March 2011 

NASA/TM–2011-214528 

Cavity Heating Experiments Supporting 

Shuttle Columbia Accident Investigation 

Joel L. Everhart, Karen T. Berger, Kim S. Bey, N. Ronald Merski, and William A. Wood 

Langley Research Center, Hampton, Virginia 
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March 2011 

NASA/TM–2011-214528 

Cavity Heating Experiments Supporting 

Shuttle Columbia Accident Investigation 

Joel L. Everhart, Karen T. Berger, Kim S. Bey, N. Ronald Merski, and William A. Wood 

Langley Research Center, Hampton, Virginia 
Available from: 

NASA Center for AeroSpace Information 

7115 Standard Drive 

Hanover, MD 21076-1320 

443-757-5802 

Acknowledgments 

The authors gratefully acknowledge the contributions of the following individuals to this 

research effort: Buddy Stotler, Grace Gleason, Roland Hatten, Kevin Hollingsworth, and 

Teck-Seng Kwa for operation of the 20-Inch Mach 6 Air Tunnel and data acquisition 

support; Gary Wainwright and Steve Nevins for model CAD design and fabrication of 

the stereolithography models; Michael Powers, Mark Griffith, Hayward C. Glaspell for 

fabrication of the ceramic test models; Ed Covington, Kathryn Kuykendoll, and Pete 

Veneris for model fidelity measurements and fiducial mark placement; and, Gregory M. 

Buck for the model oil flow visualization runs. 

The use of trademarks or names of manufacturers in this report is for accurate reporting and does not 

constitute an official endorsement, either expressed or implied, of such products or manufacturers by the 

National Aeronautics and Space Administration. 
 

5

TABLE OF CONTENTS 

PAGE 

List of Tables ................................................................................................................ 6 

List of Figures ............................................................................................................... 7 

Abstract ........................................................................................................................ 9 

Introduction ................................................................................................................. 10 

Nomenclature .............................................................................................................. 12 

Approach .................................................................................................................... 13 

Supersonic/Hypersonic Cavity Flow Physics .............................................................. 14 

Flight-to-Tunnel Scaling Methodology ........................................................................ 15 

Experimental Methods ................................................................................................ 16 

Test Facility ................................................................................................................ 17 

Test Model Description ............................................................................................... 18 

 Oil Flow Visualization Technique ..................................................................... 19 

 Phosphor Thermography Technique ............................................................... 19 

Experimental Results .................................................................................................. 20 

 Oil Flow Visualization ....................................................................................... 20 

 Surface Heating Data ....................................................................................... 22 

 Phosphor Uncertainty Analysis ........................................................................ 26 

Summary .................................................................................................................... 28 

References ................................................................................................................. 30 

Tables ......................................................................................................................... 37 

Figures ........................................................................................................................ 52 

Appendix: Derivation of Pressure Gradient Surface Geometry ................................ 104 
 6

LIST OF TABLES 

 PAGE 

TABLE 1.- Flight Scaled Cavities and Conditions. ...................................................... 37 

TABLE 2.- Nominal flow conditions in the 20-Inch Mach 6 Air Tunnel. ...................... 37 

TABLE 3.- Cavity models for Test 6857 - Columbia accident investigation. ............... 38 

TABLE 4.- Cavity models for Test 6868 - Closed cavity geometry parameters. ......... 39 

TABLE 5.- Test matrix for Test 6857 - Oil flow runs. .................................................. 40 

TABLE 6.- Test matrix for Test 6857 - Baseline model heating runs. ........................ 41 

TABLE 7.- Test matrix for Test 6857 - Cavity heating runs. ....................................... 42 

TABLE 8.- Test matrix for Test 6868 – Baseline model heating runs. ........................ 46 

TABLE 9.- Test matrix for Test 6868 - Cavity heating runs. ....................................... 47 

TABLE 10.- Test matrix for Test 6868 - Pressure gradient effects on heating 

using SLA models. ................................................................................... 50 

TABLE 11.- Test matrix for Test 6868 - Cross-flow effect on heating using 

SLA models. ............................................................................................ 51 

TABLE 12.- Phosphor uncertainty estimates. ............................................................. 51 

 
 

7

LIST OF FIGURES 

 PAGE 

Figure 1.- Combined experimental and computational simulation approach. ............. 52 

Figure 2.- Classification of supersonic/hypersonic cavity flows. ................................. 52 

Figure 3.- Sketch of model coordinate system. .......................................................... 53 

Figure 4.- LAURA predictions for Me versus a for baseline model. ............................ 53 

Figure 5.- LAURA predictions for Reθ versus a for baseline model. ........................... 54 

Figure 6.- LAURA predictions for boundary layer thickness for baseline model. ........ 54 

Figure 7.- Langley 20-Inch Mach 6 Air Tunnel. ........................................................... 55 

Figure 8.- Baseline model - 4-in. by 6-in. inserts used in Test 6857. ......................... 55 

Figure 9.- Baseline model - 4-in. by 18-in. inserts used in Test 6868. ....................... 55 

Figure 10.- Model fabrication process. ....................................................................... 56 

Figure 11. – Cavity insert photographs for Test 6857 ................................................ 57 

Figure 12. – Cavity insert model photographs for Test 6868 ...................................... 61 

Figure 13.- Oil flow images for Test 6857 ................................................................... 68 

Figure 14.- Oil flow visualization of cavity width effect - L/H=14, H/δ=1.1. ................. 70 

Figure 15.- Baseline heating runs for Test 6857. ....................................................... 71 

Figure 16.- Open flow cavity heating runs for Test 6857. ........................................... 73 

Figure 17.- Cavity alignment effect on open flow cavity heating for Test 6857. ......... 79 

Figure 18.- Closed flow cavity heating runs for Test 6857. ......................................... 81 

Figure 19.- Baseline heating runs for Test 6868. ....................................................... 88 

Figure 20.- Open (short) cavity heating runs for Test 6868. ....................................... 90 

Figure 21.- Length effect on closed cavity heating for Test 6868 – 

Reθ=300, H/δ=1.1. .................................................................................... 91 

Figure 22.- Length effect on closed cavity heating for Test 6868 – 

Reθ=300, H/δ=2.39. .................................................................................. 93 

Figure 23.- Length effect on closed cavity heating for Test 6868 – 

Reθ=500, H/δ=1.1. .................................................................................... 94 

Figure 24.- Length effect on closed cavity heating for Test 6868 – 

Reθ=500, H/δ=2.39. .................................................................................. 95 
 8

Figure 25.- Depth effect on closed cavity heating for Test 6868 – 

Reθ=503, H/δ=2.39. .................................................................................. 96 

Figure 26.- Depth effect on closed cavity heating for Test 6868 – 

Reθ=503, H/δ=5.20. .................................................................................. 98 

Figure 27- Depth effect on closed cavity heating for Test 6868 – 

Reθ=260, H/δ=0.65 ................................................................................... 99 

Figure 28.- Depth effect on closed cavity heating for Test 6868 – 

Reθ=260, H/δ=2.7. .................................................................................. 100 

Figure 29.- Pressure gradient effect on closed cavities for Test 6868. ..................... 101 

Figure 30.- Cross flow effect on closed cavities for Test 6868 – 

Reθ=300, H/δ=1.1 ................................................................................... 102 

Figure 31.- Cross flow effect on closed cavities for Test 6868 – 

Reθ=500, H/δ=2.39. ................................................................................ 103 

 
 

9

Abstract 

The two-color thermographic phosphor method has been used to map the local 

heating augmentation of scaled idealized cavities at conditions simulating the windward 

surface of the Shuttle Orbiter Columbia during flight STS-107. Two experiments initiated 

in support of the Columbia Accident Investigation were conducted in the Langley 20-

Inch Mach 6 Tunnel. Generally, the first test series evaluated open (length-to-depth less 

than 10) rectangular cavity geometries proposed as possible damage scenarios resulting from foam and ice impact during launch at several discrete locations on the vehicle 

windward surface, though some closed (length-to-depth greater than 13) geometries 

were briefly examined. The second test series was designed to parametrically evaluate 

heating augmentation in closed rectangular cavities. The tests were conducted under 

laminar cavity entry conditions over a range of local boundary layer edge-flow parameters typical of re-entry. Cavity design parameters were developed using laminar computational predictions, while the experimental boundary layer state conditions were inferred from the heating measurements. An analysis of the aeroheating caused by cavities allowed exclusion of non-breeching damage from the possible loss scenarios being 

considered during the investigation. 

 
 10

Introduction 

On February 1, 2003 the Space Shuttle Columbia, flight STS-107, was lost during 

re-entry. Many possible damage scenarios were examined during the ensuing accident 

investigation, among them the creation of impact-induced cavities in the tile thermal protection system (TPS), resulting in local augmentation of the heating and eventual burn 

through of the aluminum substructure. The purpose of this report is to document two 

wind tunnel cavity heating experiments conducted in the NASA Langley 20-Inch Mach 6 

Wind Tunnel to support the Columbia Accident Investigation Board (CAIB) and to present foundational cavity heating information supporting the follow-on Return-to-Flight 

(RTF) program established to implement CAIB recommendations. This document will 

outline the adopted testing approach and the flight-to-tunnel scaling philosophy. Data 

presented herein are in image format; accordingly, only qualitative observations are 

made concerning their behavior. 

Defining the heating environment of a damage-scenario cavity presents a significant challenge – how is the cavity defined? Particularly when few (if any) direct measurements of the cavity geometry are available and when the exact location of the cavity 

(implying local flow conditions) may be unknown. Many possible variables and parameters exist. The gross geometric parameter space includes: length (L), width (W), and 

depth (H), and planform, cross-sectional, and profile variations. Sidewall and upstream 

endwall entry angles may play an important role on the flow expansion into the cavity, 

while sidewall and endwall exit angles may affect recompression on the downstream 

wall and further expansion around the end-wall corner – all parameters having a direct 

influence on the heating. The depth distribution along the cavity profile may potentially 

affect how or if the flow enters the cavity and impinges on the floor. The impact of 

roughness caused by irregular surfaces, protuberances, gap fillers between tiles (present or missing), embedded objects, etc. is unknown, but the roughness may significantly impact the shear layer/boundary layer transition process. Local boundary layer edgeflow conditions include Mach number (Me), momentum thickness Reynolds number 

(Reθ), boundary layer thickness (δ), stream direction with respect to the cavity major axis, chemistry effects reflected in the ratio of specific heats (γ), wall temperature ratio 

(Tw/Taw), and boundary layer state (laminar, transitional, or turbulent) entering and leaving the cavity. 

The supersonic/hypersonic cavity flow literature (particularly that for laminar test 

conditions), while helpful, is particularly sparse on both the local and downstream effects of many fundamental flow parameters. Fletcher, et al. published the survey paper 

“A Review of Heat Transfer in Separated and Reattached Flows” in 1970. Nestler updated this survey with more recent work in 1985 in “The Effects of Surface Discontinuities on Convective Heat Transfer in Hypersonic Flow.” Together, these papers include 

much of the existing, pertinent cavity flow literature. They cited theoretical models of 
 

11

cavity flow developed by Burggraf (1965), Chapman (1956), Carlson (1959), Chang 

(1966), and Lamb (1967); however, none of the theories adequately capture the physics 

of laminar and turbulent three-dimensional cavity flows as will be demonstrated via 

global surface heating distributions presented herein, though they offer a framework for 

experimental studies. Numerical studies were conducted by Adams (1974), Morgenstern and Chokani (1994), and Zhang, et al. (2001). More recently, additional numerical 

studies have been added for the CAIB and the Shuttle RTF program by Wood, et al. 

(2004) and Pulsonetti, et al. (2005). Experimentally, laminar two-dimensional flows were 

addressed by Galenter (1975) and Nestler (1966); laminar axisymmetric flows were addressed by Nestler (1970), Nicoll (1963, 1964), and Wyborny, et al. (1967); and, laminar 

three-dimensional flows were examined by Cheatwood, et al. (2001), Hahn (1969), and 

Nestler (1981). Experimental studies with cavities tested in transitional boundary layers 

or to determine the onset of transition were presented by Charbonnier and Boerrigter 

(1993), Boerrigter and Charbonnier (1996), Hollis and Liechty (2002, 2006), Larson and 

Keating (1960), Liechty, et al. (2006a), and Rhudy and Magnan (1964). Turbulent twodimensional experiments were presented by Charwat, et al. (1961a, 1961b), Chin and 

Seban (1972), Emery (1969), Hunt (1977), Lamb (1968, 1980a, 1980b), Nestler (1968, 

1969, 1970a, 1970b), Shchukin, et al. (1980), and Stallings and Wilcox (1987); turbulent 

axisymmetric experiments were conducted by Hunt (1974) and Netterfield (1989); and 

turbulent three-dimensional experiments were published by Wilcox (1990, 1991). In 

general, the heating measurements were obtained with sparsely spaced discrete sensors, the exception being those global phosphor thermography measurements presented by Cheatwood, et al., and Hollis and Liechty for circular cavities. Because of this spatial measurement sparseness, much of the three-dimensional nature of the surface 

heating profile is missed, as is in many cases the important peak heating value. While 

most of the idealized cavity geometries are rectangular with a flat bottom profile (a few 

have an arc bottom), the paper by Emery is of particular significance because heating 

profiles are presented for models with notched, cutback profile geometries that may be 

more representative of potential impact damage conditions. Two papers by Blair and 

Stallings (1986) and Stallings, et al. (1991) are significant because they present oil flow 

visualizations in and around the cavity and vapor screen visualizations of the cavity 

crossflow plane, clearly showing the growth and development of the cavity vortex structure. Similar, limited oil flow results for the current experiments will be presented herein. 

