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Twenty-Third Symposium on Naval Hydrodynamics (2001)

Chapter: Validation of Tab Assisted Control Surface Computation

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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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Suggested Citation:"Validation of Tab Assisted Control Surface Computation." National Research Council. 2001. Twenty-Third Symposium on Naval Hydrodynamics. Washington, DC: The National Academies Press. doi: 10.17226/10189.
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lengths, word breaks, heading styles, and other typesetting-specific formatting, however, cannot be retained, and some typographic errors may have been accidentally inserted. Please use the print version of this publication as About this PDF file: This new digital representation of the original work has been recomposed from XML files created from the original paper book, not from the original typesetting files. Page breaks are true to the original; line VALIDATION OF TAB ASSISTED CONTROL SURFACE COMPUTATION 479 becomes much worse with the additional complication from turbulence models. Efforts to quantify the errors in RANS computations for practical problem shave produced dubious results. Nevertheless, attempts must be made to establish confidence in the RANS solutions. The most important thing to do is to assure that the solution is monotonically convergent. But this is not sufficient. It is well known that a nonphysical converged solution can be obtained. For example, one can construct a coarse grid for a turbulent flow about a body with the nondimensional first grid normal distance to the wall y+ on the order of several hundred and obtain a very nicely converged solution. But the solution will not be close to the real physics because a turbulent boundary layer can not be developed for such a large y+ for a conventional two-equation turbulence model. To resolve this difficulty, a sequence of finer grids must be used until the change in the solution due to the grid refinement is small enough to be acceptable to engineering requirement. Thus a careful check of convergence history and mesh refinement to obtain a grid-independent solution are the most effective approach to establish confidence in the results. Other researchers as in e.g., [23] have adopted a similar view. A C-grid with four blocks was used in the computation. The first block wraps around the entire control surface, the second block is on top of the control surface, the third covers the gap between the stabilizer and the flap and the final block covers the gap between the flap and the tab. The water tunnel is not modeled in the computation. A total of three meshes were considered. The coarse mesh consists of 112x28x20, 44x8x8, 8x8x12, and 8x8x12 grid cells for the first, second, third, and fourth block, respectively. This mesh consists of a total of about 65K grid cells. The medium mesh doubles the number of grid cells in each curvilinear coordinate direction of each block and has a total number of grid cells of about half a million. The fine mesh will have the number of grid cells increased by 50 percent in each direction of each block of the medium mesh, giving a total number of grid cells of about 1.6 millions. It will be seen that the solutions obtained by the fine and the medium mesh are almost identical, indicating that a grid independent solution has been achieved. The boundary conditions imposed are the following. The farfield boundary conditions at both the upper and lower wakes are zero gradient for the three components of the Cartesian velocity, the turbulence quantities and ω, and the nonreflecting boundary condition for the pressure. The nonreflecting boundary condition is important for good convergence and accuracy as mentioned earlier. On the outflow boundary at the top of the computational domain are imposed fixed values for the three components of the Cartesian velocity, the two turbulence quantities and ω, and zero gradient of the pressure. Because of the presence of the pedestal, the symmetric boundary condition is applied at the bottom of the computational domain. Non-slip boundary condition for the velocity and zero gradient for the pressure are applied at the wall boundary. The turbulent kinetic energy vanishes at the wall and the dissipation rate at the wall has been described earlier. For other boundaries such as between grid blocks and the interface between the upper and the lower wake, exact boundary conditions are applied. Some typical convergence histories of the root-mean-square of pressure for the case without flap and tab deflections and the case with 20 degrees deflections for both flap and tab are shown in Figure 2, where residue is defined as the root- mean-square value of the difference between the current calculated pressure and the last calculated one. It can be seen that flap and tab deflections do not seem to affect convergence rate. The residues for both cases drop more than three orders of magnitude in 200 multigrid cycles. The forces and moments become steady at about 200 multigrid cycles. It should be noted that the drop in the residue due to the multigrid starting procedure has not been included in Figure 2. This explains why logic10 (residue) starts at somewhere between −1 and −2, instead of 0. Fig 2. Root-mean square residue of pressure vs. multigrid cycles the authoritative version for attribution.

