Numerical Computations of Supersonic Flow Over a Square 

Cross-Section Missile 

by Sidra I. Silton and Jubaraj Sahu 

ARL-TR-3562 August 2005 

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Army Research Laboratory 

Aberdeen Proving Ground, MD 21005-5066 

ARL-TR-3562 August 2005 

Numerical Computations of Supersonic Flow Over a Square 

Cross-Section Missile 

Sidra I. Silton and Jubaraj Sahu 

Weapons and Materials Research Directorate, ARL 

Approved for public release; distribution is unlimited.
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Numerical Computations of Supersonic Flow Over a Square Cross-Section Missile 

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Sidra I. Silton and Jubaraj Sahu 

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14. ABSTRACT 

This report describes a computational study undertaken to determine the aerodynamics of a nonaxisymmetric missile with a 

square cross section. Numerical solutions have been obtained at supersonic speeds for various roll orientations and angles of 

attack using a two-equation Reynolds-averaged Navier-Stokes turbulence model. Numerical results show the qualitative 

features (vortices and cross-flow separation regions) of the flow field at various stream wise positions along the missile 

configurations. Aerodynamic coefficients have been obtained from the computed solutions and found to match well with the 

available experimental data for these configurations. These numerical results show the ability of computational fluid dynamics

techniques to accurately predict the aerodynamics of nonaxisymmetric missiles with a square cross section. 

15. SUBJECT TERMS

nonaxisymmetric missile, complex flows, computational fluid dynamics 

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Sidra I. Silton 

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iii

Contents 

List of Figures iv

List of Tables vi

Acknowledgments vii

1. Introduction 1

2. Solution Technique 3

3. Model Geometry and Flowfield Conditions 4

3.1 Model Geometry..............................................................................................................4 

3.2 Numerical Grid................................................................................................................5 

3.3 Flowfield Conditions.......................................................................................................5 

4. Results and Discussion 7

4.1 Grid Resolution Study.....................................................................................................7

4.2 Aerodynamic Coefficients.............................................................................................12 

4.3 Pressure Coefficient Distributions.................................................................................18 

4.4 Flowfield Visualizations................................................................................................23 

5. Summary and Conclusions 35

6. References 37

Distribution List 39 
iv

List of Figures 

Figure 1. Reference dimensions and orientation for the square section missile, no fins................5 

Figure 2. Sketch (inches) of side and corner mounted tail fins (17)...............................................6 

Figure 3. Corner fin configuration, side and base views. ...............................................................7 

Figure 4. Side fin configuration, side and base views. ...................................................................7 

Figure 5. (a) Computational grid for corner fin geometry with y+

 = 30 spacing. 

Close-up of (b) nose tip and (c) fins. .........................................................................................8 

Figure 6. (a) Computational grid for corner fin geometry with y+ = 1 spacing. 

Close-up of (b) nose tip and (c) fins. .........................................................................................9 

Figure 7. (a) Nine million hexahedral computational grid for side fin geometry 

with y+ = 1 spacing. Close-up of (b) nose tip and (c) fins. .....................................................10 

Figure 8. Roll orientation of 0°, 22.5°, and 45° for (a)–(c) corner fin projectile and 

(d)–(f) side fin projectile looking from projectile base............................................................11 

Figure 9. Axial force coefficient as a function of roll orientation. ...............................................14 

Figure 10. Normal force coefficient as a function of roll orientation...........................................15 

Figure 11. Side force coefficient as a function of roll orientation................................................16 

Figure 12. Rolling moment coefficient as a function of roll orientation. .....................................16 

Figure 13. Pitching moment coefficient as a function of roll orientation.....................................17 

Figure 14. Yawing moment coefficient as a function of roll orientation......................................18 

Figure 15. Reference axial body position locations on corner fin configuration..........................18 

Figure 16. Pressure coefficient distribution at x/D = 5.5, φ = 0°, α = 14°....................................19 

Figure 17. Pressure coefficient distribution at x/D = 5.5, φ = 22.5°, α = 14°...............................20 

Figure 18. Pressure coefficient distribution at x/D = 5.5, φ = 45°, α = 14°..................................20 

Figure 19. Pressure coefficient distribution at x/D = 8.5, φ = 0°, α = 14°....................................21 

Figure 20. Pressure coefficient distribution at x/D = 8.5, φ = 22.5°, α = 14°...............................21 

Figure 21. Pressure coefficient distribution at x/D = 8.5, φ = 45°, α = 14°..................................22 

Figure 22. Pressure coefficient distribution at x/D = 11.5. (a) φ = 0°, α = 14°, 

(b) φ = 22.5°, α = 4°, (c) φ = 45°, α = 14°...............................................................................23 

Figure 23. Pressure coefficient distribution at x/D = 12.5. (a) φ = 0°, α = 14°, 

(b) φ = 22.5°, α = 14°, (c) φ = 45°, α = 14°.............................................................................24 

Figure 24. Pressure coefficient distribution for Mach 2.5 and φ = 0° at (a) x/D = 5.5, 

(b) x/D = 8.5, (c) x/D = 11.5, and (d) x/D = 12.5. ..................................................................25 

Figure 25. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for corner fin configuration, Mach 2.5, φ = 0°, α = 14°. .........................................................25 
v

Figure 26. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for side fin configuration, Mach 2.5, φ = 0°, α = 14°..............................................................26 

Figure 27. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for corner fin configuration, Mach 2.5, φ = 22.5°, α = 14°. ....................................................26 

Figure 28. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for side fin configuration, Mach 2.5, φ = 22.5°, α = 14°.........................................................26 

Figure 29. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for corner fin configuration, Mach 2.5, φ = 45°, α = 14°. .......................................................27 

Figure 30. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for side fin configuration, Mach 2.5, φ = 45°, α = 14°............................................................27 

Figure 31. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for side fin configuration, Mach 2.5, φ = 0°, α = 20°..............................................................27 

Figure 32. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for corner fin configuration, Mach 4.5, φ = 0°, α = 14°. .........................................................28 

Figure 33. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for side fin configuration, Mach 4.5, φ = 0°, α = 14°..............................................................28 

Figure 34. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for corner fin configuration, Mach 4.5, φ = 22.5°, α = 14°. ....................................................28 

Figure 35. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for side fin configuration, Mach 4.5, φ = 22.5°, α = 14°.........................................................29 

Figure 36. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for corner fin configuration, Mach 4.5, φ = 45°, α = 14°. .......................................................29 

Figure 37. Three-dimensional (a) pitot pressure contours and (b) vorticity contours 

for side fin configuration, Mach 4.5, φ = 45°, α = 14°............................................................29 

Figure 38. Vorticity contour plots at x/D = 5.5, Mach 2.5, α = 14° for 

([a]–[c]) corner fin configuration φ = 0°, 22.5°, 45°, respectively, and 

([d]–[f]) side fin configuration, φ = 0°, 22.5°, 45°, respectively. ............................................30 

Figure 39. Vorticity contour plots at x/D = 8.5, Mach 2.5, α = 14° for 

([a]–[c]) corner fin configuration φ = 0°, 22.5°, 45°, respectively, and 

([d]–[f]) side fin configuration, φ = 0°, 22.5°, 45°, respectively. ............................................30 