The important class of cavities known as gaps (i.e., cavities with length-to-depth ratio 

less than 1) has been examined by Coats, et al. (1970), Johnson (1973), Throckmorton 

(1974), and Weinstein, et al. (1975). In-gap heating measurements were not acquired 

during the current experiments, though a limited number of runs were made with gaps 

extending downstream of the cavity to simulate the loss of gap fillers to evaluate their 

impact on the cavity heating levels. 

As discussed above, there exists only a limited amount of laminar flow experimental data for evaluating heating effects in supersonic/hypersonic cavities. Additionally, most of the existing parametric variations and correlations have been developed using turbulent heating measurements. Accordingly, turbulent methods have typically been 
 12

used for thermal assessments of damage on the Space Shuttle, even though much of 

the re-entry trajectory provides laminar edge conditions on the vehicle. The use of these 

turbulent methods may be overly conservative. An objective of these experiments is to 

develop a cavity-heating database with laminar entry and exit conditions that will, hopefully, provide an additional thermal assessment capability for possible damage scenarios. Because of the unknowns existing in the flow physics, a fundamental question is 

whether the cavity flow definitions presented in the literature are even valid for the hypersonic laminar cavity. An objective of these experiments is to provide information to 

enable an assessment of this question. Finally, computational modeling of chemically 

reacting hypersonic flows is difficult; for potential damage scenarios it becomes extremely challenging to separate the flow physics effects, and then validate the solution. 

A further objective of these experiments is to provide parametric information to reduce 

the risks associated with numerical simulations. 

Nomenclature 

B bias uncertainty 

BF Bump Factor 

BFmax maximum Bump Factor on the cavity end wall 

Cp pressure coefficient 

c specific heat (BTU/lbm- ̊F) 

CAIB Columbia Accident Investigation Board 

h heat transfer coefficient, h=q/(Haw-Hw), (lbm/ft2

/s) 

H enthalpy (btu/lbm) 

H, W, L cavity depth, width, length (in) 

k thermal conductivity (BTU/hr-ft- ̊F) 

M Mach number 

P precision uncertainty 

p pressure (psi) 

Re Reynolds number per foot (1/ft) 

Reθ momentum thickness Reynolds number 

RTF Return-to-Flight 

T temperature ( ̊R) 

t time (s) 

TGP thermographic phosphor 

TPS thermal protection system 

x axial distance from nose of model (in) 

y spanwise distance from centerline of model (in) 

z distance normal to model surface (in) 

α angle-of-attack, nose up is positive (deg) 

β thermal product, (ρck)1/2

γ ratio of specific heats 
 

13

δ boundary layer thickness 

Θ Normalized temperature in heating equation 

θ boundary layer momentum thickness (in) 

ρ density (lbm/ft3) 

Λ Normalized time in heating equation 

λ cavity rotation angle (deg) 

Subscripts 

∞ freestream conditions 

avg average 

aw adiabatic wall conditions 

e boundary layer edge conditions 

FR conditions from Fay-Riddell calculation for a hemisphere 

init initial 

ref reference condition 

t1 reservoir conditions 

w wall conditions 

Approach 

The combined computational and experimental approach taken for this program 

is presented in Figure 1. Since fluid physics and flow chemistry dominated the hypersonic accident environment, full computational simulation using even a simplified Orbiter 

geometry is difficult because of uncertainties inherent in viscous flow phenomena (i.e. 

separation, transition, shock-boundary layer interactions, etc.). High-temperature chemical interactions that define aerothermodynamics further compound the simulation process, particularly when structural combustion may be occurring. Issues such as actual 

damage site geometry and location, and computational grid resolution and convergence 

introduce a multitude of other uncertainties. Ground-based facilities are incapable of 

replicating the required energy levels with full geometric and fluid dynamic simulation. 

However, they can be used to conduct building block experiments wherein the fluid 

phenomena are simulated on simplified, scaled geometries. These experiments may 

then be used to develop an empirical understanding, ultimately leading to mathematical 

relationships and correlations for modeling the heating augmentation produced by a 

possible damage scenario, and they may also be used to assess and improve the ability 

of computer codes to suitably capture the flow physics. An assessment of flow chemistry effects can be made after validation of the flow physics model, which lends credence 

to flight scenario computations. 
 14

Supersonic/Hypersonic Cavity Flow Physics 

Based on the literature survey, the following overview of cavity flow physics is 

presented. Length-to-depth ratio (L/H) is typically used to distinguish between and classify different cavity flow regimes, as depicted in Figure 2. Very short or deep cavities 

with L/H < 1 are known as gaps. In this case, shearing induced by the main stream flow 

causes the development within the gap of a column of counter-rotating vortices numbering approximately H/L. Alternating hot spots are developed in the gap when the vortices 

directionally align and impinge on the sidewall. 

Two stable flow conditions exist for cavities with length L/H > 1. The first is a 

short cavity in the length range 1 ≤ L/H ≤ 10, known as an open cavity in the literature. 

The mainstream flow does not enter the cavity directly, though there may be some mass 

interchange with the low-energy vortical flow inside the cavity. Physically, the short cavity has insufficient length to support the required entry and exit turning angles, therefore 

it skims the cavity. In this case, the pressure in the cavity is typically above the ambient 

and climbs to a peak at the downstream lip. The heating drops to values significantly 

below the undisturbed value and rises slowly to a peak value on the downstream lip. 

The other stable solution is a long cavity, also known as a closed cavity, with length L/H 

≥ 14. In this case, three distinct flows may develop if the cavity is long enough. First, the 

upstream flow is now able to turn into the cavity and impinge on the floor, creating an 

aft-facing-step flow field. Next, a boundary layer on the floor may develop and recover to 

the ambient level outside the cavity. Finally, as the flow approaches the end wall it will 

turn outward and create a forward-facing-step flow field. For long, deep cavities, the 

pressure gradients may be severe where the flow turns and strong expansion and shock 

waves will be generated. Viscous shearing generated by this flow turning will augment 

the heating to levels significantly higher than the ambient levels on both the cavity floor 

and the end wall. The pressure in these long cavities will decrease below the ambient 

and steadily increase downstream, reaching large values of over-pressure behind the 

shocks. Vortices will develop on the cavity sidewalls as the flow expands around the 

corner into the cavity and on the floor after flow impingement, further augmenting the 

heating, which may extend laterally around the cavity in the most severe cases. These 

vortices will interact with the cavity end wall and spill into the downstream region beyond 

the cavity. Given laminar inflow, analysis of the heating profiles indicates that most any 

type of outflow may occur, depending on the cavity and its environmental state. The incavity flow may remain laminar, become transitional, or transition to fully turbulent flow; 

the downstream possibilities are equally varied and are currently the subject of extensive analysis and testing (Liechty 2006a) in support of Shuttle RTF. 

Cavities in the range 10 ≤ L/H ≤ 14 are typically unsteady as the flow alternates 

between the two bounding conditions; these are known as transitional cavities. Oil flow 

visualization of both open and closed conditions occurring in the same cavity has been 

presented by Stallings, et. al. (1991). Transitional cavities are avoided where possible in 
 

15

the present tests because of the complexity of the required instrumentation and test 

time necessary to address flow steadiness. 

The boundaries between the different cavity flow regimes are nominal, at best. 

For example, different researchers have measured L/H values ranging from 9 to 11 as 

the upper limit for open cavities and from 12 to 15 as the lower limit for closed cavity 

flow. These limits should therefore be taken only as a guide. Also, it is important to note 

that most of the reported cavity data were acquired in air (γ =1.4) with some in helium 

(γ=1.67). Since turning angle is a function of both Mach number and γ, it is conceivable 

that the open/closed cavity boundaries will vary during re-entry as vehicle boundary layer edge conditions change. At present, experimental data are insufficient to estimate the 

strength of this effect on the aeroheating. At the time of these experiments, the effects 

of pressure gradient on the local heating had not been addressed in the published literature. Some indication regarding this effect can be found in a recent paper by Everhart, 

et. al. (2006). 

Flight-To-Tunnel Scaling Methodology 

The procedure adopted for scaling a cavity from the flight environment to the 

wind tunnel is straightforward. Based on the literature survey, the cavity geometry is defined by its length (L), width (W), depth (H), and orientation (λ), while the pertinent flow 

conditions are the boundary layer edge conditions at the cavity entrance, i.e. the Mach 

number (Me), momentum thickness Reynolds number (Reθ), boundary layer thickness 

(δ), wall temperature ratio (Tw/Taw), and the gas chemistry effects reflected in the ratio of 

specific heats (γ). In flight, the cavity parameter space is given by (Me, Reθ, Tw/Taw, γ, 

H/δ, W/H, L/H, λ)FLIGHT. Similarly, in the wind tunnel the cavity parameter space is given 

by (Me, Reθ, Tw/Taw, γ, H/δ, W/H, L/H,λ)Tunnel. The challenge becomes one of matching 

the parameter spaces. The flight chemistry is difficult to match in a conventional wind 

tunnel, as is the wall temperature ratio. So, for modeling the effects of fluid physics these two parameters are excluded from the experiments, relying on future computational 

fluid simulations for their assessment, leaving 

(Me, Reθ, H/δ, W/H, L/H, λ )FLIGHT = (Me, Reθ, H/δ, W/H, L/H, λ )Tunnel

The accident investigation team prescribed two original candidate damage geometries (length by width by depth) as a missing carrier panel of 24x4x3 inches and 

missing/damaged tiles 18x6x2.5 inches (equivalent to 3 tiles) near the Main Landing 

Gear Door. The boundary layer edge conditions for the shuttle-accident flight environment were computationally determined at different possible damage sites by Pulsonetti 

and Thompson (2004). The flight scaling and conditions at these prescribed locations 

are presented in Table 1. As presented in the table, the largest variations occur in the 

values of Reθ and H/δ with the Me and L/H variations remaining relatively small. Fur-
 16

thermore, since the values of L/H correspond to those in the open-cavity flow regime, 

the effect of this parameter on the heating was expected to be small. The local influence 

of W/H was unknown. The procedure for matching these conditions in the wind tunnel is 

as follows. First, computational solutions of the boundary layer flow on a flat plate baseline model were determined by Wood, et al. (2004) using the LAURA Navier-Stokes 

code (Gnoffo 1990). Referring to the model sketch provided in Figure 3 for the coordinate system, the resulting variations of Me, Reθ, and δ as a function of plate incidence, 

α, for several freestream Reynolds numbers for several locations along the plate are 

presented in Figures 4 through 6. To design the test environment, including test conditions and cavity geometry, the model angle of attack was set to yield the prescribed Me, 

which was approximately constant along the plate at a given a setting. Next, the position 

on the plate yielding the required Reθ was determined from the boundary layer solution. 

Plate boundary layer thickness (δ) at this position was then used as a multiplier on the 

flight scaling of (H/δ) to determine the required depth of the cavity, HTunnel, to be installed 

on the wind tunnel model. The model-scaled cavity length and width follow by multiplying (W/H)Flight and (L/H)Flight by HTunnel, respectively. It is important to note that each resulting cavity represents a point design for the selected conditions because the plate 

boundary layer changes as the tunnel test conditions change. As seen in Table 1, the 

parameter variation is limited; therefore, a family of models was designed for Test 6857 

to allow a limited parametric variation about the prescribed damage geometry. All of the 

prescribed cases are for open cavity flow conditions, so additional models were used 

during Test 6868 to address the heating issues related to closed cavity flows. 

Experimental Methods 

This section presents the experimental methods and apparatus, beginning with a 

description and operational characteristics of the wind tunnel facility. This is followed 

with a detailed discussion of the baseplate model, the fabrication process used for the 

50 different cavity model inserts, a model geometry description, and photographic documentation. Next, an overview is given of the fluorescent oil flow technique. This qualitative measurement method combined with heating measurements on the baseplate is 

used to qualify the testing environment. Finally, the phosphor thermography method 

used to provide global heating measurements of the cavity flow field is presented. 