lengths, word breaks, heading styles, and other typesetting-specific formatting, however, cannot be retained, and some typographic errors may have been accidentally inserted. Please use the print version of this publication as About this PDF file: This new digital representation of the original work has been recomposed from XML files created from the original paper book, not from the original typesetting files. Page breaks are true to the original; line VALIDATION OF TAB ASSISTED CONTROL SURFACE COMPUTATION 480 Forces and Moments In the following discussion, the medium grid with a total of about half a million grid cells was used for the computed results. The grid consists of four blocks with grid size of 224x56x40, 88x16x16, 16x16x24, and 16x16x24. The Reynolds number based on the mean chordlength is 9.7x105. The coefficients of the forces and moments are also defined using the mean chordlength as the characteristic length. The lift and drag coefficients will be defined as the total forces applied to the entire control surface. The flap torque coefficient will include both the torque applied to the flap and the tab while the tab torque coefficient will include the torque applied to the tab alone. Figure 3 shows the comparison between measurement and computation of the lift coefficient as the angle of attack of the stabilizer varies from −6 to +15 degrees with no deflections for both flap and tab. The error bars on the data show 10% discrepancy in measurements. The lift coefficient is almost linear indicating insignificant viscous effect in this range of angles of attack. Both the predictions by the fine and the medium grids agree well with the measurement. However, the coarse grid prediction starts to deviate form the measurement by more than 10% after an angle of attack of 9 degrees, indicating insufficient grid resolution. Figure 3. Comparison of calculated and measured lift coefficients with flap and tab at zero deflection Figure 4 shows the comparison of the drag coefficient under the conditions as similar to those in Figure 3. Although a grid independent solution has been achieved between the fine and the medium grids, the drag coefficient is overpredicted by more than 20% in the neighborhood of the zero angle of attack and within 10% for greater than 10 degrees. The coarse grid prediction is even worse, again due to insufficient grid resolution. The effect on lift coefficient of varying the flap deflection from −15 degrees to +15 degrees is shown in Figure 5, where angle of attack for the stern stabilizer remains zero. Figure 4. Comparison of calculated and measured Fig 5. Comparison of calculated and measured lift drag coefficients with flap and tab at zero deflection coefficients with stern stabilizer and tab at zero deflection A grid independent solution has been achieved between the fine and the medium grid up to 10 degrees of flap deflection. At 15 degrees of flap deflection, the fine grid prediction is still within 10% of the measured values but the the authoritative version for attribution. medium grid prediction

