2. Experimental Fluid Dynamics: Create an experimental database of measured flow parameters for the validation of the CFD packages, including at full scale. The approach followed to achieve these objectives was to proceed as follows: measure flows around selected hull forms at a range of model scale Reynolds numbers · develop instrumentation to measure flows at full scale · measure flows in the propeller region at full scale · evaluate existing computational methods for the prediction of model scale Reynolds number flows for the selected hull forms · develop enhanced methods for the prediction of full scale flows with propulsors · validate the capability by comparison with measured data Project organization This project has been conducted through the EUCLID Memorandum Of Understanding (MOW) between armament directors of the WEAO defence organization. The Common European Priority Areas (CEPAs) groups cover the different areas of technology. CEPA 10 covers hydrodynamics and underwater systems. The role of the CEPA groups is to identify and promote Research and Technology Projects (RTPs). This project is identified as RTP10.12 with the following title "High Reynolds number incompressible flow". RTP10.12 is a 3 year co-operative research project based on equal cost sharing between all countries involved. The 4 countries involved are France (lead), Italy, Netherlands, United Kingdom with their respective research institutes: Bassin d'essais des carenes, INSEAN, MARIN, QinetiQ The breakdown in responsibilities in this project were as follows: · The Bassin d'esssais des carenes is leader of the consortium. It is also responsible of the model tests in the hydrodynamic tunnel, for both the surface ship and the submarine. QinetiQ is responsible for identifying the more suited code and turbulence models. In collaboration with the other partners, QinetiQ is responsible for the validation of the codes, from the data obtained during the towing tank, water tunnel, and full scale test campaigns. INSEAN is responsible for the implementation and evaluation of turbulence models in an existing home made code. INSEAN is also responsible for the model tests in the towing tank MARIN is responsible for the preparation and the validation of the full-scale measuring device. MARIN performed the measurements during the full-scale trials. The different test campaigns which have been performed are described in table 1: Free surface Re Nbr Euclid Sub De Ruyter Alliance Partner _ r END Full scale _ s.o log MARIN 1 . EFD EFD Water Towing tunnel tank / = ~ . 1.0 & 0.54 & s.o 107 o.9 107 . 1~ 1 1i . ~ 1/ Bassin INSEAN _ Table 1: Data gathering plan . ~ 1 . QinetiQ INSEAN
Mode! test programme The test programme is described in table 2 Model configuration Towing Water tank tunnel 1 Alliance bare hull X X 2 Alliance, with ~ X appendages 3 Alliance, with append. X and rotating shaft _ 4 Alliance, with X X appendages and propeller 5 De Ruyter, with X appendages and propeller 6 Submarine with X appendages 7 Submarine with X appendages and propulsor Table 2: Model test programme Configurations 1 and 2 provide information concerning the effect of appendages on the nominal wake. Configuration 3 helps in the evaluation of the rotating shaft effect. Additional tests with propeller are needed in order to obtain by comparison (configuration 4), the effect of wake on propeller performance. The study on the surface ship appendages of course depends on the ship used. Then, as a complement, the <<De Ruyter>> is tested in the towing tank fully appended and with the propeller, for one velocity only (Configuration 5) for comparison with full scale results. Submarine wake measurements are performed in the presence of the rear appendages (rudders and rear diving fins) for configuration 6 and with the propeller for configuration 7. Numerical program The numerical effort was confronted with the difficulty of developing a code in a collaborative environment. The solution adopted was to follow a two track approach. The first track concerned the development and validation of a commercially available code. This code was selected among several potential candidates through a benchmark effort which constituted a specific task. A second track was to evaluate the potential of an "in house" code for which the source is available The different tasks associated with the CFD work for both codes were to identify turbulence model, develop a propeller actuator disk model, and validate on test cases HULL FORMS The hull forms used in this study were selected in order to provide all the flow features present on current naval ships for both surface ships and submarines. For the latter, the collection of full scale data was not envisioned. For surface ship the key parameter was the availability of the ship for trials and the ability to circulate the hull form among partners. Two surface ships were finally selected: the NATO research vessel, the "Alliance" (figures 1, 2, and 3) and the Dutch frigate the "De Ruyter" (figures 4, 5 and 6~. The main particulars are summarised in table 3 Figure 1: The <<Alliance>> at sea Figure 2: <<Alliance>> underbody Figure 3: <<Alliance>> appendages and propeller
Characteristics De Ruyter Alliance Len th overall ~ 138.2 93.0 m g LBP 131.0 82.0 m Moulded beam | 14.8 | 15.2 m | | Drau kit, full load 4.58 5.2 m I g | Displacement loaded 4300 t 2920 t || Speed | 30 knots | 14 knots || Table 3: Ships' main particulars Figure 4: The <<De Ruyter>> at sea Figure 5: <<De Ruyter>> underbody Figure 6: <<De Ruyter>> appendages and propeller Figure 7: EUCLID submarine The submarine hull form, so called "EUCLID Sub", was invented to represent typical features of current conventional submarines with a relatively small L/D, a marked deck structure, a rather large sail, a SUBOFF type stern profile, forward dive planes, slender high aspect ratio vertical stern planes and more conventional low aspect ratio, thick stern dive planes. FULI, SCALE MEASUREMENTS Instrumentation Full scale measurements were conducted by MARIN (Netherlands) on two ships, the Dutch frigate the "De Ruyter" and the NATO research vessel, the "Alliance". The measurements performed concerned essentially 2D Laser Doppler velocimetry just ahead of the propeller plane. Pitot tube data was also gathered in the <<Alliance>> trials for boundary layer measurements. In addition to flow measurement general shipboard data was also measured: . . Shaft torque & RPM, rudder angle · Speed through the water DGPS: Heading, track, speed · Ship motions 6DOF · Weather and water conditions The LDV system consists in a single self contained system which can be adapted to a view port aboard the ship The 2D system is set up as to measure axial and cross flow components. A baffle system keeps both sides of the window and the outside of the optical lens immersed in water so as to eliminate the bias introduced by diffraction (figure 8). The optical system (emitting and receiving) is mounted on a traversing mechanism capable of inclining the head (typically athwartships) and raising the head radialy (figure 9). The displacement of the head allows sector