 
 

17

Test Facility 

The 20-Inch Mach 6 Air Tunnel (Fig. 7) is a blowdown facility in which heated, 

dried and filtered air is used as the test gas. It has a two-dimensional, contoured nozzle, 

which opens into a 20.5-in. by 20-in. test section, and it is equipped with a bottommounted injection system that can transfer a model from the sheltered model box to the 

tunnel centerline in less than 0.5 seconds. Run times of up to 15 minutes are possible in 

this facility, although for the current aeroheating study run times of only a few seconds 

were required. The nominal reservoir conditions of this facility are stagnation pressures 

of 30 psi to 500 psi with stagnation temperatures of 760 ̊R to 1000 ̊R, which very nearly 

produce perfect gas (γ = 1.4) freestream flows with Mach numbers between 5.8 and 6.1 

and Reynolds numbers of 0.5x106

/ft to 7.3x106/ft. The nominal flow conditions in this 

facility can be seen in Table 2. Micol (1995) presented a more detailed description of 

this facility. 

Test Model Description 

The baseline model used for this study (Figs. 8 and 9) was a strut-mounted, rectangular 10-inch wide by 28-inch long flat plate. The model is also shown in Figure 7 below the tunnel in the model injection box. Two different nose geometries were available, 

including a sharp leading edge and a blunted leading edge with a radius of 0.125 inches. Model station 0 was measured in all cases from the origin of the sharp nose configuration. The blunt nose model was used in all cases because it resulted in edge Mach 

numbers of approximately 2.9 at zero incidence as required to match the test conditions. 

Test article surface conditions on the model centerline were determined using the 

LAURA code. Test articles were inserted into a 4-inch wide by 18-inch long test bay beginning at station 7.5 inches downstream of the nose. During Test 6857 (T6857), the 

first entry, three 4-in. by 6-in. solid surface test articles were used to create the baseline 

surface (see Fig. 8), following the original design specifications. Installation, alignment, 

steps, and gap sealing issues were resolved (though not quite as well as desired) with 

0.003-inch thick kapton tape as shown in Figure 8. The test process was greatly improved during Test 6868 (T6868) where a single-piece model insert of dimension 4-in. 

by 18-in. was used (see Fig. 9). The model was equipped with surface pressure orifices; 

however, these were sealed from the backside to prevent leakage and disturbance tripping of the boundary layer. Additionally, there exists a provision for installation of 

boundary layer trips near the leading edge. This can be seen as a rectangular insert 

paralleling the leading edge in Figure 9 

Fabrication of the ceramic cavity models was a significant task, initially requiring 

a total of 50 different geometries to be developed. However, since model testing was 

performed concurrently with model manufacturing, many of the early designs used in 

T6857 were deemed unnecessary and were not built, though multiple copies of several 
 18

configurations were developed to address specific testing and fluid dynamic effects issues. Buck (2005) presented a detailed description of the cavity-model fabrication process. The first step in the process was developing the model geometry definition as required by the test conditions (recall that each model is a point design). Coordinates for 

these models were electronically transferred to the rapid-prototyping facility where an 

appropriate CAD model description was developed. Stereolithographic Array process 

(SLA) resin models (Fig. 10a) were then fabricated as patterns used for the casting of 

the silica ceramic insert models. An aluminum backing plate (Fig. 10b) was bonded to 

backside of each ceramic insert for mounting in the flat plate model. Models used to perform the oil flow investigations were the SLA patterns painted with a black hightemperature paint to enhance photographic contrast (Fig. 10c). This figure shows geometry details of the backside of the model without the aluminum backing plate. The ceramic models were spray coated with a mixture of phosphors suspended in a silicabased colloidal binder (Fig. 10d). Fiducial marks were next applied to the model surface 

at critical reference locations, such as the corners of the cavity on both the floor and the 

surrounding edges, and on the plate surface around the cavity. They can be seen as 

small, circular marks and they are used to aid in data reduction and model orientation. 

Erroneous heating measurements are realized at a fiducial mark because the inking 

process changes the calibration of the phosphor. Therefore, results at these locations 

should be excluded from all measurements. The locations of these marks can be seen 

in Fig. 10d. The fiducial marks used in T6857 were laboriously applied by hand because 

of time constraints imposed by the wide-ranging demands of the accident investigation 

on the Quality Assurance Laboratory; accordingly, the locations are not as precise and 

cleanly defined as normally desired. Subsequently, in the T6868 investigations when 

more time was available the fiducial marks were applied by the normal method using the 

direct computer controlled (DCC) coordinate measuring machine (CMM). The geometry 

description of each cavity model built for T6857 is presented in Table 3 and photographs of the painted SLA model insert and a close-up photograph of the ceramic cavity 

are presented in Figure 11. Similarly, T6868 models are presented in Table 4 and Figure 12. 

Two special case models of Configuration 3 were built, namely 3-Gap and 3-

Gouge (see Fig. 11). Gap fillers are shim material placed between the shuttle tiles for 

thermal and vibration isolation. Loss of the gap fillers can allow pressure relief in the 

cavity, leading to a change in the heating. Configuration 3-Gap was designed to assess 

this change by cutting spanwise and longitudinal 0.0143-inch wide by 0.536-inch deep 

gaps on each side of the cavity on the downstream end. The scaled width G/H=0.0267 

is similar to that on the Orbiter. Configuration 3-Gouge was a fortuitous mistake during 

fabrication that occurred when the mold improperly released during casting, providing a 

more realistic impact-damage geometry. 
 

19

Oil Flow Visualization Technique 

A new in-situ fluorescent oil flow technique developed at Langley Research Center (Buck, Pulsonetti, and Weilmeunster, 2005) was used to qualify the model surface 

flow and examine several of the cavities prior to initiating the heating measurements. 

Before each run, the model surface was cleaned and two coats of oil were applied. The 

first coat was a heavy weight oil used to reduce the surface tension; the second was a 

spritzed coating of a light weight oil that caused oil beads to form. Finally, the green fluorescent powder was sprinkled over the oil. During the run, an ultraviolet bandpass filtered mercury arc lamp (365 nanometers wavelength) was used to excite the fluorescent powder, which emits a visible light pattern. The emitted brightness is a function of 

the powder density, therefore accumulations of the powder on oil beads and along separation lines and other such features are particularly visible. The oil flow was then recorded during the run using a digital video camera. Post-run digital pictures of the run 

were also recorded. 

Phosphor Thermography Technique 

Global surface heating distributions were calculated using the two-color relativeintensity, phosphor-thermography aeroheating measurement method (Merski 1999). 

This is the standard method for obtaining aeroheating data in NASA Langleyʼs hypersonic wind tunnels, and it can be used to identify the surface heating effects of complex 

three-dimensional flow phenomena, which are difficult to examine using conventional 

discrete-sensor methods. With this method, ceramic wind tunnel models are coated with 

phosphor crystals that fluoresce in the red and green regions of the visible light spectrum when illuminated by ultraviolet (UV) light. During a wind tunnel run, the phosphorcoated model is exposed to the heated flow of the tunnel, and the resulting changes in 

fluorescence intensity of the model are recorded and digitized through a 640 by 480 

resolution color CCD (charge coupled device) camera and a state-of-the-art video acquisition system. The fluorescence intensity is dependent on both the intensity of the 

incident UV light and the local model surface temperature. The UV dependence is removed by taking the ratio of the green to red intensity images, from which surface temperature distributions are determined through prior calibrations. Images are acquired 

before the wind tunnel run and after injection of the model to the tunnel centerline during 

a run. Data are acquired, reduced and analyzed using the IHEAT code (Merski 1999) 

and may be mapped onto a 3-dimensional CAD representation of the test article using 

the MAP3D code (Merski 1999). During the current experiments, phosphor Batch 9 was 

used for T6857, while an improved lower-temperature formulation denoted phosphor 

Batch 10 was used for T6868. Results are presented herein in terms of a nondimensional heat transfer coefficient ratio, h/hFR, where hFR is the theoretical heating 

computed with the Fay-Riddell (1958) method for a 0.025-in. radius sphere, which is the 

radius of the nose of the test models, at 540 ̊R 
 20

Experimental Results 

This section includes a discussion of the measurements acquired for both T6857 

and T6868. Only image data will be presented; hence, only qualitative statements are 

made concerning their behavior. All required digital data have been transferred electronically to the Entry Aerothermodynamics Damage Assessment Team at Johnson Space 

Center. First presented are observations from oil flow visualizations used to assess the 

surface flow over the base model and around select cavity configurations. Next, global 

surface heating distributions acquired using phosphor thermography are introduced. 

Base plate heating combined with the oil flow visualizations allows general statements 

regarding the state of the cavity entry conditions. Future analysis will allow assessment 

of the exit conditions. Heating measurements acquired during geometry screening studies of various parameters when combined with the oil flow visualizations allows generalized statements regarding the importance of the parameter variation prior to establishing 

an extensive testing program. 

Oil Flow Visualization 

Test conditions for the oil flow visualizations conducted in T6857 are presented in 

Table 5 and the corresponding post-run images are presented in Figure 13. In all cases, 

flow is from left to right. The base plate surface flow is characterized in Runs 13-18 for a 

series of angles of attack and stream Reynolds numbers covering the anticipated test 

envelope. The side vortices caused by finite-plate width effects that are emanating from 

the plate leading edge are particularly evident in runs 17 and 18. The bounding track of 

these disturbances indicates that flow between is predominantly two-dimensional down 

the plate length. Run 17 is particularly revealing in that the Kapton tape sealing upstream pressure orifices has lifted at its corners, generating two vortices that track 

straight downstream through the test region with little divergence until the aft end of the 

model, as desired. This run and, also, the nonuniform flow shown in Run 16 highlight 

the required care that must be exercised with this model in the tape-sealing process at 

each joint to ensure clean undisturbed flow over the test region. 

Oil flow visualizations for several of the cavity geometries are presented in Figure 

13 for Runs 19-27. Run 19 is a deep cavity of L/H=8 and H/δ=1.76, while Run 23 is a 

medium-depth cavity of L/H=7.2 and H/δ=1.1. In both of these cases, the movement of 

the oil on the floor is almost non-existent, indicating open cavity flow. Downstream of 

each cavity endwall, the flow wipes the oil longitudinally away from the cavity with little 

lateral disturbance. The near-field impact of flow misalignment with the major axis of an 

open cavity is shown by comparing Runs 23, 26, and 24 for rotation angles of 0, 10, and 

90 degrees, respectively. The cavity geometry in each case is L/H=7.2, W/H=2.4, and 
 

21

H/δ=1.1 as measured along the major axis. Here, the flow approaches the cavity uniformly and leaves the cavity with little disturbance to the local surrounding flow field. 

Again the floor reveals no motion in the oil. It should also be highlighted again that as 

the cavity is rotated, its width changes with little local impact on the surrounding flow 

field. All of these oil flow visualizations support previously published observations that 

open cavity flows skip the cavity. 

Length effects for closed flow cavities can be examined using Runs 20 and 27 for 

L/H=14 and 17, respectively. Each of these cavities has W/H=2.4 and H/δ=1.1. According to published information the lower bound for closed flow is L/H≈13-14, therefore only 

minimal movement of the oil would be expected for Run 20 and more disturbance 

should occur with increasing length, as it does in Run 27. Though not as apparent in the 

reduced resolution of the image, the originals do indicate increased motion on the cavity 

floor, particularly near the end wall as the flow turns to exit the cavity. Movies of the developing oil motion acquired during the test even reveal upstream movement of the oil in 

the cavity corners. Compared to the open cavity cases, the closed geometries have 

augmented wiping and spreading of the oil downstream as the flow exits the cavity near 

field, indicating the occurrence of stronger interactions. A downstream view of the disturbance generated by the L/H=17 cavity of Run 27 is also shown, indicating two, strong 

persistent vortices tracking to the end of the plate, implying concern for possible heating 

augmentation due to boundary layer transition. These observations are, again, in accordance with published information concerning longer, closed cavities. 

Some closed flow cavity width effects are shown using images for Runs 22, 20, 

and 21 for L/H=14 cavities of depth H/δ=1.1. These cavities have widths, W/H, of 3.6, 

2.4, and 1.2, respectively, i.e. a range of ±50 percent about the baseline width. The images are consistently scaled, notated, and presented in Figure 14 for discussion. Surface flow external to the cavities is remarkably similar in all cases. First, the oil streaks 

above (and below) each cavity bend toward the cavity, reflecting the reduced pressure 

generated in the cavity as flow expands into it. As indicated by the solid red line marked 

“A”, these streaks reach a minimum point at approximately the same location. Next, the 

external flow then diverges and spreads at approximately the same angle as indicated 

by the dashed blue line marked “B”. Internal to the cavity, little movement of the oil is 

apparent upstream of dotted teal line marked “C”, which is a flow separation line between the main flow in the central region of the cavity and the endwall forward-facingstep flow field. Downstream of “C” the flow is a recirculation away from the endwall toward “C”. The near-field effects of the cavity appear insensitive to width, probably until 

the two sidewalls are of such proximity that their independent viscous effects can interact – then, they become a longitudinally oriented gap. 
 22

Surface Heating Data 

Figures 15-18 present the heating data acquired for T6857, the open cavity experiment. The geometry for the models in this test series are presented in Table 3 and 

photographically in Figure 11, and the test matrix for this test series are presented in 

Tables 6-10. Figures 19-30 present the heating data for T6868, the closed cavity experiment. The model geometries for this test series are presented in Table 4 and photographically in Figure 12, and the test matrix for this test series is presented in Tables 

11-16, where the runs are grouped according to objective. In each case, the heating images are unmapped, i.e. they are the two-dimensional view obtained during each run. 