lengths, word breaks, heading styles, and other typesetting-specific formatting, however, cannot be retained, and some typographic errors may have been accidentally inserted. Please use the print version of this publication as About this PDF file: This new digital representation of the original work has been recomposed from XML files created from the original paper book, not from the original typesetting files. Page breaks are true to the original; line VALIDATION OF TAB ASSISTED CONTROL SURFACE COMPUTATION 481 degenerates rapidly. The coarse grid prediction is inadequate beyond 5 degrees of flap deflection. One physical feature is worthy of mentioning. Consider a lift coefficient of 0.2 [Fig 3]. This lift can be achieved by an angle of attack of slightly less than 6 degrees of the entire control surface. It can also be achieved by a flap deflection of about 10 degrees but at a much smaller torque requirement. This is the essence of using a flap and also a tab which would be discussed later. Finally, the effect on lift, flap torque, and tab torque coefficients of varying the tab deflection from—60 degrees to +60 degrees are presented in Figures 6 through 8, respectively. Here, the stabilizer is at zero angle of attack and the flap has no deflection. Figure 6. Comparison of calculated and measured lift Figure 7. Comparison of calculated and measured coefficients with stern stabilizer and flap at zero flap torque coefficients with stern stabilizer and flap deflection at zero deflection A grid independent solution has not been obtained in the calculation of the lift coefficient as shown in Figure 6. However, the prediction of the lift from the fine grid is within 10% of measurement even at high tab deflection of 60 degrees. There is one discrepancy when tab deflection is less than 10 degrees. The slope of the measured lift is linear near zero tab deflection but is not zero. The predicted slope is almost zero when tab deflection is less than 10 degrees. If the measurement were correct, the discrepancy could be explained as insufficient grid resolution around the tab. The small increase in the lift due to a small tab deflection has not been picked up even by a grid as large as 1.6 million grid cells. It was mentioned earlier than a lift coefficient of 0.2 can be achieved by either an angle of attack of about 6 degrees of the entire control surface or by a flap deflection of about 10 degrees. This lift can also be obtained by a tab deflection of 40 degrees with even smaller tab torque requirement. The comparison of the flap torque coefficient is shown in Figure 7. It has a similar characteristic as the lift coefficient shown in Figure 6. A grid independent solution has not been achieved at high tab deflection, and the slope near zero tab deflection is much flatter than the measurement. The predicted slope of the tab torque coefficient near zero tab deflection seems to agree better with the measurement but the predicted torque coefficient at high tab deflection deviates from the measurement by more than 10%. It should be noted that the tab torque coefficient is smaller than the flap torque coefficient by approximately one order of magnitude. This is the main reason that the tab assisted control surface is of great practical interest. Figure 8. Comparison of calculated and measured tab torque coefficients with stern stabilizer and flap at zero deflection the authoritative version for attribution.