Figure 40. Vorticity contour plots for side fin configuration at Mach 2.5, φ = 0°, 

([a]–[b]) α = 14°, x/D = 5.5 and 8.5, respectively, and ([c]–[d]) α = 20°, 

 x/D = 5.5 and 8.5, respectively. ..............................................................................................31 

Figure 41. Vorticity contour plots at x/D = 8.5, α = 14° for corner fin configuration 

([a]–[c]) Mach 2.5, φ = 0°, 22.5°, 45°, respectively, and ([d]–[f]) Mach 4.5, 

φ = 0°, 22.5°, 45°, respectively. ...............................................................................................31 

Figure 42. Pressure contours for ([a]–[c]) corner fin configuration and 

([d]–[f]) side fin configuration at roll orientations indicated for Mach 2.5, 

α = 14°, x/D = 11.5..................................................................................................................32 

Figure 43. Vorticity contours for ([a]–[c]) corner fin configuration and 

([d]–[f]) side fin configuration at roll orientations indicated for Mach 2.5, 

α = 14°, x/D = 11.5..................................................................................................................32 
vi

Figure 44. Vorticity contours for corner fin configuration at roll orientations 

indicated, α = 14°, x/D = 11.5 for ([a]–[c]) Mach 2.5 and ([d]–[f]) Mach 4.5........................33 

Figure 45. Vorticity contours for corner fin configuration at roll orientations 

indicated, α = 14°, x/D = 12.5 for ([a]–[c]) Mach 2.5 and ([d]–[f]) Mach 4.5........................34 

Figure 46. Experimental vapor screen images for corner fin configuration at roll 

orientations indicated, α = 14°, x/D = 12.5 for ([a]–[c]) Mach 2.5 and 

([d]–[f]) Mach 4.5. ...................................................................................................................34 

Figure 47. Vorticity contours for side fin configuration at Mach 2.5, φ = 0°, 

x/D = 11.5, (a) α = 14° and (b) α = 20°...................................................................................35 

Figure 48. Vorticity contours for side fin configuration at Mach 2.5, α = 0°, 

x/D = 12.5, (a) α = 14° and (b) α = 20°...................................................................................35 

List of Tables 

Table 1. Grid resolution study for corner fin projectile at Mach 2.5, φ = 22.5°, 

α = 14°. ....................................................................................................................................11 

Table 2. Grid resolution study for side fin projectile at Mach 2.5, φ = 0°, α = 14°. ....................11 

Table 3. Aerodynamic coefficients for corner fin configuration at Mach 2.5, α = 14°................12 

Table 4. Aerodynamic coefficients for corner fin configuration at Mach 4.5, α = 14°................13 

Table 5. Aerodynamic coefficients for side fin configuration at Mach 2.5, α = 14°. ..................13 

Table 6. Aerodynamic coefficients for side fin configuration at Mach 4.5, α = 14°. ..................13 

Table 7. Aerodynamic coefficients for side fin configuration at Mach 2.5, φ = 0°, 

α = 20°. ....................................................................................................................................14 
vii

Acknowledgments 

This work was accomplished as part of The Technical Cooperation Program between the United 

States, Canada, United Kingdom, and Australia on Weapons Technical Panel TP-2 Key 

Technology Area 2-19. The authors wish to thank Sukumar Chakravarthy from Metacomp 

Technologies for his expert technical advice. This work was supported by a grant of computer 

time from the Department of Defense High Performance Computing Major Shared Resource 

Center at the U.S. Army Research Laboratory. 
viii

INTENTIONALLY LEFT BLANK. 
1

1. Introduction 

The prediction of aerodynamic coefficients for projectile configurations is essential in assessing 

the performance of new designs. Accurate determination of aerodynamics is critical to the lowcost development of new advanced guided projectiles, rockets, missiles, and smart munitions. 

Fins, canards, and jets can be used to provide control for maneuvering projectiles and missiles. 

The flow fields associated with these control mechanisms for the modern weapons are complex, 

involving three-dimensional (3-D) shock-boundary layer interactions and highly viscous 

dominated separated flow regions (1–3). Computational fluid dynamics (CFD) has emerged as a 

critical technology for the aerodynamic design and assessment of weapons. Improved computer 

technology and state-of-the-art numerical procedures enable solutions to complex, 3-D problems 

associated with projectile and missile aerodynamics. In general, these techniques produce 

accurate and reliable numerical results for simple projectile and missile configurations at small 

angles of attack. 

Modern weapons and the associated flow fields have increased in complexity as a result of 

changing military requirements. An increasing number of new weapon configurations are 

beyond the scope of traditional engineering prediction methods. Modern projectiles and missiles 

are expected to experience moderate to large angles of attack during flight. The flow field 

surrounding a slender missile at incidence features flow separations both from smooth bodies 

and sharp-edged lifting and control surfaces. Body and wing vortices (4, 5) can dominate the 

flow field and interact with one another resulting in a very complex flow field even for simple 

geometries. Earlier work (6, 7) on modeling of high angle of attack flow fields for finned missile 

configurations was initiated as part of The Technical Cooperation Program (TTCP) effort with 

participants from Canada, the United Kingdom, Australia, and the United States. This study 

addressed the application of Navier-Stokes methods to the prediction of flows around finned 

missile configurations and assessed the capabilities of Navier-Stokes solvers for the prediction of 

missile flow field at moderate to high angles of attack. It promoted technical interchange 

between aerodynamics specialists in the areas of grid generation, algorithms, turbulence 

modeling, and flow field visualization. The results of this study were reported in a special 

session of the AIAA Applied Aerodynamics Conference held in Denver, CO in August 2000 

(6–8). The study has shown that all Navier-Stokes methods employed were capable of providing 

accurate predictions of the overall normal force and pitching moment characteristics, even at 

very high angles of attack. However, some significant discrepancies in the prediction of the 

axial force were evident. In general, the agreement between computed and measured flow fields 

was good. Higher order turbulence models such as the k-ε models did not provide better 

accuracy compared to simpler algebraic and one-equation models. The next phase of the 

research was to extend the scope of that study and to undertake a follow-on study to apply CFD 
2

technology to the prediction of aerodynamic flow fields for nonaxisymmetric missile 

configurations. 

Maneuvering nonaxisymmetric missile configurations can be used for increased lethality and are 

being explored for future weapon systems in various scenarios. Additionally, nonaxisymmetric 

projectiles have been shown to generate large amounts of lift compared to conventional circular 

projectiles (9). The aerodynamics of such configurations at high speeds are quite complex due to 

strong cross-flow influences and flow separation. Of particular interest was the accurate 

determination of supersonic and hypersonic flow over elliptic projectiles at moderate angles of 

attack. The flow field for such projectiles with nonaxisymmetric cross sections is complex, 

especially in the presence of jets used to maneuver these projectiles. This work, initiated as part 

of another TTCP effort with participants from Canada, the United Kingdom, and the United 

States, was aimed at assessing the capabilities of the both Euler and Navier-Stokes solvers 

currently available to research scientists for supersonic flow over elliptic projectiles for both jetoff and jet-on conditions (10, 11). The TTCP research effort also included wind-tunnel testing 

and free-flight testing of these projectiles. CFD approaches that included both Euler and NavierStokes schemes, structured and unstructured techniques, and a variety of turbulence models 

were used in the numerical simulations (10, 11–15). In general, very good agreement of the 

computed aerodynamic coefficients with the experimental data was achieved at all angles of 

attack investigated for the jet-off uncontrolled conditions. Accurate prediction of the jet-on cases 

was found to be a little more challenging than the jet-off cases (16). The results of this study 

were reported in a special session of the AIAA Atmospheric Flight Mechanics Conference held 

in Montreal, Canada in August 2001. 