shape grids to be covered in cylindrical co-ordinates with maximum dimensions: 2 m radius, max angles +/- 30 deg. This system was calibrated and tested prior to installation aboard the ship. Figure 9: LDV optical head with traversing system << De Ruyter>> trials The"De Ruyter" trials were conducted in March 2000 as non dedicated trial. Only limited data could be gathered (about 24 furs) due to operational constraints of the ship. The velocity measurement were made for 2 speeds (9 and 14 knots) out of a single view port (figure 10~. ~ - ~ Figure 10: LDV grid on the De Ruyter <<Alliance>> Trials The <<Alliance>> was fitted with two view ports and two pilot tube hull fittings (figure 11~. The grid used for the measurements consist in about 130 points (13 angles and 10 radial distances) as shown in figure 12. Special grids refined in the shadow of A-bracket were generated. The pitot tube extends up to 30 cm out of hull. Figure 11: Windows and pilot tube on the Alliance
into account all the sources of error. The result is shown in table 4. Error source mean velocity Resolution ~- ~ . Figure 12: LDV Grid on the Alliance The "Alliance" trials were conducted over two dedicated test campaigns where the ship was chartered. The first campaign from one window was performed over 10 days with excellent environmental conditions (waves < 0,5 m, wind < BE 2, current < 1 knot) at speeds of 6 and 12 knots. Time available (about 6 hrs per test condition) enabled recording of three complete sets of measurements for each condition in order to assess the repeatability of the measurements. The second test campaign was performed in much harsher weather conditions which had a strong adverse influence on the LDV measurements. Generally, the results at 12 knots show much better reproducibility than at 6 knots, particularly for the axial flow component. The crossflow component accuracy is much more influenced by outside factors such as the ship motions, particularly for the second test campaign. Data analysis The large number of light bursts received were analysed to try to achieve the most accurate measurements possible. The procedure used to analyse the data is outlined below: Step 1: Determine quality of each data point Step 2: Omit all data points with poor quality Step 3: Determine the average flow Step 4: Maximum deviation from average flow Step 5: Omit data with poor reproducibility The repeat runs and the statistical analysis of the bursts used in each measurement (figure 13) were used to identify the quality of the data points and to discard poorer data points. An uncertainty analysis was conducted to try to assess the confidence level of the measurement by taking 0.2% 1.4% Processor2 0.1% Finite size mv 0.01% Sampling 0.25% Velocity bias 1.0% Alignment Total error Positioning Direction 0.3% 1.8% 0.01 mm 1 deg Table 4: Review of errors for a mean velocity of 4 m/see Full scale LDV results The results of L.DV measurements after analysis consists in two wake maps, on for each speed and represent the best estimate of the flow field based on the two test campaign (figure 14~. The final maps are composite maps which incorporate the data acquired for the different grids. The two maps show very similar patterns with a significant deficit on the upper part of the inboard sector due to the shaft bossing and inboard strut. Two interesting features which arise from the comparison of the wakes at the two speeds is on one hand the expected general increase in velocity deficit at the lower speed (Re effect), and on the other the hand the strengthening of the shaft bossing and inboard strut wake deficit. These two contradictory features are clearly present and may be the result of Froude dependent flow characteristics. Full scale pilot tube data This LDV data is supplemented with pitot tube data close to the hull surface, this for two speeds and two positions. Figure 15 shows an example of data acquired in the better weather conditions with an uncertainty band indicated by brackets which show good confidence levels. The velocity profiles are typical of thick boundary layers and can be used to check the capacity of CFD to model high Reynolds number boundary layer flows.
Good data Acceptable 20l-------- data ' Poor data ......................... , ., ~ . . ....... ....... · -:- - - --_'''' ...... _ . . 1.000 2.000 3.000 BSA1 Vel lm/s] O' ~ 1 tiff ....... 4.000 Figure 13: Statistical analysis of LDV data Y~IA6Vr~- Vrel ~ 6'~n4s -- .kleas~red date 4~"~\2 - "5 Idly ]4 ,~,;r~) JA Figure 14.a: Full scale measurements at 6 knots ~Y'X.~VrC4* 'art ].~:~ts - !~red data: AilIa''*~* FS Cr:~s lA: HI 2A Figure 14.b: Full scale measurements at 12 knots Figure 14: Full scale measurement results after final analysis Alliance Pitot tube location 11,12 knot measurement 0,0 0,1 0,2 0,3 0,4 0,5 0,6 0,7 0,8 0,9 1,0 Figure 15: Pitot tube measurements at 12 knots TESTS IN THE WATER TUNNEL Facility The tests were performed in the GTH (Grand Tunnel Hydrodynamique) of the Bassin d'essais des carenes in Val de Reuil, France. This hydroacoustic water tunnel facility has two test sections. The larger one (2.00x1.35xlO.OO) can accommodate whole ship models with flow velocities ranging from lrn/s to 12 m/s (figure 161. The flow can be seeded with microbubbles Cavitation nuclei) which are used as tracers for LDV measurements.