The narrowing (perspective) and skew of the heating image is a result of the camera 

angle with respect to the plate. For T6857, the three different 4-inch wide by 6-inch long 

inserts are readily discerned. Even though the overall length is 18 inches, the phosphor 

fluorescence is undetectable through the Kapton tape used to seal the joints and gaps 

on the model upstream, downstream and between each insert. The phosphor-imaging 

camera was typically focused to capture the entire 4-inch by 18-inch ceramic insert region; however, for select runs the camera was focused on the immediate cavity region. 

These runs are indicated in the Test-Run column of figure by the addition of an “f” to the 

value (ex. 6857-096 f indicates Test 6857 Run 096 focused). For select “screening” 

runs, the SLA models were painted with an insulating coating and sprayed with the 

phosphor. Therefore, these runs offer only a qualitative thermal foot print (analogous to 

an oil flow footprint), because the thermal properties of the model are unavailable for 

determining the heating. These runs are denoted in the Config column by appending an 

“s” the configuration identifier (ex. 9 s indicates Configuration 9 SLA). For this section, 

only qualitative remarks are presented regarding the image data. 

For the T6857 baseline plate cases presented in Figure 15, fairly uniform heating 

is observed, even though very small levels are realized because the thermal driving potential for grazing flows on a plate are small for this facility. Accordingly, larger uncertainties and scatter will be evident in all of the data acquired in lower temperature regions of the flow. 

The open (short) cavity heating images for T6857 are presented in Figure 16, 

and they are ordered first according to configuration number (note the color coding) and 

then according to test condition. Recalling from Table 1, the target depth H/δ range of 

0.72 to 1.76 and the target Reθ range of 136 to 241, it is obvious that none of the Configuration 3 or 5 cases develop any appreciable rise in the end wall or floor heating until 

considerably beyond expected limits on the vehicle, as is evident in Run 49, where the 

H/δ=3.37 and Reθ=461. In all cases, the heating drops on the cavity floor, as anticipated 

from the literature and flow physics discussions. The effect of increasing the cavity 

depth, H/δ, at matched Reθ is to increase the heat load on the cavity floor. For example, 

compare Configuration 5 Runs 60, 61, and 62 with Configuration 3 Runs 47, 48, and 49, 

respectively. Several special cases of Configuration 3 were developed: namely, 3-
 

23

Gouge, 3-Gap, and 3-Slot. The Gouge configuration offers a “more damage-like” effect 

of the ragged bottom and sides appears to reduce the average depth of the cavity, 

thereby reducing the heating (see Runs 42, 43, 44, and 45) compared to the standard 

Configuration 3. The Gap configuration examines the impact of removing the gap fillers 

from the end of the cavity, allowing some flow relief inside the cavity. These runs are 51, 

52, 57, 58, and 59. Test Runs 93 through 96 were focused views of the gap region, allowing enhanced viewing of the cavity. Finally, an attempt was made to examine the 

floor heating when a slot exists along the bottom side of the cavity, as would exist behind a carrier panel. Here, a slot was cut along the bottom side of the cavity and incavity flow was directed into a plenum attached to the back of the model. The plenum 

was evacuated into a large dump tank maintained at near vacuum. This test series includes runs 102 through 106. 

Oil flow visualizations presented in Figures 13-14 indicated that external flow 

skipped over open cavities with little effect, whether aligned with or skewed from the 

main flow. This same no-significant-effect observation can be applied to the heating of 

open cavities in Figure 17 where Configuration 8 at 0O rotation (i.e. aligned with the 

stream) is compared with Configuration 11 at 10O rotation for a large range of depths 

(1.1 to 2.11) and Reθ (205 to 394). 

A series of screening experiments was conducted during T6857 to preview the 

effects of closed (long) cavities. Unlike the skimming flow over open cavities that displayed only minor heating augmentation with changes near the target conditions, closed 

cavities are long enough to allow sufficient turning for flow entry. The turning into the 

cavity, along the floor, and out of the cavity creates substantial high shear regions that 

generate increases in the heating. Corners offer the possible development of vortices 

that, again, create high shear regions. These effects are presented in Figure 18 for 

L/H=14 and 17 cavities. Typically, stronger interactions occur with longer cavities, 

thereby enhancing the heating. Figure 18 also offers a width-effect series for heating of 

L/H=14 at matched entry conditions. As with the oil flow visualizations, these were conducted at W/H values of 1.2, 2.4, and 3.6. An example of this variation at matched conditions can be observed in Runs 91f, 81f, and 82f, respectively. As the cavity is widened, the major effect appears to be a spreading of the vortices exiting the cavity region 

(e.g. compare Runs 81f with 82f). The largest variations appear for the narrowest cavity 

(e.g. W/H=1.2, Run 91f) where the sidewalls may be close enough for the occurrence of 

side-to-side vortical interactions. 

The baseline heating runs for T6868 are presented in Figure 19 over the range of 

expected test conditions. As with T6857, fairly uniform spanwise heating is observed 

and a uniform drop in the heating occurs, as would be expected of laminar flow conditions. A significant exception occurs with Run 123 acquired at conditions where transition to turbulence would not be unexpected on the aft end of the plate, and this indeed 

appears to be the case. Two baseline cases, Runs 30 and 31, were acquired at Re= 

0.5x106

/ft. These two runs display a high degree of scatter in the heating and extremely 
 24

low heating for Run 31, again highlighting the low thermal driving potential of the facility. 

Accordingly, larger uncertainties and scatter will be evident in all of the data acquired in 

lower temperature regions of the flow. 

Figure 20 presents the results from a limited number of open cavity runs that 

were made to check the consistency of the test process and to verify consistency of observations between T6857 and T6868. Note that Configuration 5 is a standard ceramic 

model and Configuration 9s is an SLA insert. 

Cavity length effects are presented as shown in Figures 21 through 24 for lengths 

L/H=14 to 30. 

H/δ L/H 

Range 

Reθ Me

Figure 21 1.1 14-30 300 2.90 

Figure 22 2.4 14-25 300 2.90 

Figure 23 1.1 14-25 500 2.90 

Figure 24 2.4 14-20 500 2.90 

Generally, longer cavities produce greater effects on the cavity floor and on the downstream. This is in direct response to the increase of high-energy fluid entering the cavity. 

The longer cavities enable the development of stronger vortices on the top outside edges as the flow rolls into the lower-pressure regions, and development of corner vortices 

in the junction between the sidewall and floor. These stronger interactions increase the 

shearing with the cavity structure providing increased resistance to maintaining laminar 

flow, and transitional or turbulent flow may occur in the cavity. 

Deeper, long cavities greatly enhance the heating and the possibility of transition 

to turbulence. These effects are presented in Figures 25 through 28 as follows: 

H/δ L/H 

Range 

Reθ Me

Figure 25 2.39 14-30 500 2.25 

Figure 26 5.20 14-30 500 2.25 

Figure 27 0.65 14-25 260 2.25 

Figure 28 2.70 14-30 260 2.25 

The effects are strongly demonstrated by comparing Figures 25 and 26 where the 

depth, H/δ, increases from 2.4 to 5.2 and then in Figures 28 and 27 where depth decreases from a deep 2.70 to a shallow 0.65. In this latter case, it becomes difficult to see 

the cavity foot print in the heating image because of the lower-disturbance shallow cavity. 
 

25

The effect of edge Mach number, Me, can be evaluated by comparing Figures 24 

and 25. These Figures have the following conditions: 

H/δ L/H 

Range 

Reθ Me

Figure 24 2.4 14-20 500 2.90 

Figure 25 2.4 14-30 500 2.25 

Care should be exercised to ensure that the transitional-turbulent flow has not occurred 

and that a proper comparison is being conducted. 

Pressure gradient effects are examined in Figure 29. The derivation of the surface geometry is presented in the Appendix. Since these are SLA models coated, first, 

with an insulating paint and, then, with the phosphor, only the qualitative observations 

can be provided regarding their behavior, since the thermal properties of the SLA resin 

are unknown. The data are presented with caveats. First, all cavity-entry conditions 

were determined using the dcp/dx=0 results from the laminar LAURA CFD solutions — 

no CFD solutions were developed for the gradient surface. Though boundary layer edge 

conditions at the beginning of the insert plate (station 7.5) are the same for each model 

test set and though each model has a prescribed design pressure gradient, the actual 

conditions at the entry of the cavity will differ somewhat for the different gradient conditions. For example, the expansion surface will increase Mach number above the baseline while the compression surface will decrease Mach number. Without substantiation, 

this effect is believed minimal because the cavities are located upstream on the insert at 

station 10.5 before large changes in edge conditions occur. Next, spanwise gradients 

are also believed to be minimal at the cavity because the presence/absence of sidewall 

constraints at the insert edge is small (recall that the insert is only 4 inches wide); 

downstream the effect may be more pronounced. With these qualifiers, scaled dimensions (L/H, W/H, H/δ) of (20, 2.4, 1.1) are assigned to the first set of target test conditions of Me=2.895 and Reθ=300 (α=0o

). Entry and exit boundary layer states for this 

case are laminar based on analysis of the heating levels. At the second set of conditions, Me=2.245 and Reθ=503 (α=-10o

), the cavity depth is increased to 2.4. The entry 

boundary layer is laminar for this condition; however, the exit boundary layer is turbulent, indicating shear layer transition within the cavity. For the first set of test conditions, 

gradient effects are not obvious in the vicinity of the cavity with the possible exception of 

changes in vortex strength downstream. However, for the second set of test conditions, 

the gradient effects are definitely evident, particularly downstream for the negative gradient with dcp/dx=-.005 where a severe narrowing of the cavity wake occurs and where 

significant heating augmentation appears. As the gradient increases, the cavity wake 

spreads and the separation line between the wake and the surrounding “undisturbed” 

flow straightens. Also, heating augmentation appears to decrease near the cavity for increasing dcp/dx. Unfortunately, the available data are insufficient to independently assess the influences of the longitudinal and lateral gradients. 
 26

The impact of flow misalignment with the cavity major axis is examined in Figures 

30 and 31 for L/H=20 cavities. Since these are SLA models, only the impact of the rotation angle on the flow field and qualitative levels can be evaluated. Four angles are 

available 0O (the baseline case), 15O, 30O, and 45O. The conditions of these tests are: 

H/δ L/H Reθ Me

Figure 30 1.1 20 300 2.90 

Figure 31 2.4 20 500 2.25 

As is evident in the figures, a skewed cavity can have a significant impact on the heating levels and on the downstream transition process. A skewed/rotated cavity allows an 

increase in the high-energy mass flowing into the cavity, and it develops a non-axial velocity component that enhances mixing and shearing. A closed-flow, rectangular cavity 

will increase in length as it is rotated until the cavity diagonal is aligned with the stream, 

after which it will decrease and the flow will ultimately become that of an open-flow, rectangular cavity with significantly reduced heating, if the W/H<10. 

Phosphor Uncertainty Analysis 

Merski published the uncertainties for the phosphor thermography method as a 

whole in 1999. Uncertainties in that report where determined under conditions where 

much larger temperature increases were experienced. For the present test, conditions 

on the grazing-flow models presented here generate a much smaller temperature 

change from the pre-run value, implying the existence of larger uncertainties. Typically, 

the analysis metric of choice has been the heating augmentation above/below the undisturbed ambient condition for both the experimental data and computational simulations. This metric is known as Bump Factor (BF) and it is the uncertainty in this parameter that is actually desired. BF is defined as a local normalization of h/href by a reference 

average heating value obtained upstream of the cavity, havg/href. Thus, 

BF = (h/href)/(havg/href) = (h/havg)

When presented and analyzed in this format the disturbance field introduced by the cavity is readily discernable, because the undisturbed state has a nominal value of unity. 

BF uncertainty levels were presented in Everhart, et al. (2006) by adding phosphor 

heat-transfer bump-factor uncertainty algorithms to the MAP3D code, resulting in global 

uncertainty surface maps similar to the three-dimensional heat transfer maps. The process from that publication is outlined below. 