The TTCP efforts have continued recently with a focus on the application of CFD to missile 

aerodynamics which takes it to the next level of difficulty, that of multiple sets of fins on a 

elliptic cross-section missile. An additional nonaxisymmetric configuration was also 

investigated. Again, the objective here is to compare the results obtained from Navier-Stokes 

computations against experimental data, evaluate the accuracy of the predictive technologies, 

and identify priorities for future development. The present research focuses on the application of 

advanced CFD techniques for accurate numerical prediction of supersonic flow over two 

nonaxisymmetric missile configurations, one with a square cross section and the other with an 

elliptic cross section; both have multiple sets of complex fins. Numerical computations for these 

complex configurations have been performed at supersonic speeds and for several angles of 

attack. Only the results of the square cross-section axisymmetric missile configuration are 

presented in this report. 

A description of the computational techniques is presented, followed by a description of the 

applications of these techniques to the square section configurations. Computed results for these 

configurations are presented at Mach 2.5 and 4.5 and several angles of attack from 0° to 20°. 

Computed data have been compared with experimental data provided by the Defence Science 
3

and Technology Laboratory (DSTL, formerly the Defence Evaluation and Research Agency of 

the United Kingdom) (17). 

2. Solution Technique 

A commercially available code, CFD++ (18, 19), is used for the CFD simulations. The basic 

numerical framework in the code contains unified-grid, unified-physics, and unified-computing 

features. A brief synopsis of this framework and methodology is given in the following 

paragraphs. The user is referred to the previous references for details. 

The 3-D, Reynolds-averaged Navier-Stokes (20) equations are solved using the following finite 

volume method: 

[ ] –

v

dV dA dV

t ν

∂

+ ⋅=

∂ ∫∫ ∫ W FG H v , (1) 

where W is the vector of conservative variables, F and G are the inviscid and viscous flux 

vectors, respectively, H is the vector of source terms, V is the cell volume, and A is the surface 

area of the cell face. 

The numerical framework of CFD++ is based on the following general elements: 

1. unsteady compressible and incompressible Navier-Stokes equations with turbulence 

modeling (unified physics), 

2. unification of Cartesian, structured curvilinear, and unstructured grids, including hybrids 

(unified grid), 

3. unification of treatment of various cell shapes including hexahedral, tetrahedral, and 

triangular prism cells 3-D, quadrilateral and triangular cells two dimensional, and linear 

elements one-dimensional (unified grid), 

4. treatment of multiblock patched aligned (nodally connected), patched-nonaligned, and 

overset grids (unified grid), 

5. total Variation Diminishing discretization based on a new multidimensional interpolation 

framework, 

6. Riemann solvers to provide proper signal propagation physics, including versions for 

preconditioned forms of the governing equations, 

7. consistent and accurate discretization of viscous terms using the same multidimensional 

polynomial framework, 
4

8. pointwise turbulence models that do not require knowledge of distance to walls, 

9. versatile boundary condition implementation including a rich variety of integrated 

boundary condition types for the various sets of equations, and 

10. implementation on massively parallel computers based on the distributed-memory 

message-passing model using native message-passing libraries or MPI, PVM, etc. (unified 

computing). 

The code has brought together several ideas on convergence acceleration to yield a fast 

methodology for all flow regimes. The approach can be labeled as a “preconditioned-implicitrelaxation” scheme. It combines three basic ideas: implicit local time-stepping, relaxation, and 

preconditioning. Preconditioning the equations ideally equalizes the eigenvalues of the inviscid 

flux Jacobians and removes the stiffness arising from large discrepancies between the flow and 

sound velocities at low speeds. Use of an implicit scheme circumvents the stringent stability 

limits suffered by their explicit counterparts, and successive relaxation allows update of cells as 

information becomes available and thus aids convergence. 

Second-order discretization was used for the flow variables and the turbulent viscosity equation. 

Turbulence closure is based on topology-parameter-free formulations. The realizable k-ε

turbulence model (21) was used for the computation of turbulent flows for the missile 

configurations considered here. The next section describes the model geometries and the flow 

conditions for the square-section missile (17). 

3. Model Geometry and Flowfield Conditions 

3.1 Model Geometry 

Two configurations were utilized for the square-section missile (17)—corner fins and side fins. 

The nose and body of the configurations were the same for each configuration (figure 1). The 

nose is 340.36 mm long and transitions from a pointed ogive to a square cross section. The 

remainder of the projectile has a square cross section (93.98 mm on a side) for a total projectile 

length of 1221.74 mm. All moments are referenced from the nose and are positive in the 

direction indicated in figure 1. 

The fin geometry, an elongated pyramid, was the same for each configuration (figure 2). The 

root chord was 74.42 mm, with the leading edge of the fin placed 1040.13 mm from the nose tip. 

Each fin was 74.42 mm high and had a sweep angle of 45°. The fin had a taper angle of 6.83° 

with a maximum thickness of 7.87 mm. The only difference between the two configurations was 

the circumferential placement of the fins on the body. For the corner fin configuration, the fin 

tip of each fin was placed on the edge of the square (figure 3). For the side fin configuration, the 

fin tip of each fin was placed on the centerline of one side of the square (figure 4). 
5

Figure 1. Reference dimensions and orientation for the square section missile, no fins. 

3.2 Numerical Grid 

Four unstructured grids were created, two for each geometry, using ICEM CFD (22). For the 

corner fin configuration, a grid containing 9.9 million hexahedral cells and 30,000 triangular 

prisms was constructed to have a y+ of ~30 (figure 5). A second, more efficiently constructed 

grid containing 6.2 million hexahedral cells was constructed to have a y+ of ~1.0 (figure 6). For 

the side fin configuration, both grids were constructed with a y+ of ~1.0, but had different cell 

distributions. The resulting grids had 6.7 million and 9.0 million hexahedral cells (figure 7), 

respectively. A grid resolution study was completed between these grids and is presented in 

section 4.1. 