Instrumentation The measurements are performed using two 2D LDV systems to measure at the same position simultaneously (figure 17~. The LDV equipment on the side of the test section measures the axial and vertical components of the velocity. The LDV equipment under the test section measures the axial and horizontal transversal components of the velocity. The two measurements of axial velocity usually differ in high velocity gradient areas because of the elongated ellipsoidal shape of the measurement volume. The value to be used is that which is obtained with the measurement volume oriented near the normal to the direction of the gradient. The seeding system allows a data rate of about 1 kHz which varies with tunnel velocity. At least 500 samples were used per grid point in the analysis of the data. EUCLID Submarine in the GTH The experimental set up consists of a submarine modular test set up called "COLIBRI" which comprises a massive central cylindrical section to which the strut is attached and which includes the hydraulic connections to the motor in the model aft sections. The fore and aft sections were manufactured in GRP to the hull, deck and appendage. Given the two extreme velocities which were planned, two motor drives were used: an electric motor for low speed and a hydraulic motor for the higher speed. The EUCLID submarine was fitted in the large test section of the GTH, and LDV measurements have been performed, mostly 3D, at over 20600 points for 5 planes during this test campaign: · far upstream of the propeller, · in the middle of ship, · upstream of appendages, · upstream of propeller, · downstream of propeller. An XI n Figure 16: GTH facility ~ it' Figure 17: LDV system for the EUCLID Sub measurements Figures 18 and 19 show examples of axial wake maps in the propeller plane for the two extreme tunnel velocities. Both maps reveal the typical flow features downstream of the stern planes of a submarine with stronger wake deficit between adjacent stern planes and lesser deficit in the wake of the fins, particularly in the case of the dive plane which is thicker. This is the result of the effect of axial horseshoe vortices created by the fins which draw higher velocity flow along the fins inwards radialy, and similarly draw lower velocity flow radialy outward. This vorticity is clearly put in evidence in the map in figure 21 which is a zoom downstream of a vertical tail fin root. Comparing figure 18 and 19, the effect of Reynolds number is apparent with larger wake deficit at low speed. This effect is clearly illustrated in figure 20 which is a point by point difference of the figures 18 and 19. Differences on the order of 8 to 15% of upstream velocities are put in evidence in the areas between fins. i;11 -TOO Figure 18: Nominal wake at 2 rn/s (P4) no OF
Alliance in the GTH .~ _ . . ~ red 4,iy D ~ 1DO Yam Figure 19: Nominal wake at 12 m/s (P4) 41 m n ff3 'Day O? ' :D Hi nss Os Q~i no it: tom Figure 20: Re number effect (P4) Bier 'fr '.~ ~ \4hirt irtenst' .90 0 so :Y t~) IDES O 1~ ~.n~ ~.as 0.~ Hi.] ~.12 ~~.~:4 H 10 4 hi' 1 n ~1 .d Figure 21: Zoom on rudder wake flow (transverse flow and axial vorticity A half model of the <<Alliance>> was built in GRP at scale 1/13th and installed in the large tests section of the GTH attached to a vertical splitter plate and the roof of the tests section. The model was built up to a waterline higher than design by 31.5 mm to reduce the perturbation from the tunnel roof boundary layer. In this case the tunnel flow is split in two streams, one runs in the section where the model is (starboard), the other behind the plate which extends fore and aft of the model (figure 22~. The reference velocity is measured in the section where the model is present. In this case also, two motorizations and dynamometers were used, one adapted to each velocity. The model was tested bare hull, with appendages (static and rotating shaft) and with the propeller operating to achieve the full scale thrust coefficient. The bilge keels were not fitted since they would not have been aligned with the flow in the absence of the free surface. A photograph of the model in the test section is shown in figure 24. The instrumentation is identical to the submarine case (figure 23), except that in this the measurements had to be performed in backscattering mode. The 5 planes investigated are shown in figure 25. ~:.A A. ~ a a ~ - . A ~ ~ ..A ~ . i Figure 22: <<Alliance>> in GTH test section = at. ... . . ~ i. i . a ... Figure 23: LDV System for the <<Alliance>> measurement
Figure 24: Photograph of the <<Alliance>> model in GTH 90 - 80 ~ , i __ 70 ~ , _ . 1 _ 60 - ~1 _: ~ m. ~ 40 30 - 20 10 O Planes 5 4 3 2 1 Figure 25: Definition of measurement planes in GTH In these different planes velocity profiles along normals to the hull were also measured with closely spaced points as well as wake maps for the different configurations and velocities, hence about 50 data sets. The profile measurements in the bare hull case are illustrated in figures 26 where two profiles are shown in each plane for the two tests velocities. The large differences in the profiles are due to the 3D nature of the flow. The Reynolds number effect accounts for the small modification of the profile which tends to reduce the boundary layer thickness as the Re increases, particularly in the after plane (figure 26.b). Most of the data concerns wake maps at different positions, for the 4 configurations and the two different velocities. This data is illustrated in figure 27 for the 9.9 m/s test velocity where the bare hull, appended, and propelled cases are shown. The bare hull data shows the 3D effect on boundary layer development. The appended case shows the effect of the shaft bossing and shaft line which extract low velocity flow from the boundary layer into the free stream. The data with the propeller active shows the presence of a asymmetric jet due to the non uniform propeller disk loading, and the effect of the rudder. 90 ~ 80 ~ 70 60 E50 40 ~ 30 ~ 20 ~ 10 ~ o 0,8 1,0 1,2 VxNref Figure 26.a: Axial velocity profiles in plane 1 - e: == O_ . ~ ~ 0,8 1,0 1,2 1,4 Vx/\/ref Figure 26.b: Axial velocity profiles in plane 5 Figure 26: Axial velocity profiles at 1.4 and 9.9 m/s The resolution of the wake maps is illustrated in figure 28 where the transverse components of the velocity vectors are plotted for all measured data points in plane P5, aft of the rudder. The propeller swirl is clearly identifiable and so is the rudder flow straightening effect. Uncertainty assessment Efforts to identify error sources and to quantify them were also conducted. One possible source of errors is the model geometry which was checked on 64 different points with a mean error of 0.06 mm and a standard deviation of 0.65 mm. The histogram is shown on figure 29. Another source of error is the actual position of the centre of the measurement volume which is due to the positioning procedure used. These errors and other LDV system equipment noise (about 0.3%) were identified and documented.