lengths, word breaks, heading styles, and other typesetting-specific formatting, however, cannot be retained, and some typographic errors may have been accidentally inserted. Please use the print version of this publication as About this PDF file: This new digital representation of the original work has been recomposed from XML files created from the original paper book, not from the original typesetting files. Page breaks are true to the original; line VALIDATION OF TAB ASSISTED CONTROL SURFACE COMPUTATION 482 CONCLUSIONS A numerical procedure for the prediction of the forces and moments of a tab assisted control surface has been developed. The procedure is based on solving the incompressible Reynolds-averaged Navier-Stokes equations coupled with a -ω, turbulence model. Computed results of lift, flap, and tab torque coefficients were compared with the measured data at a Reynolds number of 9.7x105 based on the mean chordlength. Three meshes with grid size of 65K, one half million, and 1.6 millions were used to investigate the grid independent solution. A grid independent solution was achieved in most of the cases except for some cases with high flap and tab deflections. The trend of the changes in the forces and moments due to the variations in the angle of attack of the stabilizer and the deflection of the flap and tab has been completely captured. In most cases investigated, the predictions are within 10% of the measurements. Some exceptions are the tab torque coefficients at high tab deflections and the slopes of the lift and flap torque coefficients near zero tab deflection. It is suggested that both the turbulence model and the grid resolution need to be improved. The fact that even with a grid as large as 1.6 million cells, a grid independent solution can only be achieved in most, but not all cases, indicates that more efficient numerical schemes and turbulence models are urgently needed. Despite all these limitations, the predictive procedure presented here is already a useful tool for the design of efficient control surfaces. ACKNOWLEDGMENTS This work is funded by the Office of Naval Research, Code 333, under the Mechanics and Energy Conversion Science and Technology Division (PE0602121). Dr Patrick Purtell is the technical monitor of this program. Dr. Nguyen Thang is the monitor at David Taylor Model Basin. Helpful discussions of experiment and measured data with Mr. David Bochinski at David Taylor Model Basin are gratefully acknowledged. Computer resources provided by the Department of Defense High Performance Computing Modernization Office (DOD-HPCMC) at NAVO and the Arctic Region Supercompting Center in Fairbank, AK are also gratefully acknowledged. REFERENCES 1. AGARD Conference Proceedings 515 on “High-Lift System Aerodynamics”, September, 1993. 2. Richard I.Sears and Robert B.Liddel, “Wind-Tunnel investigation of Conrol-Surface Characteristics, VI—A 3 Percent-Chord Plain Flap On the NACA 0015 Airfoil”. NACA Wartime Report 454, June 1942. 3. Whicker, L.Folger and Leo F.Fehlner, “Free-Stream Characteristics of a Family of Low-Aspect Ratio, All-Movable Control Surfaces For Application to Ship Design”, David Taylor Model Basin Report 933, December 1958. 4. Bowers, Allen, “Wind Tunnel Investigation of the Characteristics of a Flapped Control Surface Mounted on a Simulated Submarine Hull”, University of Maryland Wind Tunnel Report No. 259, June, 1959. 5. Kerwin, Justine E., Philip Mandel and S.Dean Lewis, “An Experimental Study of a Series of Flapped Rudders”, Journal of Ship Research, December, 1972 6. Goodrich, G.J. and A.F.Molland, “Wind Tunnel Investigation of Semi-Balanced Ship Skeg-Rudders”, The Royal Institute of Naval Architects, pp. 285–307, 1979. 7. Soding, H., “Limits of Potential Theory in Rudder Flow Predictions”, Twenty-Second Symposium on Naval Hydrodynamics, Washington, D.C., pp. 264–276, August 9–14, 1998. 8. Chau, Shiu-Wu, “Computation of Rudder Force and Moments in Uniform Flow”, Ship Technology Research Vol. 45, pp. 3–13, 1998. 9. Gowing, Scott, Thang Nguyen and David Bochinski, “T.A.C. Test Static Results in the 24” Water Tunnel”, NSWC, CD, not yet published, 1999. 10. Wilcox, D.C., Turbulence Modeling for CFD, DCW Industries, Inc. CA, 1993 11. Speziale, Charles G., “Comparison of Explicit and Traditional Algebraic Stress Models of Turbulence”, AIAA Journal vol. 35, No. 9, September 1997. 12. Chorin, A.J., “A Numerical Method for Solving Incompressible Viscous Flow Problem”, Journal of Computational Physics, vol. 2, 275, 1967. 13. Turkel, E., “Preconditioned Methods for Solving the Incompressible and Low Speed Compressible Equations”, Journal of Computational Physics, vol. 72, 277, 1987. the authoritative version for attribution.

lengths, word breaks, heading styles, and other typesetting-specific formatting, however, cannot be retained, and some typographic errors may have been accidentally inserted. Please use the print version of this publication as About this PDF file: This new digital representation of the original work has been recomposed from XML files created from the original paper book, not from the original typesetting files. Page breaks are true to the original; line VALIDATION OF TAB ASSISTED CONTROL SURFACE COMPUTATION 483 14. Yee, H.C., “A Class of High-Resolution Explicit and Implicit Shock-Capturing Methods”, NASA Technical Memorandum 101088, February, 1989. 15. James on, A., “Time Dependent Calculations Using Multigrid with Applications to Unsteady Flows Past Airfoils and Wings”, AIAA 91–1596, June 1991. 16. Liu, C., X.Zheng and C.H.Sung, “Preconditioned Multigrid Methods for Unsteady Incompressible Flows”, Journal of Computational Physics, vol. 139, 35–57, 1998. 17. Brandt, A., “Multigrid Techniques: 1984 Guide, with Applications to Fluid Dynamics”, 1984, 191 pages, ISBN-3–88457–081–1; GMD-Studien Nr 85; Available from GMD-AIW, Postfach 1316, D-53731, St. Augustin 1, Germany, 1984. 18. Jameson, A., “Multigrid Algorithms for Compressible Flow Calculations”, Lecture Notes in Mathematics, No. 1228, Proceedings of the Second European Conference on Multigrid Methods, Cologne, pp. 166–201, October 1–4, 1985. 19. Brandt, A., “Barriers to Achieving Textbook Multigrid Efficiency (TME) in CFD”, NASA/CR-1998–207647, ICASE Interim Report No. 32, April 1998. 20. Hedstrom, G.W., “Nonreflecting Boundary Conditions for Nonlinear Hyperbolic System”, Journal of Computational Physics, vol. 30, pp. 222–237, 1979. 21. Rudy, D.H., and J.C.Strikwerda, “Boundary Conditions for Subsonic Compressible Navier-Stokes Equations”, Computers and Fluids, vol. 9, pp. 327–338, 1981. 22. Sung, C.H., “An Explicit Runge-Kutta Method for 3D Incompressible Turbulent Flows”, DTNSRDC/SH—1244–01, July 1987. 23. Jameson, A. and L.Martinelli, “Mesh Refinement and Modeling Errors in Fluid Simulation”, AIAA Journal vol. 36, No. 5, May 1998. the authoritative version for attribution.