3.3 Flowfield Conditions 

A series of computations at both Mach 2.5 and Mach 4.5 was performed for each geometry at 

14° angle of attack for roll orientations of 0°, 22.5°, and 45° (figure 8). An additional 

computational series was performed at Mach 2.5 for the side fin configuration at 20° angle of 

attack for the 0° roll orientation. These computations were completed in the blind as the 

experimental test was conducted concurrently. In order to accurately compare the CFD results to 

the experimental results after they became available, wind tunnel Reynolds number, Re, 

stagnation temperature, T0, and stagnation pressure, P0, were matched; Re = 1.31 × 107

 per meter 

for both Mach 2.5 and 4.5, T0 = 324.82 K and P0 = 153.22 kPa for Mach 2.5, and T0 = 338.71 K 

and P0 = 446.83 kPa for Mach 4.5. These conditions correspond to static temperatures and 

pressures of 144.36 K and 8.92 kPa, respectively, at Mach 2.5 and 67.07 K and 1.54 kPa, 

respectively, at Mach 4.5.
6

Figure 2. Sketch (inches) of side and corner mounted tail fins (17). 
7

Figure 3. Corner fin configuration, side and base views. 

Figure 4. Side fin configuration, side and base views. 

4. Results and Discussion 

Numerical computations using viscous Navier-Stokes methods were performed to predict the 

flow field and aerodynamic coefficients for wind tunnel test conditions. Thus, computations 

were conducted at Mach 2.5 and 4.5 for each configuration. Calculations were completed for 

each configuration at 0°, 22.5°, and 45° roll orientation, φ (figure 8). For each Mach number and 

roll orientation, calculations were completed at 14° angle of attack, α. For the side fin 

configuration at a roll orientation of 0° (figure 8d), a calculation was also performed for Mach 

2.5 at 20° angle of attack. Full 3-D computations were performed and no symmetry was used. 

4.1 Grid Resolution Study 

A grid resolution study was conducted for the corner fin configuration at Mach 2.5, φ = 22.5°, 

α = 14°. The results of the y+

 = 30 grid, with the wall function option employed, were compared 

to that of the y+

 = 1 grid (table 1). Almost all coefficients were found to be within 1% or less. 

Only the rolling moment coefficient, Cl, had a greater discrepancy at 7%. This shows that either 

grid would be adequate for determining the aerodynamic coefficients. The y+

 = 30 grid was used 

for the Mach 2.5, φ = 0°, α = 14° case. The y+

 = 1 grid was used for the remainder of the cases 

as the grid was smaller and required less computation time. 

A grid resolution study was also conducted for the side fin configuration for the Mach = 2.5, 

φ = 0°, α = 14° case (table 2). In this configuration, no side force, yawing moment, or rolling 

moment is expected. Again, almost all coefficients agreed to within 1%. For the coefficients 

expected to be near zero, small changes in value correspond to an unrealistically large percent 

error. Thus, percent difference is not calculated for these coefficients. The second grid did 
8

Figure 5. (a) Computational grid for corner fin geometry with y+ = 30 spacing. Close-up of (b) nose tip and 

(c) fins. 
9

Figure 6. (a) Computational grid for corner fin geometry with y+ = 1 spacing. Close-up of (b) nose tip and 

(c) fins. 
10

Figure 7. (a) Nine million hexahedral computational grid for side fin geometry with y+ = 1 spacing. Close-up 

of (b) nose tip and (c) fins. 

better in predicting the experimental normal force than the first grid. Also, CFD++ would not 

run for the first grid in some of the roll configurations. Therefore, the second grid (9.0 million 

hexahedral cells) was used for the remainder of the study.
11

Figure 8. Roll orientation of 0°, 22.5°, and 45° for (a)–(c) corner fin projectile and (d)–(f) side fin 

projectile looking from projectile base. 

Table 1. Grid resolution study for corner fin projectile at Mach 2.5, 

φ = 22.5°, α = 14°. 

Coefficient y+

 = 30 y+ = 1 % difference 

CN 3.745 3.757 0.307 

CS 0.864 0.873 1.075 

CA 0.466 0.463 –0.683 

CL 3.521 3.533 0.338 

CD 1.358 1.358 –0.020 

CL/CD 2.593 2.602 0.358 

Cm –24.685 –24.730 0.181 

Cn –7.598 –7.643 0.588 

Cl 0.059 0.055 –7.054 

Xcp 6.591 6.583 –0.127 

Table 2. Grid resolution study for side fin projectile at Mach 2.5, φ = 0°, 

α = 14°. 

Coefficient Grid 1 Grid 2 % difference 

CN 4.017 4.014 0.065 

CS 0.000 0.000 — 

CA 0.492 0.488 0.814 

CL 3.778 3.777 0.041 

CD 1.449 1.444 0.312 

CL/CD 2.608 2.615 –0.271 

Cm –26.014 –25.982 0.122 

Cn 0.003 0.000 — 

Cl 0.000 0.000 — 

Xcp 6.477 6.473 0.057 
12

4.2 Aerodynamic Coefficients 

Aerodynamic coefficients including forces, moments, and center of pressure locations were 

determined for all cases. The moment reference is the nose-tip of the projectile. Comparing the 

computational results to the experimental results shows remarkably good agreement (tables 3–7). 

The values of axial force coefficient, CA, normal force coefficient, CN, and side force coefficient, 

CS, are obtain directly from the experimental data and the numerical results. The values of lift 

coefficient, CL, drag coefficient, CD, and axial center of pressure, XCP, are obtained from the 

following equations: 

 cos sin CC C LN A = α − α , (2) 

 sin cos CC C DN A = α + α , (3) 

and 

m

CP

N

C X

C = − , (4) 

where α is the flow angle of attack and Cm is the pitching moment. 

As expected, CA does not change significantly with fin location or roll orientation for either the 

experimental or the computational results (figure 9). CA increases slightly for a given Mach 

number when comparing the corner fin configuration to side fin configuration, which is likely 

due to shock interaction effects due to fin location. For a given configuration, CA decreases with 

increasing Mach number. The increase in angle of attack for the side fin configuration with a 

Table 3. Aerodynamic coefficients for corner fin configuration at Mach 2.5, α = 14°. 

φ = 0° φ = 22.5° φ = 45° 

Coefficient Exp CFD Exp CFD Exp CFD 

CN 4.125 4.111 3.808 3.758 3.196 3.179 

CS –0.0224 0.000 0.972 0.873 –0.049 0.001 

CA 0.480 0.467 0.475 0.463 0.476 0.466 

CL 3.886 3.876 3.580 3.534 2.986 2.972 

CD 1.464 1.448 1.382 1.358 1.235 1.221 

CL/CD 2.654 2.677 2.590 2.602 2.418 2.434 

Cm –27.126 –27.289 –24.865 –24.738 –19.688 –19.809 

Cn 0.167 0.002 –8.271 –7.645 0.375 –0.006 

Cl 0.003 0.000 0.061 0.055 –0.004 0.000 

Xcp 6.576 6.638 6.530 6.583 6.160 6.231 
13

Table 4. Aerodynamic coefficients for corner fin configuration at Mach 4.5, α = 14°. 