- o- 100 - _ -300 - ...... -400 - ...... .... o 500 0 -500 - 1 Coo: Figure 27.a: Bare hull (P 3,4) -100 - -200 - -300 - ~oo - n- -100 - 2no -300 - 1 2 : 10 it' o I. On. 1-' ,,,,,, Hi- ~ ~ Figure 29: Histogram of geomtrical deviations measured on the model Figure 27.b: Appended hull (P 3,4) ._ hi_ _ Figure 27.c: Appended with propeller (P 3,4,5) Figure 27: Axial velocity maps at 9.9 m/s Figure 28: Transverse velocity field in plane 5 for the <<Alliance>> in GTH (1.4 m/s) Figure 30: Positioning bias of the measurement volume from the bottom and side LDV systems TESTS IN THE TOWING TANK Facility The tests were performed in the INSEAN towing tank in Rome, Italy (dimensions are 476x13.5x6.5). The model fixed at the equilibrium sinkage and trim and LDV equipment are installed on the carriage. The test program included the following arrangements of the ship model: . · hull with appendages without propeller hull with appendages and propeller
The tests performed concerned both the <<De Ruyter>> and the <<Alliance>> with similar test program. Only the work on the <<Alliance>> is reported here. Instrumentation Measurements were performed using a Laser Doppler Velocimeter along five transverse sections identical to the GTH tests (figure 25~. Three dimensional velocity field measurements were carried out in two separate steps by means of two different optical configurations with the laser radii coming from below and one side, in order to measure the axial-transversal (XY) and axial-vertical (XZ) velocity components respectively (figure 31~. The number of samples used per point is between 200 to 400. Phase sampling was executed by using a rotary 3600 pulse/revolution encoder which supplied the actual propeller position with an angular accuracy of 0.1°. The encoder signals were processed by a synchroniser which provided the propeller angular position. In order to improve the Doppler signal processor data rate and to reduce the acquisition time at point, the tunnel water was seeded with loom Titanium dioxide (TiO2) particles, provided downstream the ship model by using a special seeding rake device. Seeding was systematically carried out at the beginning of each test day. ~~ ~ ~~ Id. Figure 31: LDV system in the towing tank Alliance in the towing Tank An towing tank model of the <<Alliance>> was built at the same scale as the GTH ( 1/1 3th). The same propellers were used. The test program called for measurements at 0.856 m/s and 1.427 m/s with and without propeller and appendages, with the propeller angular velocity of 3.14 rps and 5.2 rps respectively corresponding to full scale running conditions at 6 and 12 knots (see table 51. Tests were performed at the full load draught by locking the ship model at the dynamic trim. Measurement were carried out by arranging the hubs in place of the propellers. Upstream Velocity (m/s) . 0.856 1.427 _ Propeller velocity (rps) 3.14 5.2 Advance ratio l 1.058 1.065 Table 5: Test conditions 3 6.85 1 3 The measurement grids used during the tests are presented in figure 32 and the results of axial flow measurements in figure 33 for the case with propeller operating. These measurements are time averaged results. The wakes in plane 5 without and with propeller operating are shown in figures 34 and 35. The patterns are similar to those obtained in the GTH. rJ~--~ I'm.... Y:~ - I~¢ ~s --:~0 :. ~3~.~ ~ w.i~ 0~ Figure 34: Axial wake and transverse streamlines in plane 5 without propeller At_ Am_ ~ .~ Em. I~18 ~.:~ f.m O%~ lo. - 0." 0," 01` ! 0" As, at', 90~8 1_~r ~~ - Ia~ ~ pi Lh ~~al sux~ell~ Figure 35: Axial wake and transverse streamlines in plane 5 with propeller
Laser beams from below . , I i +~ +;; ;~4 |' ~ ~ ·~- ; · ~~ a ~ e Lr L~.ci~r harms nv .ci'n'? . +: :~: Sty: ~4 :;t::: I; +~ ~~ r4~ ~ ; ·,l e ~ _ ~ · r~ r~r- r~ _ ! ·'~ ^ V ~ _ ! ~ ~^J ~ h^J LJ J ~& J ~J ~ a~ ~ s ~ e ~ a ~_ ~ r _ _ _ b ~ .. · - ^ ~ ~ ~ ^ a ~ ~ ~ Ir ~ ~ ~ · r7 ]! ItItIl.~.~1.1~. .. r l~i Figure 32: Definition of the grid density on the <<Alliance>> in the towing tank U'Uoo _~ ~ ~ 9- - Cl.4O ~ ~lr O.S~ C .l~i 066 0.~; O.g~ O 8~' O.97 1.0~1 Figure 33: Mean axial velocities an transverse streamlines on the <<Alliance>> in the towing tank
Aside from the classical time averaging technique which gives the mean flow quantities, an ensemble phase averaged technique was implemented to capture the flow features which depend on the propeller revolution, and in particular the blade passing frequency. This type of analysis allows the identification of: · the morphology and the angular evolution of the propeller suction effect on the inflow; · the evolution of the tip vortices and their interaction with the hull wake, especially downstream the rudder. In this technique velocity samples are acquired when Doppler signal is detected on the corresponding LDV system channel. This process is repeated independently in the two LDV channels because it is experienced that Doppler burst detection is not necessary simultaneous. A slotting technique is applied because LDV samples arrive randomly and classical ensemble averaging is not possible. Therefore, each sample, tagged with the angular propeller position at the measurement time, provided by the encoder-synchroniser system, is rearranged inside N angular slots of constant width. Statistical analysis is performed inside each slot to obtain mean flow field and turbulence intensity information. The slotting parameter choice is rather critical for this kind of analysis [elk et al., 2000), where a compromise solution should be obtained between the need to increase the angular resolution, required to capture the velocity fluctuations (smaller slots), and to have an adequate number of samples inside the slot for the consistency of the statistical estimators (larger slots). In the present analysis, statistical evaluation was performed by using a slotting technique with 360 overlapped slots of amplitude ~=2° and weighted averaging by gaussian law. In such way 200 to 400 samples per slot (obtained with 30 seconds of time acquisition at a point and a data rate of about 700 samples per second), are collected. In addition angular resolution is adequate to describe flow regions with high gradients. The flow is then reconstituted as described in figure 36 to give at each time step a complete image of the flow. This allows the identification of the time dependent flow upstream of the propeller (figure 37) and the generation of instantaneous wake maps of axial velocity upstream of the propeller (figure 38) and down stream of the rudder (figure 39~. A 3D presentation of these figures in shown in figure 40. These images can then be compiled in animations which reveal the 3D unsteady flow features of the hull/propeller interactions. The data both ahead and more markedly downstream of the propeller shows the acceleration areas created by the action of the blades. However, the presence of the viscous wake is perceptible, and the asymmetry in propeller loading due to global shaft to flow inclination can also be noticed in the form of stronger accelerations for blade positions corresponding to higher local angles of attack. Hence, the whole complexity of this flow field is well described. Although CFD development and validation concern only steady flows, this type of data will be particularly useful in further stages of CFD development, namely Unsteady RANSE codes. Adolf. flak_ rO t t | ~ it _ _ ~~i~ ~ ~ -fit ~-~^ 7~-, Figure 36: Flow reconstitution ~,`~0 Figure 37: Flow velocities upstream of the propeller