lengths, word breaks, heading styles, and other typesetting-specific formatting, however, cannot be retained, and some typographic errors may have been accidentally inserted. Please use the print version of this publication as About this PDF file: This new digital representation of the original work has been recomposed from XML files created from the original paper book, not from the original typesetting files. Page breaks are true to the original; line VALIDATION OF TAB ASSISTED CONTROL SURFACE COMPUTATION 484 DISCUSSION Y.Tahara Oskaka Prefecture University Japan In your computation, laminar-to-turbulent flow transition was not considered, although your experimental condition (i.e., Re≈106) apparently implies that there exists laminar-flow region on the wing (stabilizer in your definition) surface. Inclusion of the effects is generally essential for accurate prediction of hydrodynamic forces especially for drag component [Tahara et al., 1998, 2000]. In addition, the conventional two equation model used in your work may not be suitable for the purpose. REFERENCES: Tahara, Y., et al., “An Application of RaNS Equation Method to Strut/Bulb Configuration of America's Cup Sailing Yacht and Comparison with Experiments,” J. Kansai Society of Naval Architects, No. 230, 1998, pp. 163–171. Tahara, Y., et al., “Development of Ballast Bulb for IACC Sailing Yacht—Especially for Investigation on Basic Low Drag Form,” J. Kansai Society of Naval Architects, No. 234, 2000, pp. 51–59 AUTHOR'S REPLY Due to a relatively high turbulence level in a water tunnel, early experimental tests indicated that the flow was turbulent at a Reynolds number of about one million based on a mean chordlength of 9.53 inches. For this reason, computations were made assuming the flow was completely turbulent. Transition from laminar to turbulence was not considered. Admittedly, turbulence models are not perfect for a complex flow such as the one investigated here. However, the −ω turbulence model used here worked quite satisfactorily in our opinion. It is believed that further improvement of accuracy can be made by increasing the grid size, particularly in the leeward side of the flow region. This will be investigated in the future. the authoritative version for attribution.

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"Vive la Revolution!" was the theme of the Twenty-Third Symposium on Naval Hydrodynamics held in Val de Reuil, France, from September 17-22, 2000 as more than 140 experts in ship design, construction, and operation came together to exchange naval research developments. The forum encouraged both formal and informal discussion of presented papers, and the occasion provides an opportunity for direct communication between international peers.

This book includes sixty-three papers presented at the symposium which was organized jointly by the Office of Naval Research, the National Research Council (Naval Studies Board), and the Bassin d'Essais des Carènes. This book includes the ten topical areas discussed at the symposium: wave-induced motions and loads, hydrodynamics in ship design, propulsor hydrodynamics and hydroacoustics, CFD validation, viscous ship hydrodynamics, cavitation and bubbly flow, wave hydrodynamics, wake dynamics, shallow water hydrodynamics, and fluid dynamics in the naval context.

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