φ = 0° φ = 22.5° φ = 45° 

Coefficient Exp CFD Exp CFD Exp CFD 

CN 3.108 3.082 2.949 2.904 2.778 2.743 

CS 0.011 0.000 0.260 0.219 –0.001 0.000 

CA 0.432 0.419 0.426 0.412 0.422 0.407 

CL 2.911 2.889 2.758 2.718 2.593 2.563 

CD 1.171 1.152 1.127 1.102 1.082 1.059 

CL/CD 2.486 2.507 2.447 2.467 2.396 2.421 

Cm –20.315 –20.588 –18.924 –18.973 –17.263 –17.411 

Cn –0.124 0.002 –2.387 –2.029 –0.034 –0.001 

Cl 0.000 0.000 0.062 0.066 0.001 0.000 

Xcp 6.536 6.680 6.417 6.534 6.214 6.348 

Table 5. Aerodynamic coefficients for side fin configuration at Mach 2.5, α = 14°. 

φ = 0° φ = 22.5° φ = 45° 

Coefficient Exp CFD Exp CFD Exp CFD 

CN 4.056 4.014 3.770 3.709 3.166 3.155 

CS –0.020 0.000 0.885 0.814 –0.068 0.000 

CA 0.501 0.488 0.500 0.486 0.499 0.488 

CL 3.814 3.777 3.537 3.481 2.951 2.944 

CD 1.467 1.444 1.397 1.369 1.250 1.237 

CL/CD 2.560 2.615 2.532 2.543 2.361 2.380 

Cm –26.134 –25.982 –24.286 –24.082 –19.286 –19.459 

Cn 0.141 0.000 –7.287 –6.978 0.527 –0.001 

Cl –0.002 0.000 –0.053 –0.047 0.001 0.000 

Xcp 6.443 6.473 6.442 6.494 6.092 6.167 

Table 6. Aerodynamic coefficients for side fin configuration at Mach 4.5, α = 14°. 

φ = 0° φ = 22.5° φ = 45° 

Coefficient Exp CFD Exp CFD Exp CFD 

CN 3.110 3.087 2.950 2.906 2.773 2.735 

CS 0.004 0.000 0.255 0.204 –0.005 0.000 

CA 0.442 0.428 0.441 0.424 0.435 0.419 

CL 2.911 2.891 2.756 2.718 2.585 2.552 

CD 1.181 1.162 1.142 1.115 1.093 1.069 

CL/CD 2.465 2.487 2.413 2.438 2.365 2.389 

Cm –20.301 –20.634 –18.873 –18.955 –17.137 –17.271 

Cn –0.064 –0.001 –2.343 –1.883 –0.007 0.000 

Cl 0.000 0.000 0.0253 0.023 –0.003 0.000 

Xcp 6.528 6.685 6.398 6.522 6.180 6.315 
14

Table 7. Aerodynamic coefficients for side fin configuration at Mach 

2.5, φ = 0°, α = 20°. 

Coefficient Exp CFD 

CN 6.784 6.714 

CS –0.017 0.000 

CA 0.526 0.510 

CL 6.195 6.135 

CD 2.815 2.776 

CL/CD 2.201 2.210 

Cm –45.751 –45.543 

Cn 0.138 –0.001 

Cl 0.000 0.000 

Xcp 6.744 6.783 

Figure 9. Axial force coefficient as a function of roll orientation. 

constant Mach number of 2.5 caused an increase in CA. For the roll angles investigated 

computationally, regardless of Mach number or angle of attack, there is a small, but consistent, 

under prediction of the experimental CA. Perhaps the boundary layer is not completely resolved 

or a different turbulence model should have been used. 

The computed CN agrees very well with the experimental data (figure 10). The maximum 

normal force occurs for φ = 0°, regardless of the fin configuration and decreases as the body is 

rolled. Fin location has a negligible effect on CN, as CN for a given Mach number and angle of 
15

Figure 10. Normal force coefficient as a function of roll orientation. 

attack is nearly identical regardless of the fin location. This implies that body orientation has a 

greater influence on CN that does fin configuration. The increase in CN with angle of attack 

was expected. 

At Mach 4.5, CS does not vary significantly with fin location (figure 11). At Mach 2.5, CS

increases a greater amount (to a 10% maximum) for the corner fin projectile than the side fin 

projectile as it rolls from the φ = 0° orientation to the φ = 22.5° position. The difference then 

decreases as the roll progresses from φ = 22.5° to φ = 45°, suggesting that the flow is affected by 

the asymmetries. The computational results show the same trends as the experimental data. 

CFD accurately predicts the lack of side force at φ = 0° and φ = 45°. However, there seems to be 

an offset from the experimental data at φ= 22.5°. Perhaps the shock interactions are not fully 

resolved. 

The rolling moment coefficient, Cl, shows good agreement between the CFD and experimental 

data at Mach 4.5 (figure 12). At Mach 2.5, however, there seems to be a small discrepancy 

φ = 22.5°. The Mach 2.5, α = 14° side fin configuration shows roll in the opposite direction 

from the rest of the cases investigated. This is likely due to large differences in vortex 

development for this case. The remainder of the data is as expected, with larger rolling moment 

coefficients existing for higher Mach numbers, angles of attack, and deviation from a symmetric 

roll orientation. 
16

Figure 11. Side force coefficient as a function of roll orientation. 

Figure 12. Rolling moment coefficient as a function of roll orientation. 
17

The pitching moment coefficient, Cm, also shows good agreement between the CFD and 

experimental data (figure 13). The trend of these coefficients appears to be a mirror image of the 

normal force coefficient trend, which would be expected, and there is little variation with fin 

location. It is interesting that the largest magnitude pitching moment coefficients occur at φ = 0° 

However, it is not unexpected that the largest variation with roll orientation occurs for the largest 

angle of attack. Unfortunately, the experimental data was provided with a moment reference 

point of the nose. This does not allow for a good determination of the dominating flow 

characteristics (most likely the tail) or the true sensitivity of the pitching moment coefficient. 

The yawing moment coefficient, Cn (figure 14), show trends similar to those of the side force 

coefficient. The largest magnitude coefficient occurs at φ = 22.5° and decreases as the 

symmetric roll orientations are approached. There is a larger difference in the coefficient at the 

same roll orientation between the two fin locations for Mach 2.5 than Mach 4.5. There is less 

yawing moment at Mach 4.5 than Mach 2.5 for a given angle of attack. Although not very 

noticeable in figure 14, tables 3–7 show that the CFD is underpredicting the yawing moment 

coefficient by as much as 0.5. This is as likely due to experimental error (there appears to be an 

offset in Cn as its value is not exactly zero when one expects it to be) as due to discrepancies in 

the computations. 

Figure 13. Pitching moment coefficient as a function of roll orientation. 
18

Figure 14. Yawing moment coefficient as a function of roll orientation. 

4.3 Pressure Coefficient Distributions 

Surface pressure coefficient, Cp, distributions were obtained for each simulation at four axial 

body positions with respect to the reference diameter of the body, D; x/D = 5.5, 8.5, 11.5, and 

12.5. These locations correspond to one-third and two-thirds of the way down stream on the 

body, the midpoint of the root chord of the fin and aft of the fin, respectively (figure 15). The 

numerical results were compared to the results of the wind tunnel tests for each α = 14° 

simulation at the x/D = 5.5 and 8.5 locations with fair success. For each configuration, two 

experimental data sets were available. As a measure of experimental error, both data sets are 

plotted (labeled exp1 and exp2 in all Cp figures). 