The solution of the heat conduction equation (i.e. the data reduction equation) 

used in the IHEAT code is given by 

Θ

haw (Tw / hw ) − Tinit

= 1− e

Λ2

erfcΛ
 

27

where 

and 

Λ = h(haw / Tw ) t

β

Also, h is the convective heat transfer coefficient, haw is the adiabatic wall enthalpy, hw is 

the wall enthalpy, Tw is the corresponding wall temperature, and t is the effective time of 

data acquisition. β is the thermal product of the substrate materials and it is the square 

root of the product of the material density, specific heat, and thermal conductivity. Bias 

and precision uncertainty values developed at the 95 percent confidence level for each 

parameter appearing in the data reduction equation are given in Table 12. 

Each of these individual parameter uncertainties was individually inserted into the 

data reduction equation to determine a component uncertainty in heat transfer coefficient. The bias uncertainty, B, and the precision uncertainty, P, for the heat transfer coefficient were determined by obtaining the root-sum-square (RSS) from each of the 

component uncertainties using 

B = Bj

2

( ) ∑

1

2

 and 

P = Pj

2

( ) ∑

1

2

where j is the number of component uncertainties. Total uncertainties are obtained by 

taking the RSS of the bias and precision uncertainties. Bias and total uncertainties of 

the heat transfer coefficients are determined at every pixel point imaged on the model. 

Bump factors extracted from the mappings were used in two ways: 1) as line cuts 

when selecting specific data, and 2) as regions of interest over which all of the data 

were averaged. In the first case, the total uncertainties are applicable. In the latter case, 

bias uncertainties are applicable because precision uncertainties are removed during 

averaging, since they are primarily due to random pixel scatter. Accordingly, the average heat transfer coefficient havg used to compute bump factors was assumed to have 

only a bias uncertainty. Typically, total uncertainties are determined using the quotient 

rule by adding the total uncertainty in hlocal to the bias uncertainty in havg. For nominally 

flat models of this type, very low heating rates are obtained during a run and the resulting uncertainties were initially very high. Yet, comparisons of line cuts were more consistent than the uncertainty analysis seemed to suggest. While conservatism is important, excessive conservatism is undesirable. Therefore, the phosphor temperature 

lookup table data were re-examined and it was determined that since there was minimal 

variation of incident UV intensities on the models during the tunnel runs, and because of 

the low heating measurements, the data were typically confined to one very small segment of the temperature range. Thus, the bias uncertainties for the data used in calculating hlocal and havg were uni-directional and determinate. Therefore, the bias uncertainties in hlocal and the bias uncertainties in havg had to be subtracted via the quotient rule, 

Θ( ) xyt , , Txyt ( ) , , – Txy ( ) , , 0 Txyt ( ) , , Tinit = = –
 28

instead of added, when computing bias uncertainties for the augmentation factors. Similarly, total uncertainties in augmentation factors were then obtained simply by taking the 

RSS of the bias and precision uncertainties in hlocal and by subtracting the bias uncertainties in havg (since the precision uncertainties were negligible). 

Following this procedure, representative within-run total and bias uncertainties for 

open cavity flows are typically less than 10% over the plate surface. In the cavity where 

the lowest temperatures are experienced, the bias uncertainties are about 10%, while 

the total uncertainties approach 20%. For the closed conditions, the temperature rise is 

generally greater, resulting in lower with-in run bias uncertainties (3-4 percent or lower) 

with the exception of the low temperature regions on the cavity floor. Total within-run 

uncertainties are in the 10-20 percent range. An assessment of the within-test uncertainties, including run-to-run, model-to-model, etc. can be found in the “The Cavity Heating Tool” documentation OEAN-0305-001 by Anderson, et al. (2005). 

Summary 

Two cavity heating experiments have been conducted in support of the Columbia 

Accident Investigation Board in the NASA Langley 20-Inch Mach 6 Tunnel at laminarentry fluid-dynamic boundary-layer edge conditions typical of the Columbia Orbiter vehicle at Mach 18. The first experiment concentrated on open (short) cavity geometries, 

while the second focused on closed (long) cavity geometries. Rectangular planform geometries with flat bottoms and nearly vertical side walls and end walls were used in the 

modeling, representing what is believed to be a worst case, missing tile(s), nonbreeching, damage configuration. Global surface heating distributions were acquired for 

each test configuration via the two-color thermographic phosphor method and oil flow 

visualizations were used to assess the surface flow characteristics of select configurations. The present report summarizes the philosophy developed to scale from flight-totunnel and testing processes used to acquire the data. Results are presented in image 

format and were provided during the test to the Aeroheating Damage Assessment Team 

for further analysis. 

The traditional definitions for cavity flow regimes are L/H<10 for open cavities and 

L/H>13-14 for closed cavities. These definitions appear consistent with the present observations in hypersonic flow, though insufficient data are available to assess the validity of the regime limits. The present cavities of L/H=7.2 and 8 with depths ranging from 

H/δ=0.92 to 3.37 exhibit characteristics prescribed to open flows where the mainstream 

flow skips over the top of the cavity. Both oil flow and heating observations of these geometries indicate little disturbance to the surrounding environment with only a heating 

peak and recovery zone on the downstream endwall top edge. This is further supported 

by the observation that rotation of the cavity off the stream axis has little impact on sur-
 

29

rounding surface flow. Virtually no oil-flow motion was observed on the cavity floor, implying a low shear region and accordingly a drop in the surface heating. Since the flow 

skips over the cavity, the cavity floor is isolated from the high-energy external stream. 

Closed cavities of lengths L/H=14 to 30 were tested over a depth range H/δ=0.65 

to 5.2. Global phosphor thermography of closed cavities has revealed highly threedimensional surface heating details and vortical interactions in both the near-field and 

far-field that would have been impossible to observe using discrete-sensor measurement techniques. These cavities are highly vortical; they have large regions of high 

shear and significant energy may be transferred not only to the cavity floor from the external stream, but also to the endwall and the surrounding side and downstream surfaces. Longer cavities allow sidewall and floor vortex strengths to increase. These enhanced vortices increase the shearing and energy transfer to the cavity surfaces. The 

increased vortical mixing enhances the possibility of transition-to-turbulence within the 

cavity and downstream. Deeper closed cavities allow more backflow in the cavity in 

separation regions and allow greater flow-turning angles to be realized. The increased 

flow-turning promotes stronger shock and expansion waves, thereby augmenting the 

heating in these regions. When cavity width was increased the major local effect was to 

spread the vortices that develop on the sidewalls and their impingement points on the 

endwall. The major flow features in the cavity and on the endwall, and the resulting 

heating does not appear to be affected to first order with these width changes. Significant heating changes due to decreasing the width will probably appear only when the 

sidewall vortices are close enough to mutually interact. 

The effects of external pressure gradient on the cavity heating and cavity misalignment with the main stream were screened for closed cavities using phosphorcoated SLA models. These tests allow only a thermal footprint of the disturbance to be 

examined, much like an oil flow visualization. Longitudinal and lateral pressure gradient 

effects were observed in the images, but because of model constraints their individual 

effects can not be separated in the data. Rotating the closed cavity to create a misalignment with the stream introduces a lateral velocity in the cavity, increases the shearing and vorticity, and results in augmented heating and flow disturbance and increased 

the probability of transition-to-turbulence with and downstream of the cavity. 
 30

References 

Adams, J.C., Jr.: “Numerical Calculation of Hypersonic Laminar Cavity Flows,” AIAA 

Paper 74-707, ASME 74-HT-27, AIAA/ASME 1974 Thermophysics and Heat Transfer 

Conference, Boston, Mass., July 1974. 

Anderson, John D., Jr.: “Hypersonic and High Temperature Gas Dynamics,” McGrawHill Book Co., New York, 1989. 

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37

Tables 

Table 1.- Flight Scaled Cavities and Conditions. 

Location Me Reθ

δ

(inch) H/ δ W/H L/H 

Carrier Panel 6 3.0713 136.2 1.7643 1.70 1.33 8.0 

Carrier Panel 7 3.1107 132.9 1.7084 1.76 1.33 8.0 

Tile PT-T 3.1376 163.4 2.2742 1.10 2.40 7.2 

Tile PT-O 3.0591 172.6 2.3628 1.06 2.40 7.2 

Tile PT-M 3.0417 219.9 3.2046 0.78 2.40 7.2 

Tile PT-I 3.0122 241.3 3.4737 0.72 2.40 7.2 

Table 2.- Nominal flow conditions in the 20-Inch Mach 6 Air Tunnel. 

Re∞x10-6

(1/ft) 

M∞

pt1 

(psi)

Tt1 

( ̊R)

.56 5.93 30.63 856.33 

.84 5.91 46.20 863.70 

1.08 5.94 61.10 874.65 

1.39 5.95 81.33 886.94 

1.72 5.96 101.73 859.87 

2.09 5.98 126.52 903.14 

2.49 5.98 151.40 903.82 

2.98 5.99 181.96 904.78 

3.46 6.00 212.63 905.85 

4.12 6.00 252.61 905.93 

4.65 6.01 292.75 920.01 

5.41 6.02 343.57 924.47 

6.13 6.02 393.28 929.03 

6.98 6.03 450.89 933.62 

7.38 6.03 476.62 933.80 
Table 3.- Cavity models for Test 6857 - Columbia accident investigation. 

, 

del 

ch) X, cavity (inch) 

Me Reθ δ 

(inch) 

H/δ W/H L/H H 

(inch) 

W 

(inch) 

L 

(inch) 

Rotation 

(deg) 

dCp/dX 

(per ft) 

 0.0 0.0 0.0

4 6.858 2.895 241 0.305 1.8 1.3 8.0 0.536 0.714 4.285 0 0 

4 6.858 2.895 241 0.305 1.8 1.3 8.0 0.536 0.714 4.285 0 0 

4 6.858 2.895 241 0.305 1.8 1.3 8.0 0.536 0.714 4.285 0 0 

4 7.883 2.895 461 0.159 1.8 1.3 8.0 0.279 0.372 2.234 0 0 

6 2.050 2.895 205 0.240 1.1 2.4 7.2 0.264 0.633 1.900 0 0 

6 2.050 2.895 205 0.240 1.1 2.4 7.2 0.264 0.633 1.900 10 0 6 2.050 2.895 205 0.240 1.1 2.4 7.2 0.264 0.633 1.900 90 0 

7 1.153 2.895 205 0.240 1.1 1.2 14.0 0.264 0.317 3.694 0 0 7 1.153 2.895 205 0.240 1.1 2.4 14.0 0.264 0.633 3.694 0 0 

7 1.153 2.895 205 0.240 1.1 3.6 14.0 0.264 0.950 3.694 0 0 

3 0.757 2.895 205 0.240 1.1 2.4 17.0 0.264 0.633 4.485 0 0 
39

Table 4.- Cavity models for Test 6868 - Closed cavity geometry parameters. 

, 

del 

ch) X, cavity (inch) Me Reθ δ (inch) H/δ W/H L/H H (inch) W (inch) L (inch) 

Rotation 

(deg) 

dCp/dX 

(per ft) 

5 - - - - - - - - - 18.000 - 0 

.5 3.00 2.895 300 0.148 1.1 2.4 14.0 0.163 0.391 2.279 0 0 .5 3.00 2.895 300 0.148 1.1 2.4 17.0 0.163 0.391 2.768 0 0 .5 3.00 2.895 300 0.148 1.1 2.4 20.0 0.163 0.391 3.256 0 0 .5 3.00 2.895 300 0.148 1.1 2.4 25.0 0.163 0.391 4.070 0 0 .5 3.00 2.895 300 0.148 1.1 2.4 30.0 0.163 0.391 4.884 0 0 .5 3.00 2.895 300 0.148 1.1 2.4 20.0 0.163 0.391 3.256 15 0 

.5 3.00 2.895 300 0.148 1.1 2.4 20.0 0.163 0.391 3.256 30 0 

.5 3.00 2.895 300 0.148 1.1 2.4 20.0 0.163 0.391 3.256 45 0 .5 3.00 2.895 300 0.148 2.4 2.4 14.0 0.354 0.849 4.952 0 0 .5 3.00 2.895 300 0.148 2.4 2.4 17.0 0.354 0.849 6.013 0 0 .5 3.00 2.895 300 0.148 2.4 2.4 20.0 0.354 0.849 7.074 0 0 .5 3.00 2.895 300 0.148 2.4 2.4 25.0 0.354 0.849 8.843 0 0 .5 3.00 2.895 300 0.148 2.4 2.4 30.0 0.354 0.849 10.612 0 0 .0 10.50 2.895 500 0.139 1.1 2.4 14.0 0.153 0.367 2.143 0 0 .0 10.50 2.895 500 0.139 1.1 2.4 17.0 0.153 0.367 2.602 0 0 .0 10.50 2.895 500 0.139 1.1 2.4 20.0 0.153 0.367 3.061 0 0 .0 10.50 2.895 500 0.139 1.1 2.4 25.0 0.153 0.367 3.826 0 0 .5 3.00 2.895 300 0.148 1.1 2.4 20.0 0.163 0.391 3.256 0 -0.005.5 3.00 2.895 300 0.148 1.1 2.4 20.0 0.163 0.391 3.256 0 0.005.5 3.00 2.895 300 0.148 1.1 2.4 20.0 0.163 0.391 3.256 0 0.011.0 10.50 2.895 500 0.139 2.4 2.4 14.0 0.333 0.798 4.655 0 0 .0 10.50 2.895 500 0.139 2.4 2.4 17.0 0.333 0.798 5.653 0 0 .0 10.50 2.895 500 0.139 2.4 2.4 20.0 0.333 0.798 6.650 0 0 
Table 5.- Test matrix for Test 6857 - Oil flow runs. 