NUMERICAL SIMULATION USING COMMERCIAL CODE ALL I, .DD .~ 0~ 78 0.~ ~ r:,` 0~ D.33 : 0= ~ 11 O.m . 4 X' . :10- -71:>O - 1 DO ~ ~ (mm) Figure 38: Instantaneous axial flow map in plane 4 with propeller Ian ,2m ha 1W _ 4 _ 400 330 'an08" ml. ·.8 1.= I1. 4 1.1 19 1 a:~. C 0. D.8 a.~ O.~ ace 1 w ~ 1~) Figure 39: Instantaneous axial flow map in plane 5 with propeller C 70 O.~6 ~ 81 O8. 4.9Z ~ Da ~ '.~ '. ~ ~ 20 Figure 40: Instantaneous axial flow map in plane 4 and 5 with propeller QinetiQ conducted numerical studies using a commercial code in order to establish the conditions under which satisfactory results could be obtained considering the geometries and the data gathered in this project. Code selection The first task has been to select of code through a benchmarking effort by computing the flow around a representative geometry The basic requirements for the codes to be tested were as follows: Standard k-e turbulence model with wall function Optional alternative turbulence models: Reynolds Stress or non-linear models, Pressure gradient wall functions, Near wall models Recommended second order accurate discretisation (alternative higher order discretisation possible) The range of grid resolutions required to test wall functions were 0.20, 0.40, and 0.8 million cells and 0.25, 0.50, and 1.0 million cells for near wall models The different CFD packages tested were CFX4, CFX5, and Fluent with different grid resolution: 500K to 2M cells, y+ from 1 to 100 . Turbulence mode] implementation The implementation of turbulence models concentrated on co equation: k-co (Wilcox two equations model) k-~-BSL (Menter blend between ce and c) · k-co-SST (Menter modification to reduce overproduction) Stress (Wilcox Reynolds stress model) Stress-c}BSL (Menter blend between ce and s) The turbulence models were implemented so that the near wall boundary conditions automatically blend between laminar sub-layer and logarithmic layer. Furthermore, the models had to be compatible with rotating coordinate systems. As the experimental program had not started yet, the SUBOFF case was chosen to be used as benchmark since it had already been used for similar exercises. Furthermore it is a representative test case (body of revolution with 4 fins, Re = 1.2x107. A further study on models for turbulence implemented in the enhanced version of CFX5 has
been carried out using the DARPA SUBOFF test case. The turbulence models that were used are: Eddy · Eddy viscosity transport (EVT) model (one Viscosity equation model) k-e (KE) model (Standard two equation model) k-o (KW) model (Standard Wilcox two equation model) k-co-SST (SST) model Venter modification to the Wilcox model to account for the over production of k and the sensitivity to the far field) Launder Reece Rodi (LRR) Reynolds stress model (Standard Reynolds stress model with £ k-e Model equation) Speziale Sarkar Gatski (SSG) Reynolds stress model (Reynolds stress model with £ equation and quadratic pressure strain relationship) Wilcox (WCX) Reynolds stress model (Reynolds stress model with ~ equation) Menter (MEN) Reynolds stress model (Menter modification to the Wilcox stress model to blend between ce in the near field and £ in the far field) 00,05 -0,1 - 0,05 - O - The results of four of the best performing models are plotted in figure 41 were the same flow features as seen in the EUCLID Sub measurements. -0.1' In order to compare the results more clearly, the difference between experiments and calculations at a given radius is plotted against angular position over a 90° sector for all four turbulence models (figure 42~. Only one turbulence model gives a good level of accuracy (+/- 2.5%), it is the Launder, Reece and Rodi RSM model. EUCLID Sub calculation Once the EUCLID sub geometry and the GTH tests parameters and boundary conditions were defined (Reynolds number, confinement), this geometry was considered for calculation using CFX5 at Re = 0.67x107 & 3.97x107. The turbulence models used systematically were the k-e & RSM. The grid generated is multibloc with O-O grid around appendages with a total of 1M to 4M cells. A view of the axial velocity contours at several positions along the model is presented in figure 43. Turning to the GTH tests, one comparison illustrating the capacity to model Reynolds number effect is shown on figure 44 where the axial velocity computed (RSM) and measured in the propeller disk at a given radius are compared for two Reynolds Number. Two important comments can be made here. First the two sets of data agree rather well and second the trend in Reynolds number is very well predicted. RSM -cm BSL RSM -~ Figure 41: Comparison of axial velocities at the stern of the SUBOFF model U-Axial, SUBOFF AFF- 4 r/Rmax=0.25, y+=30 ._ ~ ~ ~ ,~ . ~+ ~ ~=~/ 0 10 20 30 40 50 60 70 80 90 Thetn ~ =A ~~ ~.O~i=~;; ~ k-epsilon swf (SM Lrr swf ~-omega k-omega bsl k-omega sst ~ +Uv Uv Figure 42: Comparison of axial velocities computed and measured at the stern of the SUBOFF model ~ ~ . ~ i, . 1 1 1 Figure 43: Flow computed on the EUCLID Sub model
1 _ 0,8 <~ O,6 . id, 0,4 0,2 O- RSM lrr Model, EUCLID submarine r/Rmax=0.34, y+=150 _ ~ 0 20 40 60 80 100 120 | · Data (2 m/s) · Data (12 m/s) | Model (2 m/s) Model (12 m/s, Figure 44: Comparison of measured and computed data at the stern of the EUCLID Sub Validation on the "Alliance" After these very encouraging results on the EUCLID Sub, attention turned to the simulation of the <<Alliance>> flow in the GTH with the objective of validating the code. The overall process for validation is defined in three stages · Documentation · Verification · Validation The difficulty in comparing numerical results against experimental data for the purpose of validation is to have some knowledge of their respective uncertainties in order to be able to draw any conclusions as to the validity of a method. The uncertainties in the numerical predictions from simulations are broadly categorized into numerical uncertainties and modelling uncertainties. For a typical Computational Fluid Dynamics (CFD) method the numerical uncertainties consist of uncertainties due to the numerical solution of the mathematical equations. These include such items as discretisation, dissipation, iterative and grid convergence, local and global mass, momentum and energy conservation, numerical roundoff, etc. The modelling uncertainties are those due to the assumptions and approximations in the mathematical equations. These include such items as turbulence models, propulsor models, free surface, wall boundary conditions, etc. They also include uncertainties due to the incorporation of previous data, such as the fluid properties and 'constants' in the turbulence models. A five-stage process for grid, iteration and time convergence provides the information on the numerical uncertainty for the CFD code. . Grid design and identification of important parameters Convergence studies Determination of effects of explicit artificial dissipation (if used) Estimation of overall order of accuracy Order of accuracy and Richardson extrapolation The range of grid resolutions used in this analysis are as follows: · Bare hull: 1, 2, and 4M cells · Appended hull: 2.5, 5 and lOM cells It must be emphasized that the generation of computational grids for such complicated geometries remains a difficult and time consuming exercise. Despite the large effort which has been dedicated to mesh generation in order to obtain a satisfactory mesh, it must be recognized that the solution will depend on grid topology to some extent until the grid generation software incorporates the grid criteria in an automatic fashion such as in adaptive grids. An exemple of the grid resolution in plane 5 is shown in figure 45. In the grid and boundary conditions, the physical tunnel geometry was simulated to includes tunnel blockage and tests conditions: with/without propeller and shaft Figure 45: Numerical grid in plane 5 (10 Mcells) This numerical uncertainty analysis was conducted for 2 turbulence models: k-e and Reynolds Stress Models and for the two Reynolds numbers corresponding to the GTH tests. The results of the numerical uncertainty analysis are as follows: · grid convergence uncertainty for the unappended case is less than 1% on average with a peak of 1.5%; for appended cases the average uncertainty is 4% with a peak of 8.5%.