Figure 15. Reference axial body position locations on corner fin configuration. 
19

Figures 16–18 show these results for the x/D = 5.5 location. As expected, the pressure 

distributions of the side and corner fin geometries are the same for each configuration at these 

x/D locations as they are located upstream up stream of the fins. At φ = 0° (figure 16), the 

pressure distribution is quite accurately predicted on the leeward side regardless of Mach 

number. On the windward side of the projectile, the trends are fairly well predicted, but the 

magnitude is slightly high. For the φ = 22.5° (figure 17) and φ = 45° (figure 18) roll orientations, 

the pressure distributions are more accurately predicted at Mach 4.5 than at Mach 2.5. This is 

likely because the gradients are steeper for Mach 2.5. The trends for the x/D = 8.5 location are 

nearly the same at for x/D = 5.5. Figures 19–21 show the results for the x/D = 8.5 location. 

Figure 16. Pressure coefficient distribution at x/D = 5.5, φ = 0°, α = 14°. 

For x/D = 11.5 and x/D = 12.5, no wind tunnel data was provided for the projectile with fins at 

α = 14° in any roll orientation. The numerical results for each roll orientation at x/D = 11.5 are 

shown in figure 22. At this location, Cp was obtained only on the body surface and not on the 

fins. As such, data is plotted with symbols and the discontinuity in the data is the location of the 

fins. For the φ = 0° roll orientation (figure 22a), the surface pressures are consistently greater at 

Mach 2.5 than at Mach 4.5 for the side fin configuration, while the inverse in true for the corner 

fin configuration. Additionally, the pressure coefficient for the side fin configuration is 

consistently greater than the corner fin configuration regardless of Mach number except in the 

vicinity of the fin. As the roll orientation is changed to φ = 22.5° (figure 22b), the flow becomes 

very complicated as it is completely asymmetric. This complexity carries over to the surface 
20

Figure 17. Pressure coefficient distribution at x/D = 5.5, φ = 22.5°, α = 14°. 

Figure 18. Pressure coefficient distribution at x/D = 5.5, φ = 45°, α = 14°.
21

Figure 19. Pressure coefficient distribution at x/D = 8.5, φ = 0°, α = 14°. 

Figure 20. Pressure coefficient distribution at x/D = 8.5, φ = 22.5°, α = 14°. 
22

Figure 21. Pressure coefficient distribution at x/D = 8.5, φ = 45°, α = 14°. 

pressure distributions as can be seen in figure 22b. The one noticeable difference at this roll 

orientation is that the surface pressure coefficients at Mach 4.5 are almost always greater than 

those at Mach 2.5 for given configuration. At φ = 45° , the flow becomes more symmetrical. In 

terms of surface pressure distribution (figure 22c), symmetry exists about θ = 145° if one looks 

at 0° < θ < 270°. 

As one moves downstream of the fins to x/D = 12.5, the surface pressure coefficient distributions 

once again become easier to interpret, though the influence of the fins is still present as would be 

expected (figure 23). Areas of increased surface pressure remain at the surface locations where 

the fin were located. This causes “bulges” in the curves. 

Experimental Cp distributions were not available for the Mach 2.5, φ = 0°, α = 20° side fin 

configuration either. The numerical results are compared to the Mach 2.5, φ = 0°, α = 14° side 

fin configuration at each x/D location in figure 24. While the surface pressure distributions 

remain symmetric due to body orientation, there is a significant change in Cp along the entire 

length of the body as the angle of attack is increased from 14° to 20° for the same flow 

conditions and model orientation. Over the top side of the body (0° ≤ θ < 90° and 270° ≤ θ < 

360°), the pressure coefficient is more negative for the higher angle of attack, while on the 

bottom side of the body (90° ≤ θ ≤ 270°), the pressure coefficient is more positive. This 

confirms the increase in lift that is generated at the higher angle of attack. 
23

Figure 22. Pressure coefficient distribution at x/D = 11.5. (a) φ = 0°, α = 14°, (b) φ = 22.5°, α = 4°, 

(c) φ = 45°, α = 14°. 

4.4 Flowfield Visualizations 

Flow field visualizations were completed for each configuration, orientation, and flow conditions 

at the same four axial body position locations for which the pressure coefficient distributions 

were obtained (i.e., x/D = 5.5, 8.5, 11.5, and 12.5). Figures 25–30 show 3-D plots of pitot 

pressure and vorticity for both fin configurations at all three roll angles for the Mach 2.5 flow. 

The vortex cores are seen in both the pitot pressure contours as well as the vorticity contours. 

The contours are symmetric, as expected for φ = 0° and φ = 45° and asymmetric for φ = 22.5°, 

regardless of fin location. For all configurations, the vortex cores grow and move away from the 

body as they move downstream. 

Figure 31 shows the results of the Mach 2.5 flow at 20° angle of attack. The vortex cores are 

much larger and more elongated than the 14° angle-of-attack case. The vortex cores also appear 

to become more disorganized and weaker as they pass over the fins than at the lower angle of 

attack. 

Figures 32–37 show the results of the Mach 4.5 flow. The overall flow structure remains 

practically the same as for the Mach 2.5 flow. The biggest difference that one notices is that the 
24

Figure 23. Pressure coefficient distribution at x/D = 12.5. (a) φ = 0°, α = 14°, (b) φ = 22.5°, α = 14°, 

(c) φ = 45°, α = 14°. 

vortex cores are no longer circular upstream of the fins (x/D = 5.5 and 8.5) as they were at Mach 

2.5, α = 14°. As these figures again show that the flow structure can easily be seen in either the 

pitot pressure contours or the vorticity contours, either vorticity contours or pressure contours are 

presented for the remainder of this section. 

The visualizations for x/D = 5.5 and 8.5 (upstream of the fins) do not vary with fin location 

(figures 38 and 39, respectively) as the flow is supersonic and therefore the downstream 

geometric differences do not effect the upstream flow. However, the flow structure does vary 

with roll (figures 38 and 39), angle of attack (figure 40), and Mach number (figure 41) as would 

be expected. The visualizations at the two further aft positions are much more interesting due to 

the influence of the geometric influences. 

If one first compares the effect of fin location on pressure and vorticity contours at x/D = 11.5, it 

is easy to see that the areas of high pressure are limited to a very small region near the fins, 

regardless of roll orientation (figure 42), while the areas of low pressure, large circulation, and 

vorticity depend more on roll orientation than fin location (figure 43). The pitch-plane symmetry 

of the flow at φ = 0° and φ = 45° is quite noticeable as is the complete asymmetry at φ = 22.5° 

regardless of fin position.
25

Figure 24. Pressure coefficient distribution for Mach 2.5 and φ = 0° at (a) x/D = 5.5, (b) x/D = 8.5, 

(c) x/D = 11.5, and (d) x/D = 12.5. 

Figure 25. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for corner fin 

configuration, Mach 2.5, φ = 0°, α = 14°. 
26

Figure 26. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for side fin 

configuration, Mach 2.5, φ = 0°, α = 14°. 