α 

(deg) 

Re∞x10-6

(ft-1) X, cavity (inch) 

Me Reθ δ 

(inch) 

H/δ W/H L/H Rot 

(deg) 

BL State Objective 

0 1 2.895 L Baseline 

-5 1 2.569 L Baseline -5 1 2.569 L Baseline 

3 1 3.100 L Baseline 0 4 2.895 L Baseline 0 4 2.895 L Baseline 

0 1 14.358 2.895 241 0.305 1.76 1.33 8.0 0 L Large H/

δ

1 8.653 2.830 205 0.240 1.10 2.40 14.0 0 L W/H, L/H 

0 1 8.653 2.895 205 0.240 1.10 1.20 14.0 0 L W/H, L/H 0 1 8.653 2.895 205 0.240 1.10 3.60 14.0 0 L W/H, L/H 

0 1 9.550 2.895 205 0.240 1.10 2.40 7.2 0 L Mid-Range H/

δ

0 1 9.550 2.895 205 0.240 1.10 2.40 7.2 90 L Rotation 0 1 9.550 2.895 205 0.240 1.10 2.40 7.2 10 L Rotation 

0 1 8.257 2.895 205 0.240 1.10 2.40 17.0 0 L W/H, L/H 
41

Table 6.- Test matrix for Test 6857 - Baseline model heating runs. 

α (deg) Re∞x10-6

(ft-1) X0, cavity (inch) Me Reθ δ (inch) H/δ W/H L/H Rot (deg) 

BL State Objective 

0 4 2.895 L Baseline 0 4 2.895 L Baseline 0 2 2.895 L Baseline 0 2 2.895 L Baseline 0 2 2.895 L Baseline 0 2 2.895 L Baseline 0 2 2.895 L Baseline 0 2 2.895 L Baseline 0 2 2.895 L Baseline 0 2 2.895 L Baseline 0 1 2.895 L Baseline 
Table 7.- Test matrix for Test 6857 - Cavity heating runs. 

g

α 

(deg) 

Re∞x10-6

(ft-1) X0, cavity (inch)

Me Reθ δ 

(inch) 

H/δ W/H L/H Rot 

(deg)

BL State Objective

uge 0 4 14.358 2.895 461 0.159 3.37 1.33 8.0 0 L Gouge uge 0 2 14.358 2.895 337 0.220 2.43 1.33 8.0 0 L Gouge 

uge 0 1 14.358 2.895 241 0.305 1.76 1.33 8.0 0 L Gouge uge 0 0.5 14.358 2.895 193 0.348 1.54 1.33 8.0 0 L Gouge 

0 0.5 14.358 2.895 193 0.348 1.54 1.33 8.0 0 L Large H/d 

0 1 14.358 2.895 241 0.305 1.76 1.33 8.0 0 L Large H/d 0 2 14.358 2.895 337 .220 2.43 1.33 8.0 0 L Large H/d 0 4 14.358 2.895 461 0.159 3.37 1.33 8.0 0 L Large H/d 
43

Table 7.- Test matrix for Test 6857 - Cavity heating runs (continued). 

α (deg) Re∞x10-6

(ft-1) X0, cavity (inch) Me Reθ δ (inch) H/δ W/H L/H Rot (deg) 

BL State Objective 

 0 1 14.358 2.895 241 0.305 1.76 1.33 8.0 0 L Filler Gaps  0 2 14.358 2.895 337 0.220 2.43 1.33 8.0 0 L Filler Gaps  0 4 14.358 2.895 461 0.159 3.37 1.33 8.0 0 L Filler Gaps  0 1 14.358 2.895 241 0.305 1.76 1.33 8.0 0 L Filler Gaps  0 2 14.358 2.895 337 0.220 2.43 1.33 8.0 0 L Filler Gaps 

 -5 2 14.358 2.569 337 0.220 2.43 1.33 8.0 0 L Filler Gaps 

 0 2 14.358 2.895 337 0.220 2.43 1.33 8.0 0 L Filler Gaps 

 0 4 14.358 2.895 461 0.159 3.37 1.33 8.0 0 L Filler Gaps 

 0 4 14.358 2.895 461 0.159 3.37 1.33 8.0 0 L Filler Gaps 

 -5 4 14.358 2.569 461 0.159 3.37 1.33 8.0 0 L Filler Gaps 

0 1 15.383 2.895 241 0.305 0.92 1.33 8.0 0 L Large H/d 0 2 15.383 2.895 337 0.220 1.27 1.33 8.0 0 L Large H/d 0 4 15.383 2.895 461 0.159 1.76 1.33 8.0 0 L Large H/d 

 0 0 14.358 0 0 0.159 3.37 1.33 8.0 0 L No Flow 

 0 1 14.358 2.895 241 0.305 1.76 1.33 8.0 0 L No Suction 

 0 1 14.358 2.895 241 0.305 1.76 1.33 8.0 0 L Suction 

 0 4 14.358 2.895 461 0.159 3.37 1.33 8.0 0 L No Suction 

 -5 4 14.358 2.569 461 0.159 3.37 1.33 8.0 0 L No Suction 

 -5 4 14.358 2.569 461 0.159 3.37 1.33 8.0 0 L Suction 

0 1 8.653 2.895 205 0.240 1.10 1.20 14.0 0 L W/H, L/H 0 4 8.653 2.895 394 0.125 2.11 1.20 14.0 0 L W/H, L/H 

-5 4 8.653 2.569 394 0.125 2.11 1.20 14.0 0 L W/H, L/H 
Table 7.- Test matrix for Test 6857 - Cavity heating runs (continued). 

α 

(deg) Re∞x10-6

(ft-1) X0, cavity (inch) 

Me Reθ δ 

(inch) 

H/δ W/H L/H Rot 

(deg) 

BL State Objective 

-3 4 8.653 2.690 394 0.125 2.11 1.20 14.0 0 L W/H, L/H -1 4 8.653 2.830 394 0.125 2.11 1.20 14.0 0 L W/H, L/H 

0 4 8.653 2.895 394 0.125 2.11 1.20 14.0 0 L W/H, L/H 

0 2 8.653 2.895 288 0.173 1.53 1.20 14.0 0 L W/H, L/H 

0 2 8.653 2.895 288 0.173 1.53 1.20 14.0 0 L W/H, L/H 

0 0.5 8.653 2.895 164 0.274 0.96 1.20 14.0 0 L W/H, L/H 

0 1 8.653 2.895 205 0.240 1.10 2.40 14.0 0 L W/H, L/H 0 2 8.653 2.895 288 0.173 1.53 2.40 14.0 0 L W/H, L/H 0 4 8.653 2.895 394 0.125 2.11 2.40 14.0 0 L W/H, L/H 

0 4 8.653 2.895 394 0.125 2.11 2.40 14.0 0 L W/H, L/H 

0 1 8.653 2.895 205 0.240 1.10 3.60 14.0 0 L W/H, L/H 

0 2 8.653 2.895 288 0.173 1.53 3.60 14.0 0 L W/H, L/H 0 4 8.653 2.895 394 0.125 2.11 3.60 14.0 0 L W/H, L/H 

0 4 8.653 2.895 394 0.125 2.11 3.60 14.0 0 L W/H, L/H 

-1 4 8.653 2.830 394 0.125 2.11 3.60 14.0 0 L W/H, L/H -1 4 8.653 2.830 394 0.125 2.11 3.60 14.0 0 L W/H, L/H 

1 4 8.653 2.960 394 0.125 2.11 3.60 14.0 0 L W/H, L/H 5 4 8.653 2.569 394 0.125 2.11 3.60 14.0 0 L W/H, L/H 

-3 4 8.653 2.690 394 0.125 2.11 3.60 14.0 0 L W/H, L/H 

0 1 8.653 2.895 205 0.240 1.10 2.40 17.0 0 L W/H, L/H 0 2 8.653 2.895 288 0.173 1.53 2.40 17.0 0 L W/H, L/H 
45

Table 7.- Test matrix for Test 6857 - Cavity heating runs (Concluded). 

α (deg) Re∞x10-6

(ft-1) X0, cavity (inch) Me Reθ δ (inch) H/δ W/H L/H Rot (deg) 

BL State Objective 

0 2 8.653 2.895 288 0.173 1.53 2.40 17.0 0 L W/H, L/H 

0 4 8.653 2.895 394 0.125 2.11 2.40 17.0 0 L W/H, L/H 

0 4 8.653 2.895 394 0.125 2.11 2.40 17.0 0 L W/H, L/H 

0 1 9.550 2.895 205 0.240 1.10 2.40 7.2 0 L Mid-Range H/d 0 2 9.550 2.895 288 0.173 1.53 2.40 7.2 0 L Mid-Range H/d 0 4 9.550 2.895 394 0.125 2.11 2.40 7.2 0 L Mid-Range H/d 

0 4 9.550 2.895 394 0.125 2.11 2.40 7.2 0 L Mid-Range H/d 

0 1 9.550 2.895 205 0.240 1.10 2.40 7.2 10 L Rotation 0 2 9.550 2.895 288 0.173 1.53 2.40 7.2 10 L Rotation 0 4 9.550 2.895 394 0.125 2.11 2.40 7.2 10 L Rotation 
Table 8.- Test matrix for Test 6868 – Baseline model heating runs. 

α 

(deg) Re∞x10-6

(ft-1) X0, cavity (inch) 

Me Reθ δ 

(inch) 

H/δ W/H L/H Rot 

(deg) 

BL State Objective 

-10 4 7.5 2.245 0 0 0 0 L Baseline 

0 2 7.5 2.895 0 0 0 0 L Baseline 0 4 7.5 2.895 0 0 0 0 L Baseline 0 1 7.5 2.895 0 0 0 0 L Baseline 

-10 1 7.5 2.245 0 0 0 0 L Baseline -10 2 7.5 2.245 0 0 0 0 L Baseline 

-10 0.5 7.5 2.245 0 0 0 0 L Baseline 

0 0.5 7.5 2.895 0 0 0 0 L Baseline 

0 8 7.5 2.895 0 0 0 0 L Baseline 
47

Table 9.- Test matrix for Test 6868 - Cavity heating runs. 

α (deg) Re∞x10-6

(ft-1) X0, cavity (inch) Me Reθ δ (inch) H/δ W/H L/H Rot (deg) 

BL State Objective 

0 2 14.88 2.895 346 0.178 1.6 1.3 8 0 L Short Cavity 

0 2 9.82 2.895 288 0.148 1.3 2.4 7.2 0 L Short Cavity 

-10 2 9.82 2.245 368 0.094 2.0 2.4 7.2 0 L Short Cavity -10 4 9.82 2.227 503 0.068 2.8 2.4 7.2 0 L Short Cavity 

0 2 10.5 2.895 300 0.148 1.1 2.4 14 0 L Cavity Length 

0 2 10.5 2.895 300 0.148 1.1 2.4 14 0 L Cavity Length 

0 2 10.5 2.895 300 0.148 1.1 2.4 17 0 L Cavity Length 0 2 10.5 2.895 300 0.148 1.1 2.4 20 0 L Cavity Length 0 2 10.5 2.895 300 0.148 1.1 2.4 20 0 L Cavity Length 

0 2 10.5 2.895 300 0.148 1.1 2.4 20 0 L Cavity Length 

0 2 10.5 2.895 300 0.148 1.1 2.4 20 0 L SLA vs. 

Ceramic 

0 2 10.5 2.895 300 0.148 1.1 2.4 25 0 L Cavity Length 0 2 10.5 2.895 300 0.148 1.1 2.4 30 0 L Cavity Length 

0 2 10.5 2.895 300 0.148 1.1 2.4 30 0 L Cavity Length 

0 2 10.5 2.895 300 0.148 2.4 2.4 14 0 L Cavity Length 0 2 10.5 2.895 300 0.148 2.4 2.4 17 0 L Cavity Length 

0 2 10.5 2.895 300 0.148 2.4 2.4 20 0 L Cavity Length 

0 2 10.5 2.895 300 0.148 2.4 2.4 20 0 L Cavity Length 

0 2 10.5 2.895 300 0.148 2.4 2.4 25 0 L Cavity Length 

0 2 10.5 2.895 300 0.148 2.4 2.4 30 0 L Cavity Length 

0 4 18 2.895 500 0.139 1.1 2.4 14 0 L Cavity Length 
Table 9.- Test matrix for Test 6868 - Cavity heating runs (Continued). 