iteration convergence uncertainty is less than 1% for both appended and unappended cases. · uncertainty due to numerical parameters (background turbulence, coefficients etc) is assumed negligible. The uncertainty of the comparison error can be estimated by considering all of the contributions to the uncertainties for both the numerical simulation method and the measured data point. The uncertainty in the comparison error UE can be expressed as: U2 (3E ) U2 +(3E ) U2 U2 +U2 where Us is the simulation uncertainty and UD is the data uncertainty. If the simulation uncertainty is broken down into its constituent parts then: a,- \~ ~ \~,,6t ~~ ~ ~ ~~ ~ ~ f ~9~ -~.r~e I: =~.r~ Am with ~.~r^~ - GAIL ,'~130: 46.a: Water tunnel data (77::~ tempt. (£~ ::~ >{ ~T>:~ ~ 1~.~-.:f~ relI.- ~ :r~rntr~ x11~ `;~:tt ~st<:h ~F~>~;I~ ~~d ;.~l.~r 46.b: Computational result Figure 46: Comparison of axial data for the propeller inflow plane at 1.4 m/s UE = UD + USN + USPD + USMA where USN is the simulation numerical uncertainty, USPD is the simulation modelling uncertainty arising from previous data and USMA is the simulation modelling uncertainty arising from modelling assumptions. However, the objective of a validation exercise is to define an estimate for the unknown uncertainty in a modelling assumption within a numerical simulation method. Defining the validation uncertainty Uv as the combination of all uncertainties that can then be estimated as: U2 = UEUSMA = UD + USN + USPD _ . . ~ ~ .. a._ .. i_ _ __ _- ~ i' ~ N~ - f ~ i} f-) - ,:~ .. 7,l~s~ ~,~11 (~ 11 J ~ by- `X L~ ~ ~ ~ }] ~ :< Irk ~ ') ~ it ~~ beet ,~ I ~!?~ I l 47.a: Water tunnel data v~ ~l.~r~ir arms -a ~ ~ COLA ~ ~ t0.~ I ~ ~~11~~ . ~llJ~£~1~ Wry =~.:t ~rIplf]~l~:~-Y' ~1 173'.~1~:t: 47.b: Computational result Figure 47: Comparison of axial data for the propeller inflow plane at 9.9 m/s
This gives the key metric in the validation process. The quantity UVis the validation 'noise level' imposed by the uncertainties inherent in the measurements, the numerical solution and the previous data used in the model.Based on the different data reports, the measured uncertainty was estimated to be ~2% and the positional uncertainty estimated at to be on the order of 1% (especially near walls). Obviously the estimation of this value is in fact very difficult since it will depend on the measurement position, and orientation of measurement volume which is much more critical in some areas than others. However, this approach to be applicable can not reflect these details and a global estimate had to be chosen. Hence, for this exercise a validation uncertainty 2.5% for unappended, 4.7% for appended cases was used. Comparisons between the measured and computed axial velocity components are given in figures 46 and 47 for the inflow conditions to the propeller for the <<Alliance>> geometry for the two tunnel velocities (1.4 and 9.9 m/s). These figures show the extent of the Reynolds number scaling of the flow features achieved within the EUCLID 10.12 programme. The overall axial flow parameters are well predicted by the computational methods. Full scale computations The ultimate phase of the CFD work has been to compute the flow at full scale Reynolds number for comparison with full scale measurements on the Alliance. Although this cannot be regarded as a validation exercise since the boundary conditions are clearly different (free surface, etc...), it is a comparison which can be performed as a "reality check" to put in evidence differences which will have to be addressed in future work. The axial flow computed in the same plane as measured at full scale on the <<Alliance>> is shown on figures 48 with the full scale measurement grid superimposed. These figures can be compared to the full scale measurements on figures 49 (same as 14~. The comparison is quite encouraging considering the differences in boundary conditions. Since the calculations are performed without free surface, the 6 knot case should fare better (low Froude number). The trend in the velocity deficit on the upper left sector is opposite between computations and measurements, most likely due to free surface effects which create stronger velocity deficits at high speed. This may be due to a change in flow alignment of the bossing at higher En which creates a stronger wake. NUMERICAL SIMULATION USING INSEAN CODE In parallel with the application and validation of CFX, INSEAN applied its own code to the validation exercise. Code description The code solves the Reynolds averaged Navier- Stokes equations using a Pseudo-compressible formulation discretized using a finite volume cell centered method. The convective terms are solved using E.N.O. (Harten et al. 1983) approach with artificial dissipation (Jameson). The viscous terms are also solved at the center of the cells. Although not used in these calculations, the free surface can be modeled using either surface fitting (moving grid) or level set (fixed grid) methods. The code can solve multi-block grids with either one- to-one matching at block boundaries or general matching with local refinement. The resolution is based on a time marching algorithm with local time step and multi-stage Runge-kutta with residual smoothing. Turbulence models Different turbulence models have been implemented: · Algebraic: Baldwin-Lomax · 1 equation model: Spalart-Allmaras · 2 equations models: k-£ The one equation Spalart and Allmaras model was found to give good agreement with experimental data on the velocity field although the turbulence field is affected by grid density. Mesh Generation Again, mesh generation is one of the most time consuming task in the numerical process. It is also critical in the validation process since the solution depends on the quality of grid. Some examples of the grid details around the appendages of the <<Alliance>> are shown on figure 50. Propeller mode] The propeller model is identical in both QinetiQ and INSEAN computations. The action of the propeller (torque and thrust) is modeled by a force field which is distributed in the cells which are at the propeller position. The strength of the forces assigned to each cell is determined based on a propeller calculation which is performed bv a unsteady lifting body