Figure 27. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for corner fin 

configuration, Mach 2.5, φ = 22.5°, α = 14°. 

Figure 28. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for side fin 

configuration, Mach 2.5, φ = 22.5°, α = 14°.
27

Figure 29. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for corner fin 

configuration, Mach 2.5, φ = 45°, α = 14°. 

Figure 30. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for side fin 

configuration, Mach 2.5, φ = 45°, α = 14°. 

Figure 31. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for side fin 

configuration, Mach 2.5, φ = 0°, α = 20°. 
28

Figure 32. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for corner fin 

configuration, Mach 4.5, φ = 0°, α = 14°. 

Figure 33. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for side fin 

configuration, Mach 4.5, φ = 0°, α = 14°. 

Figure 34. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for corner fin 

configuration, Mach 4.5, φ = 22.5°, α = 14°.
29

Figure 35. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for side fin 

configuration, Mach 4.5, φ = 22.5°, α = 14°. 

Figure 36. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for corner fin 

configuration, Mach 4.5, φ = 45°, α = 14°. 

Figure 37. Three-dimensional (a) pitot pressure contours and (b) vorticity contours for side fin 

configuration, Mach 4.5, φ = 45°, α = 14°. 
30

Figure 38. Vorticity contour plots at x/D = 5.5, Mach 2.5, α = 14° for ([a]–[c]) corner fin configuration 

φ = 0°, 22.5°, 45°, respectively, and ([d]–[f]) side fin configuration, φ = 0°, 22.5°, 45°, 

respectively. 

Figure 39. Vorticity contour plots at x/D = 8.5, Mach 2.5, α = 14° for ([a]–[c]) corner fin configuration 

φ = 0°, 22.5°, 45°, respectively, and ([d]–[f]) side fin configuration, φ = 0°, 22.5°, 45°, 

respectively. 
31

Figure 40. Vorticity contour plots for side fin configuration at Mach 2.5, φ = 0°, ([a]–[b]) α = 14°, 

x/D = 5.5 and 8.5, respectively, and ([c]–[d]) α = 20°, x/D = 5.5 and 8.5, respectively. 

Figure 41. Vorticity contour plots at x/D = 8.5, α = 14° for corner fin configuration ([a]–[c]) Mach 2.5, 

φ = 0°, 22.5°, 45°, respectively, and ([d]–[f]) Mach 4.5, φ = 0°, 22.5°, 45°, respectively. 
32

Figure 42. Pressure contours for ([a]–[c]) corner fin configuration and ([d]–[f]) side fin configuration at 

roll orientations indicated Mach 2.5, α = 14°, x/D = 11.5. 

Figure 43. Vorticity contours for ([a] – [c]) corner fin configuration and ([d] – [f]) side fin configuration at 

roll orientations indicated for Mach 2.5, α = 14°, x/D = 11.5.
33

Comparing vorticity contours of Mach 2.5 and 4.5 flows for α = 14° at x/D = 11.5 (figure 44) 

indicates that once again the magnitude of the circulation increases, but the basic vortex structure 

does not. At x/D = 12.5 (figure 45), the vorticity magnitude as well as the asymmetry for 

φ = 22.5° has further increased. At φ = 0° and φ = 45°, pitch-plane asymmetries do not occur 

downstream of the fins as the fins themselves are symmetric. Experimental vapor screen images 

of the corner fin configuration were available for comparison to the computational vorticity 

contours at x/D = 12.5 and are shown in figure 46 (17). The similarities in the vortex cores for 

all three roll orientations and Mach numbers strongly suggest that the CFD is correctly predicting 

the flow field for the corner fin configuration. Comparison of the CFD vorticity contours at 

x/D = 12.5 and the experimental vapor screen images (17) for the side fin configuration show 

similarly good agreement. 

Figure 44. Vorticity contours for corner fin configuration at roll orientations indicated, α = 14°, x/D = 11.5 

for ([a]–[c]) Mach 2.5 and ([d]–[f]) Mach 4.5. 

At both x/D = 11.5 (figure 47) and x/D = 12.5 (figure 48), an increase in α from 14° to 20° for 

Mach 2.5, α = 0°, did not cause the basic vortex structure around most of the projectile to vary, 

only the sphere of influence. However, the leeward fins cause the vortex cores to become less 

concentrated at the higher angle of attack as they pass over the fins (figure 47). This 

disorganization remains apparent as the flow moves downstream of the fins (figure 48b). This 

indicated that one possible way to increase the aerodynamic forces acting on the fins would be to 

increase the angle of attack for a given Mach number. 
34

Figure 45. Vorticity contours for corner fin configuration at roll orientations indicated, α = 14°, x/D = 12.5 

for ([a]–[c]) Mach 2.5 and ([d]–[f]) Mach 4.5. 

Figure 46. Experimental vapor screen images for corner fin configuration at roll orientations indicated, 

α = 14°, x/D = 12.5 for ([a]–[c]) Mach 2.5 and ([d]–[f]) Mach 4.5. 
35

Figure 47. Vorticity contours for side fin configuration at Mach 2.5, φ = 0°, x/D = 11.5, (a) α = 14° and (b) 

α = 20°. 

Figure 48. Vorticity contours for side fin configuration at Mach 2.5, α = 0°, x/D = 12.5, (a) α = 14° and 

(b) α = 20°. 

These changes in the flow field (or lack thereof) explain the similarities and differences seen in 

the surface pressure distributions. It would appear that a square cross-section missile with 

pyramid fins, placed either at the corners or on the sides, could be expected to perform quite 

well. Further studies need to be conducted to determine the dynamic behavior of the projectile 

before a conclusion could be fully stated. 

5. Summary and Conclusions 

Numerical computations using viscous Navier-Stokes methods were performed to predict the 

flow field and aerodynamic coefficients on all missile configurations for wind tunnel test 

conditions. Full 3-D computations were performed and no symmetry was used. Computational 
36

results were obtained for the square cross-section missile at supersonic speed for various roll 

orientations and angles of attack using a two-equation k-ε turbulence model. Numerical results 

show the qualitative features (vortices and cross-flow separation regions) of flow field at various 

stream wise positions along the missile configurations. Aerodynamic coefficients have been 

obtained from the computed solutions and found to match well with the available experimental 

data for these configurations. These numerical results show that modern CFD technique such as 

the one used in the present computational study is capable of accurately predicting the 

aerodynamics of complex missile configurations with a square cross section and multiple of fins. 
37

6. References 

1. Sahu, J.; Heavey, K. R.; Ferry, E. N. Computational Fluid Dynamics for Multiple Projectile 

Configurations. Proceedings of the 3rd Overset Composite Grid and Solution Technology 

Symposium, Los Alamos, NM, October 1996. 

2. Sahu, J.; Heavey, K. R.; Nietubicz, C. J. Time-Dependent Navier-Stokes Computations for 

Submunitions in Relative Motion. 6th International Symposium on Computational Fluid 

Dynamics, Lake Tahoe, NV, September 1995. 

3. Meakin, R. L. Computations of the Unsteady Flow About a Generic Wing/Pylon/FinnedStore Configuration. AIAA 92-4568-CP, August 1992. 