α 

(deg) Re∞x10-6

(ft-1) X0, cavity (inch)

Me Reθ δ 

(inch) 

H/δ W/H L/H Rot 

(deg)

BL State Objective 

0 4 18 2.895 500 0.139 1.1 2.4 17 0 L Cavity Length 0 4 18 2.895 500 0.139 1.1 2.4 20 0 L Cavity Length 

0 4 18 2.895 500 0.139 1.1 2.4 20 0 L Cavity Length 

0 4 18 2.895 500 0.139 1.1 2.4 25 0 L Cavity Length 

0 4 18 2.895 500 0.139 2.4 2.4 14 0 L Cavity Length 0 4 18 2.895 500 0.139 2.4 2.4 17 0 L Cavity Length 0 4 18 2.895 500 0.139 2.4 2.4 20 0 L Cavity Length 

0 4 18 2.895 500 0.139 2.4 2.4 20 0 L Cavity Length 

-10 4 10.5 2.245 503 0.068 2.4 2.4 14 0 L Deep Cavity 

-10 4 10.5 2.245 503 0.068 2.4 2.4 14 0 L Deep Cavity 

-10 4 10.5 2.245 503 0.068 2.4 2.4 17 0 L Deep Cavity -10 4 10.5 2.245 503 0.068 2.4 2.4 20 0 L Deep Cavity -10 4 10.5 2.245 503 0.068 2.4 2.4 20 0 L Deep Cavity 

-10 4 10.5 2.245 503 0.068 2.4 2.4 20 0 L Deep Cavity 

-10 4 10.5 2.245 503 0.068 2.4 2.4 20 0 L Deep Cavity 

-10 4 10.5 2.245 503 0.068 2.4 2.4 25 0 L Deep Cavity -10 4 10.5 2.245 503 0.068 2.4 2.4 30 0 L Deep Cavity 

-10 4 10.5 2.245 503 0.068 2.4 2.4 30 0 L Deep Cavity 

-10 4 10.5 2.245 503 0.068 5.2 2.4 14 0 L Deep Cavity 

-10 4 10.5 2.245 503 0.068 5.2 2.4 17 0 L Deep Cavity -10 4 10.5 2.245 503 0.068 5.2 2.4 20 0 L Deep Cavity 

-10 4 10.5 2.245 503 0.068 5.2 2.4 20 0 L Deep Cavity 
49

Table 9.- Test matrix for Test 6868 - Cavity heating runs (Concluded).

g

α (deg) Re∞x10-6

(ft-1) X0, cavity (inch) Me Reθ δ (inch) H/δ W/H L/H Rot (deg) 

BL State Objective 

 -10 4 10.5 2.245 503 0.068 5.2 2.4 25 0 L Deep Cavity 

 -10 4 10.5 2.245 503 0.068 5.2 2.4 30 0 L Deep Cavity 

 -10 2.18 10.5 2.228 351 0.0883 4.0 2.4 14 0 L Deep Cavity 

 -10 2.11 10.5 2.244 346 0.0883 4.0 2.4 17 0 L Deep Cavity  -10 2.11 10.5 2.238 345 0.0897 3.9 2.4 20 0 L Deep Cavity 

 -10 2.11 10.5 2.238 345 0.0897 3.9 2.4 20 0 L Deep Cavity 

 -10 2.11 10.5 2.226 347 0.089 4.0 2.4 25 0 L Deep Cavity 

 -10 2.11 10.5 2.26 347 0.089 4.0 2.4 30 0 L Deep Cavity 

 -10 0.5 18 2.245 260 0.236 0.6 2.4 14 0 L Shallow Cavity  -10 0.5 18 2.245 260 0.236 0.6 2.4 17 0 L Shallow Cavity  -10 0.5 18 2.245 260 0.236 0.6 2.4 20 0 L Shallow Cavity 

 -10 0.5 18 2.245 260 0.236 0.6 2.4 25 0 L Shallow Cavity 
.- Test matrix for Test 6868 - Pressure gradient effects on heating using SLA models. 

Config. 

α 

(deg) Re∞x10-6

(ft-1) X0, cavity (inch) 

Me Reθ δ 

(inch) 

H/δ W/H L/H Rot 

(deg) 

BL State Objective dp/dx 

5s 0 2 10.5 2.895 300 0.148 1.1 2.4 20 0 L -0.005 6s 0 2 10.5 2.895 300 0.148 1.1 2.4 20 0 L 0.005 7s 0 2 10.5 2.895 300 0.148 1.1 2.4 20 0 L 0.011 

5s -10 4 10.5 2.245 503 0.068 2.4 2.4 20 0 L -0.005 6s -10 4 10.5 2.245 503 0.068 2.4 2.4 20 0 L 0.005 7s -10 4 10.5 2.245 503 0.068 2.4 2.4 20 0 L 0.011 
51

e 11.- Test matrix for Test 6868 - Cross-flow effect on heating using SLA models. 

Config. α (deg) Re∞x10-6

(ft-1) X0, cavity (inch) Me Reθ δ (inch) H/δ W/H L/H 

BL State Objective Rotation (deg) 

33s 0 2 10.5 2.895 300 0.148 1.1 2.4 20 L 15 33s 0 2 10.5 2.895 300 0.148 1.1 2.4 20 L 15 34s 0 2 10.5 2.895 300 0.148 1.1 2.4 20 L 30 34s 0 2 10.5 2.895 300 0.148 1.1 2.4 20 L 30 35s 0 2 10.5 2.895 300 0.148 1.1 2.4 20 L 45 35s 0 2 10.5 2.895 300 0.148 1.1 2.4 20 L 45 33s -10 4 10.5 2.245 503 0.068 2.4 2.4 20 L 15 33s -10 4 10.5 2.245 503 0.068 2.4 2.4 20 L 15 34s -10 4 10.5 2.245 503 0.068 2.4 2.4 20 L 30 34s -10 4 10.5 2.245 503 0.068 2.4 2.4 20 L 30 35s -10 4 10.5 2.245 503 0.068 2.4 2.4 20 L 45 35s -10 4 10.5 2.245 503 0.068 2.4 2.4 20 L 45 

Table 12.- Phosphor uncertainty estimates. 

Uncertainty Type Bias Uncertainty

Value 

Precision Uncertainty Value

Initial model wall temperature 1.43 oC 1.0 oC 

Model run wall temperature at reference

location 

1.93 oC n/a 

Model run wall temperature 3.36 oC 1.0 oC 

Effective time 0.02 s 0.05 s 

Thermal properties 5.9% n/a 
52 

Figures 

Figure 1.- Combined experimental and computational simulation approach. 

Figure 2.- Classification of supersonic/hypersonic cavity flows. 

 
 

53

Figure 3.- Sketch of model coordinate system. 

Figure 4.- LAURA predictions for Me versus α for baseline model. 

 
54 

Figure 5.- LAURA predictions for Reθ versus α for baseline model. 

Figure 6.- LAURA predictions for boundary layer thickness for baseline model. 

 
 

55

Figure 7.- Langley 20-Inch Mach 6 Air Tunnel. 

 

Figure 8.- Baseline model - 4-in. by 6-in. 

inserts used in Test 6857. 

Figure 9.- Baseline model - 4-in. by 

18-in. inserts used in Test 6868. 

 
56 

a) Cavity model patterns were fabricated 

using sterolithographic array process 

(SLA). 

b) Ceramic castings were fabricated using 

the SLA models and an aluminum mounting plate was bonded to the back side of 

the model 

c) SLA patterns were spray painted black 

for oil flow visualizations (typical backside 

shown without mounting plate). 

d) Ceramic castings were spray coated with 

thermographic phosphors and marked 

with fiducial indicators for wind tunnel 

testing. 

Figure 10.- Model fabrication process.
 

57

Config 

View 1 View 2 

1 

3 

3 

G 

o 

u 

g 

e 

3 

G 

a 

p 

Figure 11. – Cavity insert photographs for Test 6857.
58 Config 

View 1 View 2 

4 

5 

8 

9 No comparable model 

Figure 11. – Cavity insert photographs for Test 6857 (Continued).
 

59

Config 

View 1 View 2 

11

14

15

16

Figure 11. – Cavity insert photographs for Test 6857 (Continued).
60 Config 

View 1 View 2 

17

18

19

Figure 11. – Cavity insert photographs for Test 6857 (Concluded). 

 
 

61

Config 

Full View Close-up View 

27 

A 

27 

B 

28 

29 

Figure 12. – Cavity insert model photographs for Test 6868. 
62 Config 

Full View Close-up View 

30 

30 

B 

30 

G 

a 

p 

31 

Figure 12. – Cavity insert model photographs for Test 6868 (Continued).

 
 

63

Config 

Full View Close-up View 

32 

33 

34 

35 

Figure 12. – Cavity insert model photographs for Test 6868 (Continued).

 
64 Config 

Full View Close-up View 

36 

37 

38 Not available. 

39 

Figure 12. – Cavity insert model photographs for Test 6868 (Continued).

 
 

65

Config 

Full View Close-up View 

40 

41 

42 

43 

Figure 12. – Cavity insert model photographs for Test 6868 (Continued).

 
66 Config 

Full View Close-up View 

44 

45 

46 

47 

Figure 12. – Cavity insert model photographs for Test 6868 (Continued). 

 
 

67

Config 

Full View Close-up View 

48 

49 

50 

Figure 12. – Cavity insert model photographs for Test 6868 (Concluded). 

 
68 Test-Run

Test 6857 Post-Run Image 

Test-Run

Test 6857 Post-Run Image 

6857-0136857-015

6857-016 6857-017

6857-018 6857-019

Figure 13.- Oil flow images for Test 6857.
 

69

Test-Run

Test 6857 Post-Run Image 

Test-Run

Test 6857 Post-Run Image 

6857-020 6857-021

6857-022 6857-023

6857-024 6857-026

6857-027 

6857-027 — Far Field View

Figure 13.- Oil flow images for Test 6857 (Concluded). 

 
70 

T6857 

R22 

W/H=3.6

T6857 

R21 

W/H=1.2

T6857 

R20 

W/H=2.4

A B 

B 

B 

C 

Figure 14.- Oil flow visualization of closed-cavity width effect - L/H=14, H/δ=1.1. 

 
 

71

Test-Run 

Config. 

α 

(deg) 

Re

∞x10-6

(ft-1) Me Heating Image h/href

Color Scale 

6857-029 

1 0 4 2.895

6857-031 

1 0 2 2.895

6857-032 

1 0 2 2.895

6857-034 

1 0 2 2.895

6857-036 

1 0 2 2.895

6857-037 

1 0 2 2.895

6857-038 

1 0 1 2.895

Figure 15.- Baseline heating runs for Test 6857. 
72 Test-Run Config. 

α 

(deg) 

Re

∞x10-6

(ft-1) Me Heating Image h/href

Color Scale 

6857-033 

1 0 2 2.895

6857-035 

1 0 2 2.895

Figure 15.- Baseline heating runs for Test 6857 (Concluded). 

 
 

73

Test-Run 

Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image 

h/href

6857-046 

3 0 193 1.54 1.33 8 0

6857-047 

3 0 241 1.76 1.33 8 0

6857-048 

3 0 337 2.43 1.33 8 0

6857-049 

3 0 461 3.37 1.33 8 0

6857-042 

3 

Gouge 0 461 3.37 1.33 8 0

6857-043 

3 

Gouge 0 337 2.43 1.33 8 0

Figure 16.- Open flow cavity heating runs for Test 6857. 
74 

TestRun

Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image 

6857-044 

3 

Gouge 0 241 1.76 1.33 8 0

6857-045 

3 

Gouge 0 193 1.54 1.33 8 0

6857-050 

3 

Gap 0 241 1.76 1.33 8 0

6857-051 

3 

Gap 0 337 2.43 1.33 8 0

6857-052 

3 

Gap 0 461 3.37 1.33 8 0

6857-057 

3 

Gap 0 241 1.76 1.33 8 0

6857-058 

3 

Gap 0 337 2.43 1.33 8 0

6857-059 

3 

Gap 0 461 3.37 1.33 8 0

Figure 16.- Open flow cavity heating runs for Test 6857 (Continued).
 