potential flow panel (BEM) code. The grid for the <<Alliance>> propeller is shown in figure 51. The wake used in the computation is extracted form the RANSE calculation. The method is outlined in figure 52. Few iterations are required to converge to force field which reflects the unsteady operation of the propeller. An exemple of the discretized force field in the RANSE mesh is shown in figure 53 where the color correponds to the fraction of the cell volume assigned to constitute the propeller disk. .~q .: Verification and vali(lation In this section, only the results concerning the <<Alliance>> are presented. Similar calculations have also been performed for the <<De Ruyter>> but are not presented here. The computations were carried out fo the towing tank configurations (without confinement). The free surface however was not modelled. Hence the two different flow velocity cases differ only by the Reynolds number in the computation while the Froude number in the experimental data is actually different. As an exemple of the calculations performed, the results obtained for the appended hull without propeller are shown on figure 54. lift \r:rEr. Vre~ = Into 5~: oe1J~ ~ ~s~r A:~;~¢ :~S with ¢~nd^~: and pr~p~;ll~:r Figure 48.a: Full scale computations at 6 knots vx'~. vr~t _ ta:~s _ ~ OM cel]s _ ~.~.f~llr.r IL:~ne~ :~S wish ~Lpp~ndege~ D~:~ p~rop~:ll~r: Figure 48.b: Full scale computations at 12 knots Figure 48: Axial velocity computed at full scale (QinetiQ) iX,/Vre:~. 17~! = 6~s - ~.~Bured: data: 61~r:.~ :Fs Chide i' a=d 7A Figure 49.a: Full scale measurements at 6 knots ~ Ha. ~$6 ~5 l'~..~- c,'~,$>, c::~q ~ A ~ ~ d-. ~ ~ Amp. errs - ~~ ~ Pleasured data AlL-as:~-~ FS Cr d~ 1A arid 2A Figure 49.b: Full scale measurements at 12 knots Figure 49: Full scale velocity measured at full scale
Figure SO.a: View of the grid around the rudder ... . ~ . ~ ~ . ~ ~ . .~ _._=. __ . . . .. . ~ Figure SO.b: View of the grid around the shaft Figure 50: Grid topology around the appendages - _ _ _ _ _ _ _ _ ~ i ~ 1 ~ __ _ __ ~ ; ~ _ ~ _ .. . _ _ ~ ~ _ . i i __ I_ 1 _ ~ ~ _ ~ . 1 Figure 52: Flow chart for the propeller force field Figure 51: <<Alliance>> propeller grid Figure 53: <<Alliance>> force field discretisation Figure 54: Axial flow field on the unappended <<Alliance>> hull
The numerical uncertainty (U) was evaluated based on estimated values of grid uncertainty (Uh) and iterative uncertainty (Ui~) U =4Uh MU,, The iterative uncertainty was found to be negligible. The grid uncertainty was evaluated based on grid refinement studies and Richardson extrapolation. U = 6 =log _ h r~-~ Rogers (f25-f~ ) Where f represents any dependent variable and h, 2h and 4h the grid densities of the three different meshes used, each obtained by taking one cell out of two each time the grid was made coarser. Hence the refinement ratio r = 2 and the theoretical convergence order (c,) is also equal to 2. The uncertainty evaluation was performed on the most challenging case, which is the appended case with the propeller operating. The observed convergence orders U) for the two cases (lower and higher speed) are given below: 0=1.72 Fx =2.94+0.33xlO ~ Re=5.4xlO6 6=1.58 Fx = 2.~8 +0.31x10-4 Re = 9.0x106 where Fx is the computed axial force. For he grid tested, the uncertainty remains relatively high (about 10%) however, the variability of the uncertainty over the computational domain is large, with large values in local areas. The validation is performed by comparison with experimental data gathered in the towing tank for different axial position of the measurement planes as shown on figure 55 for the slow speed case where free surface effects are small. The comparison is fairly good for the most forward plane shown and deteriorates as the propeller action becomes stronger. However the uncertainty of the comparison is well within the estimated numerical uncertainty estimated in the case with propeller. CONCLUSIONS The RTP 10.12 project on high Reynolds number viscous flows, represents an extensive European effort to improve computational tools through the acquisition of a large experimental data base at model scale and at full scale. This data has been analysed to yield the best estimates of velocities fields on different hull shapes (2 surface ships and one submarine) with estimates of uncertainty. At the end of this program, the CFD codes have been validated for naval vessel flows with an accuracy which can be considered sufficient for engineering purposes. They also have demonstrated the capacity to capture the principal features of the flow at full scale in a realistic fashion and the agreement is very encouraging. Although the analysis of this data is not complete it has already contributed to significant improvements in the capacity of CFD to model the complex flow over an open shaft vessel with an operational propeller. Further improvements in the flow model will be required, in particular free surface and rotating blades simulation, to achieve better agreement with the full scale data. O" `, an, -. ~= ~ _ DD7 0.~ 0.m Dot X=0 B517 ~ I. Figure 55.a: Plane through shaft bossing Figure 55.b: Plane aft of the propeller on= v D lo _ 0.m Din 0.m OlD1 . .~ ~ 0.~5 y Figure 55.c: Plane aft of the rudder Figure 55: Comparison of numerical results INSEAN (right) with experimental data (left) on the Alliance, Re=5.4 106