4. Pagan, D.; Molton, P.; Delery, J. Basic Experiment on a Supersonic Vortex Flow Around a 

Missile Body. Journal of Spacecraft and Rockets 1992, 29 (3), 373–378. 

5. Grasso, F.; Iaccarino, G. Influence of Crossflow and Turbulence on Vortex Flow Around a 

Supersonic Missile. Journal of Spacecraft and Rockets 1998, 35 (1), 37–45. 

6. Sahu, J.; Heavey, K. R. Application of CFD to High Angle of Attack Missile Flowfields. 

AIAA Paper No. 2000-4210, 18th Applied Aerodynamics Meeting, Denver, CO, August 

2000. 

7. Birch, T.; Wrisdale, I. E.; Prince, S. CFD Predictions of Missile Flowfields. AIAA Paper 

No. 2000-4211, 18th Applied Aerodynamics Meeting, Denver, CO, August 2000. 

8. Birch, T.; Allen, J.; Wilcox, F. J. Force, Surface Pressure and Flowfield Measurements on 

Slender Missiles at Supersonic Speeds. AIAA Paper No. 2000-4207, 18th Applied 

Aerodynamics Meeting, Denver, CO, August 2000. 

9. Edwards, J. A.; Roper, J. J. A Computational Assessment of Static and Dynamic 

Coefficients for the H3 Hypervelocity Projectile. AIAA Paper 97-0640, January 1997. 

10. Orchard, D.; Fournier, E.; Dupuis, A.; Edwards, J. Wind Tunnel Tests on the H3P78, Power 

Law, Elliptic Section Flared Projectile From Mach 2.5 to 4. AIAA Paper No. 2001-4321, 

Montreal, Canada, August 2001. 

11. Fournier, E.; Orchard, D. Testing of a Novel Maneuvering Projectile in the DREV Trisonic 

Wind Tunnel; technical memorandum report, DREV TM - 2000-108; Canada, July 2000. 

12. Edge, H. L.; Clarke, J. Graphical User Interface for ZEUS; ARL-TR-1443; U.S. Army 

Research Laboratory: Aberdeen Proving Ground, MD, June 1996. 
38

13. Edge, H. L.; Sahu, J.; Sturek, W. B.; Pressel, D. M.; Heavey, K. R. Common High 

Performance Computing Software Support Initiative (CHSSI) Computational Fluid 

Dynamics (CFD)-6 Project Final Report: ARL Block-Structured Gridding Zonal NavierStokes (ZNSFLOW) Solver Software; ARL-TR-2084; U.S. Army Research Laboratory: 

Aberdeen Proving Ground, MD, February 2000. 

14. Sahu, J.; Heavey, K. R.; Edge, H. L. Numerical Computations of Supersonic and 

Hypersonic Flow over Elliptical Projectiles. AIAA Paper No. 2001-4320, Atmospheric 

Flight Mechanics Conference: Montreal, Canada, August 2001. 

15. Orchard, D.; Fournier, E.; Dupuis, A.; Edwards, J. Wind Tunnel Tests on the H3P78, Power 

Law, Elliptic Section Flared Projectile with Jet Interaction. AIAA Paper No. 2001-4322, 

Montreal, Canada, August 2001. 

16. Sahu, J. Numerical Simulations of Supersonic Flow over an Elliptic-Section Projectile with 

Jet Interaction. AIAA Paper No. 2002-3260, AIAA Applied Aerodynamics Conference: 

St. Louis, MO, June 2002. 

17. Wilcox, F.; Birch, T.; Allen, J. Force, Surface Pressure, and Flowfield Measurements on a 

Slender Missile Configuration with Square Cross- Section at Supersonic Speeds. AIAA 

Paper No. 2004-5451, Applied Aerodynamics Meeting: Providence, RI, August 2004. 

18. Peroomian, O.; Chakravarthy, S.; Goldberg, U. A ‘Grid-Transparent’ Methodology for 

CFD. AIAA Paper 97-07245, 1997. 

19. Peroomian, O.; Chakravarthy, S.; Palaniswamy, S.; Goldberg, U. Convergence Acceleration 

for Unified-Grid Formulation Using Preconditioned Implicit Relaxation. AIAA Paper 98-

0116, 1998. 

20. Pulliam, T. H.; Steger, J. L. On Implicit Finite-Difference Simulations of ThreeDimensional Flow. AIAA Journal 1982, 18 (2), 159–167. 

21. Goldberg, U. C.; Peroomian, O.; Chakravarthy, S. A Wall-Distance-Free k-ε Model 

with Enhanced Near-Wall Treatment. ASME Journal of Fluids Engineering 1998 120, 457–

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22. ICEM CFD Engineering, ICEM CFD v3.5.1 User’s Manual, Berkeley, CA, 2000. 
NO. OF 

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 AFATL FXA 

 DAVE BELK 

 EGLIN AFB FL 32542-5434 

 3 COMMANDER 

 US ARMY TACOM ARDEC 

 AMSTA AR FSF T 

 BLDG 382 

 H HUDGINS 

 J GRAU 

 W KOENIG 

 PICATINNY ARSENAL NJ 

 07806-5000 

1 COMMANDER 

 US ARMY TACOM 

 AMSTA AR CCH B 

 P VALENTI 

 BLDG 65S 

 PICATINNY ARSENAL NJ 

 07806-5001 

 1 COMMANDER 

 US ARMY ARDEC 

 SFAE FAS SD 

 J HEDDRICTL 

 PICATINNY ARSENAL NJ 

 07806-5001 

1 NAWC 

 D FINDLAY 

 MS 3 BLDG 2187 

 PATUXENT RIVER MD 20670 

 

 1 UNIV OF ILLINOIS AT URBANA 

 CHAMPAIGN 

 DEPT OF MECHANICAL AND 

 INDUSTRIAL ENGRNG 

 J C DUTTON 

 URBANA IL 61801 

 1 COMMANDER 

 US ARMY TACOM-ARDEC 

 BLDG 162S 

 AMCPM DS MO 

 P BURKE 

 PICATINNY ARSENAL NJ 

 07806-5000 

ABERDEEN PROVING GROUND

 19 DIR AMSRD ARL 

 AMSRD ARL WM 

 J SMITH 

 AMSRD ARL WM B 

 T KOGLER 

 W CIEPIELA 

 AMSRD ARL WM BA 

 D LYON 

 AMSRD ARL WM BD 

 B FORCH 

 AMSRD ARL WM BC 

 P PLOSTINS 

 J DESPIRITO 

 B GUIDOS 

 K HEAVEY (3 CPS) 

 J NEWILL 

 J SAHU 

 S SILTON 

 P WEINACHT 

 AMSRD ARL WM BF 

 S WILKERSON 

 H EDGE 

 AMSRD ARL SL BE 

 A MIKHAIL 

 AMSRD ARL CI HC 

 R NOAK 
NO. OF 

COPIES ORGANIZATION

41

 1 J EDWARDS 

 MCM DEPARTMENT 

 DSTL FORT HALSTEAD 

 SEVEN OAKS 

 KENT NT14 7BP 

 UK 
42 

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