REFERENCES Cenedese, A, Accardo, L, Milone, R. 195, "Phase sampling in the analysis of a propeller wake", International Conference on Laser Anemometry Advances and Application; Manchester, UK. Davis J.C., "Statistics and Data Analysis in Geology", 1973, John Wiley & Sons, New York, 1986. Di Felice, F. Felli, M, Ingenito, G., 2000, "Propeller wake analysis in non uniform inflow by LDV", Propeller Shafting Symposium, Virginia Beach, USA. Felli, M, Di Felice, F. Romano, GP, 2000, "Installed Propeller wake analysis by LDV: phase sampling technique", 9th International Symposium on Flow Visualisation, Edimburgh. Hoshino, T. Oshima, A, 1987, "Measurement of flow field around propeller by using a 3-component laser Doppler velocimeter", Mitsubishi Technical Review, Vol.24,No. 1. Briancon-Marjollet L., Cordier S., Laurens J-M., Raulo J., "Effect of wake scaling on the prediction of propeller cavitation", CAV'95, Deauville, France, May 1995 Cordier S., Legrand F., Pinard J-C., "Hull and shaft wake interaction", Propeller and Shafting 1997, September 1997 Pinard J.C., "Etude experimentale et numerique du sillage en amont d'une helice" Thesis, Ecole Central de Nantes, 1997 Cordier S., Descotte L., "Control of the Turbulent Wake of an Appended Streamlined Body", 23 Symposium on Naval Hydrodynamics, Val de Reuil,, France, 2002 Jessup, S. 1976, "An experimental investigation of viscous aspects of propeller blade flow", NIIT Departments of Ocean Engineering, report 76-6. Kobayashi, S. 1982, "Propeller wake survey by Laser-Doppler Velocimeter", 4th International Symposium on Application of Laser Anemometry to Fluid Mechanics, Lisbona. Stella, A, Guy, G. Di Felice, F. 2000, "Propeller flow field analysis by means of LDV phase sampling techniques", Experiments in Fluids, Vol 28, p. 1-10. HW Coleman, F Stern. "Uncertainties and CFD Code Validation" Journal of Fluids Engineering, Dec 1997 vol 1 19, pp795-803 PJ Roache, "Quantification of uncertainty in CFD". Annual Review Fluid Mech. 1997, vol 29 pp 123-160 HW Coleman, WG Steele. "Experimentation and Uncertainty Analysis for Engineers". 1989 John Wiley & Sons, Inc. DC Wilcox, "Turbulence Modelling for CFD", DCW Industries, 2000, La Canada, CA 9101 1 NC Groves, CW Jiang, YN Liu, "Turbulence at the Stern of an Axisymmetric Model With and Without Appendages" DTRC/ SHD-1355-03 January 1992 CFX5.5. 1 Solver Guide, CFX International, AEA Technology Engineering Software Ltd. Gemini Building, Fermi Avenue, Harwell International Business Centre, Didcot, Oxfordshire OX11 OQR "Fully Three-dimensional Reynolds Averaged Navier-Stokes Flow Calculations for Two Ship Models." PW Bull, SJ Watson, AJ Musker. Proc. of the 1990 SSPA-CTH-IIHR Workshop on Ship Viscous Flows. Research Report, Flowtech Int. AB No. 2, Goteborg June 1991. "A Comparison of Two Navier-Stokes Methods Applied to the Stern Region of the HSVA Tanker". AJ Musker, DJ Atkins, SJ Watson, PW Bull. Proc. of the 2nd Int. Colloquium on Viscous Fluid Dynamics in Ship and Ocean Technology, Japan September 1991. "A Multi-block Capability for Submarine Stern Flows." PW Bull, P Gallagher. Proc. of DSAC Seminar on Naval Computational Fluid Dynamics, Portsmouth, September 1992. "The Prediction of Nominal Wake Using CFD". AJ Musker, SJ Watson, PW Bull, CC Patis, C Richardsen. Proc. of 6th Int. Conf. on Numerical Ship Hydrodynamics. Iowa, USA, August 1993. "A Brief Study of the Effects of Simple Changes to the Turbulence Modelling for Ship Flow Predictions". PW Bull, AJ Musker, SJ Watson. CFD Workshop, Ship Research Institute, Japan, March 1994. "Grid Generation for Hydrodynamic Vehicles". PW Bull. Presented at CFD Community Club, Recent Developments in Grid Generation. University of Bristol. September 1995. "Generation of Grids for Viscous Flows around Hydrodynamic Vehicles". CC Patis, PW Bull. 5th International Conference on Numerical Grid Generation in Computational Field Simulation. Mississippi State University. April 1996. "The Validation of CFD predictions of Nominal Wake for SUBOFF Fully Appended Geometry". PW Bull. 21st Symposium on Naval Hydrodynamics. Trondheim, Norway. June 1996.
"Viscous grids for Hydrodynamic Vehicles". JA Chappell PW Bull 6th Int. Conf. on Numerical Grid Generation in Computational Field Simulation. Greenwich, UK, March 1998 "The Scaling of High Reynolds' Number Viscous Flow Prediction for Appended Submarine Geometries". PW Bull SJ Watson. 22nd Symposium on Naval Hydrodynamics. Washington DC USA. August 1998.
DISCUSSION Ki-Han Kim Office of Naval Research, USA Would the detailed geometry of the hulls and propellers used for this project be available to the community for computational codes validation? AUTHORS' REPLY We wish to thank Dr. Kim for his question. At this time, the data base created in this project is the shared property of the MODs of the four participating nation. The decision to render this data public has not been taken at this time and will require the consent of all parties involved. DISCUSSION Arthur M. Reed Naval Surface Warfare Center, Carderock, USA The authors are to be congratulated on an interesting presentation on an ambitious project. I listened with a sense of deja vu. I believe that it would be useful to remind the attendees that at the 14th ONR Symposium in 1982, I presented a paper on the same subject. In that paper, we provided a full-scale and model- scale boundary layers and propeller-plane flow on the R/V Athena. Our measurements were made using 5-hole Pitot tubes and Prandt tubes. In the case of the CFD, we were restricted to boundary-layer theory, which was then the state of the art in viscous flow prediction (RANS was an as-of-yet unheard of technical in naval hydrodynamics). For those researchers who are interested, the model/ship geometry is available and the reports containing the measured model- and full-scale data can also be made available. AUTHORS' REPLY We wish to thank Dr. Reed for his discussion on our paper and the reminder that previous efforts have been made on this subject. We are well aware of the Athena data; however, our motivation was to use experimental technology available today, in particular LDV, to obtain more extensive and reliable data, at full scale but also at high Reynolds number at model scale. In particular, the capability offered by the large tunnel to investigate in the same conditions a significant of Reynolds number provides a challenging data base for turbulence models. DISCUSSION Ismail B. Celik West Virginia University, USA Have the authors compared the CPU time and convergence problems of RSTM versus 1- and 2- equation models? It is shown that RSTM performs by far the best among all turbulence models. Why then at INSEAN do they still follow the practice of 1- and 2-equation models? AUTHORS' REPLY We wish to thank Prof. Celik for raising a point which is clarified in the final version of the paper. In this collaborative effort, numerical work has been conducted by QinetiQ and INSEAN, using different codes with different capabilities. INSEAN has not implemented the RSTM model in its code and hence only shows results for 1- and 2-equation models.