75

TestRun

Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image 

6857-095 

f 

3 

Gap 5 337 2.43 1.33 8 0

6857-096 

f 

3 

Gap 0 337 2.43 1.33 8 0

6857-093 

f 

3 

Gap 0 461 3.37 1.33 8 0

Figure 16.- Open flow cavity heating runs for Test 6857 (Continued).
76 

TestRun

Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image 

6857-094 

f 

3 

Gap -5 461 3.37 1.33 8 0

6857-102 f 

No Suction 3 

Slot 0 241 1.76 1.33 8 0

6857-103 f 

Suction Applied

3 

Slot 0 241 1.76 1.33 8 0

Figure 16.- Open flow cavity heating runs for Test 6857 (Continued).
 

77

TestRun

Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image 

6857-106 f 

No Suction

3 

Slot 0 461 3.37 1.33 8 0

6857-105 f 

No Suction

3 

Slot -5 461 3.37 1.33 8 0

6857-104 f 

Suction Applied

3 

Slot -5 461 3.37 1.33 8 0

Figure 16.- Open flow cavity heating runs for Test 6857 (Concluded).

 
78 

TestRun

Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image 

6857-060 

5 0 241 0.92 1.33 8 0

6857-061 

5 0 337 1.27 1.33 8 0

6857-062 

5 0 461 1.76 1.33 8 0

Figure 16.- Open flow cavity heating runs for Test 6857 (Concluded). 
 

79

Test-Run 

Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

h/href

6857-075 

8 0 205 1.10 2.4 7.2 0

6857-076 

8 0 288 1.53 2.4 7.2 0

6857-077 

8 0 394 2.11 2.4 7.2 0

6857-078 

f 

8 0 394 2.11 2.4 7.2 0

Figure 17.- Cavity alignment effect on open flow cavity heating for Test 6857. 

 
80 Test-Run Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

6857-063 

11 0 205 1.10 2.40 7.2 10

6857-064 

11 0 288 1.53 2.4 7.2 10

6857-065 

11 0 394 2.11 2.4 7.2 10

Figure 17.- Cavity alignment effect on open flow cavity heating for Test 6857 (Concluded). 
 

81

Test-Run 

Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

h/href

6857-053 

16 0 205 1.10 1.2 14 0

6857-054 

16 0 394 2.11 1.2 14 0

6857-091 

f 

16 0 394 2.11 1.2 14 0

Figure 18.- Closed flow cavity heating runs for Test 6857.

 
82 Test- Run Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

6857-090 

f 

16 -1 1.2 14 0

6857-089 

f

16 -3 1.2 14 0

6857-088 

f 

16 -5 1.2 14 0

Figure 18.- Closed flow cavity heating runs for Test 6857 (Continued).
 

83

TestRun 

Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

6857-055 

16 0 288 1.53 1.2 14 0

6857-092 

f 

16 0 288 1.53 1.2 14 0

6857-056 

16 0 164 0.96 1.2 14 0

6857-066 

17 0 205 1.10 2.4 14 0

6857-067 

17 0 288 1.53 2.4 14 0

6857-068 

17 0 394 2.11 2.4 14 0

Figure 18.- Closed flow cavity heating runs for Test 6857 (Continued).
84 Test- Run Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

6857-081 

f 

17 0 394 2.11 2.4 14 0

6857-069 

18 0 205 1.10 3.6 14 0

6857-070 

18 0 288 1.53 3.6 14 0

6857-071 

18 0 394 2.11 3.6 14 0

6857-085 

f 

18 1 3.6 14 0

Figure 18.- Closed flow cavity heating runs for Test 6857 (Continued).
 

85

TestRun 

Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

6857-082 

f 

18 0 394 2.11 3.6 14 0

6857-083 

f 

18 -1 3.6 14 0

6857-084 

f

18 -1 3.6 14 0

Figure 18.- Closed flow cavity heating runs for Test 6857 (Continued).
86 Test- Run Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

6857-086 

f 

18 -3 3.6 14 0

6857-087 

f 

18 -5 3.6 14 0

6857-072 

19 0 205 1.10 2.4 17 0

6857-073

19 0 288 1.53 2.4 17 0

Figure 18.- Closed flow cavity heating runs for Test 6857 (Continued).

 
 

87

TestRun 

Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

6857-080 

f 

19 0 288 1.53 2.4 17 0

6857-074 

19 0 394 2.11 2.4 17 0

6857-079 

f 

19 0 394 2.11 2.4 17 0

Figure 18.- Closed flow cavity heating runs for Test 6857 (Concluded). 

 
88 Test-Run Config. 

α 

(deg) 

Re

∞x10-6

(ft-1) Me Heating Image h/href

Color Scale 

6868-019 27 0 2 2.90 6868-020 27 0 4 2.90 6868-021 27 -10 4 2.25 6868-022 27 -10 4 2.25 6868-023 27 0 2 2.90 6868-123 27 -10 8 2.25 6868-152 27 0 8 2.90 

Figure 19.- Baseline heating runs for Test 6868.
 

89

TestRun

Config. 

α 

(deg) 

Re

∞x10-

6 (ft-1) Me Heating Image h/href

Color Scale 

6868-024 27 0 4 2.90 6868-026 27 0 1 2.90 6868-027 27 -10 1 2.25 6868-028 27 -10 2 2.25 6868-030 27 -10 .5 2.25 6868-031 27 0 .5 2.90 

Figure 19.- Baseline heating runs for Test 6868 (Concluded). 

 
90 Test-Run Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image 

h/href

6868-017

5 0 346 1.57 1.3 8 0

6868-131

9 

s

0 288 1.1 2.4 7.2 0

6868-132

9 

s

-10 2.4 7.2 0

6868-133

9 

s

-10 2.4 7.2 0

Figure 20.- Open (short) cavity heating runs for Test 6868. 

 
 

91

TestRun 

Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

h/href

6868-049

28 0 300 1.1 2.4 14 0

6868-053

29 0 300 1.1 2.4 17 0

6868-071

30 0 300 1.1 2.4 20 0

6868-083

30 0 300 1.1 2.4 20 0

6868-103

30 0 300 1.1 2.4 20 0

6868-042

31 0 300 1.1 2.4 25 0

Figure 21.- Length effect on closed cavity heating for Test 6868 – Reθ=300, 

H/δ=1.1.
92 Test- Run Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image 

6868-066

32 0 300 1.1 2.4 30 0

6868-0140 

f

28 0 300 1.1 2.4 14 0

6868-0124 

f

30 0 300 1.1 2.4 20 0

6868-0128 

f

32 0 300 1.1 2.4 30 0

Figure 21.- Length effect on closed cavity heating for Test 6868 – Reθ=300, H/δ=1.1

(Concluded). 

 
 

93

Test-Run 

Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image

h/href

6868-077

36 0 300 2.39 2.4 14 0

6868-099

37 0 300 2.39 2.4 17 0

6868-108

38 0 300 2.39 2.4 20 0

6868-094

39 0 300 2.39 2.4 25 0

6868-0146 

f

38 0 300 2.39 2.4 20 0

6868-0115 

f

40 0 300 2.39 2.4 30 0

Figure 22.- Length effect on closed cavity heating for Test 6868 – Reθ=300, 

H/δ=2.39. 

 
94 Test-Run Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image

h/href

6868-058

41 0 500 1.1 2.4 14 0

6868-074

42 0 500 1.1 2.4 17 0

6868-063

43 0 500 1.1 2.4 20 0

6868-111

44 0 500 1.1 2.4 25 0

6868-151 

f

43 0 500 1.1 2.4 20 0

Figure 23.- Length effect on closed cavity heating for Test 6868 – Reθ=500, 

H/δ=1.1. 

 
 

95

Test-Run 

Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image

h/href

6868-060

48 0 500 2.39 2.4 14 0

6868-068

49 0 500 2.39 2.4 17 0

6868-080

50 0 500 2.39 2.4 20 0

6868-0147 

f

50 0 500 2.39 2.4 20 0

Figure 24.- Length effect on closed cavity heating for Test 6868 – Reθ=500, 

H/δ=2.39. 

 
96 Test- Run Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image

h/href

6868-051

28 -10 503 2.39 2.4 14 0

6868-052

29 -10 503 2.39 2.4 17 0

6868-072

30 -10 503 2.39 2.4 20 0

6868-084

30 -10 503 2.39 2.4 20 0

6868-105

30 -10 503 2.39 2.4 20 0

6868-041

31 -10 503 2.39 2.4 25 0

Figure 25.- Depth effect on closed cavity heating for Test 6868 – Reθ=503, H/δ=2.39. 
 

97

6868-067

32 -10 503 2.39 2.4 30 0

6868-142 

f

28 -10 503 2.39 2.4 14 0

6868-126 

f

30 -10 503 2.39 2.4 20 0

6868-127 

f

32 -10 503 2.39 2.4 30 0

Figure 25.- Depth effect on closed cavity heating for Test 6868 – Reθ=503, H/δ=2.39

(Concluded). 

 
98 Test-Run Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image

h/href

6868-078

36 -10 503 5.20 2.4 14 0

6868-102

37 -10 503 5.20 2.4 17 0

6868-107

38 -10 503 5.20 2.4 20 0

6868-098

39 -10 503 5.20 2.4 25 0

6868-144 

f

38 -10 503 5.20 2.4 20 0

6868-113 

f

40 -10 503 5.20 2.4 30 0

Figure 26.- Depth effect on closed cavity heating for Test 6868 – Reθ=503, H/δ=5.20. 

 
 

99

Test-Run 

Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image

h/href

6868-056

41 -10 260 0.648 2.4 14 

6868-073

42 -10 260 0.648 2.4 17

6868-062

43 -10 260 0.648 2.4 20 

6868-112

44 -10 260 0.648 2.4 25

Figure 27- Depth effect on closed cavity heating for Test 6868 – Reθ=260, H/δ=0.65.

 
100 Test-Run Config. 

α

(deg) Reθ H/δ W/H L/H λ(deg) Cavity Heating Image

h/href

6868-079

36 -10 264 2.7 2.4 14 0

6868-101

37 -10 264 2.7 2.4 17 0

6868-106

38 -10 264 2.7 2.4 20 0

6868-096

39 -10 264 2.7 2.4 25 0

6868-143 

f

38 -10 264 2.7 2.4 20 0

6868-114 

f

40 -10 264 2.7 2.4 30 0

Figure 28.- Depth effect on closed cavity heating for Test 6868 – Reθ=260, H/δ=2.7. 

 
 

101

TestRun 

Config. 

α

(deg) 

Reθ H/δ W/H L/H dp/dx 

(ft-1) Cavity Heating Image

6868-035

45 

s

0 300 1.1 2.4 20 -0.005 

6868-037

46 

s

0 300 1.1 2.4 20 0.005 

6868-045

47 

s

0 300 1.1 2.4 20 0.011 

6868-036

45 

s

-10 500 2.4 2.4 20 -0.005 

6868-038

46 

s

-10 500 2.4 2.4 20 0.005 

6868-046

47 

s

-10 500 2.4 2.4 20 0.011 

Figure 29.- Pressure gradient effect on closed cavities for Test 6868. 
102 

TestRun 

Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image

6868-117

33 

s

0 300 1.1 2.4 20 15

6868-121

34 

s

0 300 1.1 2.4 20 30

6868-120

35 

s

0 300 1.1 2.4 20 45

6868-135 

f

33 

s

0 300 1.1 2.4 20 15

6868-136 

f

34 

s

0 300 1.1 2.4 20 30

6868-139 

f

35 

s

0 300 1.1 2.4 20 45

Figure 30.- Cross flow effect on closed cavities for Test 6868 – Reθ=300, H/δ=1.1. 
 

103

TestRun 

Config. 

α

(deg) 

Reθ H/δ W/H L/H λ

(deg) Cavity Heating Image

6868-118

33 

s

-10 503 2.394 2.4 20 15

6868-122

34 

s

-10 503 2.394 2.4 20 30

6868-119

35 

s

-10 503 2.394 2.4 20 45

6868-134 

f

33 

s

-10 503 2.394 2.4 20 15

6868-137 

f

34 

s

-10 503 2.394 2.4 20 30

6868-138 

f

35 

s

-10 503 2.394 2.4 20 45

Figure 31.- Cross flow effect on closed cavities for Test 6868 – Reθ=500, H/δ=2.39. 
104 

Appendix: Derivation of Pressure Gradient Surface 

Geometry 

An inviscid approximation for a constant pressure gradient surface can be obtained using Modified Newtonian Theory (Anderson, 1989). Accordingly, 

cp = cp,max sin2θ

where for M→∞

cp,max → ( ) γ + 1 2

4γ

⎡

⎣

⎢

⎢

⎤

⎦

⎥

⎥

4

γ + 1

⎡

⎣

⎢ ⎤

⎦

⎥

For γ=1.4, cp,max→1.839. Assuming small surface deflections, it follows that 

cp ≈ cp,maxθ2 = cp,max

dy

dx

⎛

⎝

⎜ ⎞

⎠

⎟

2

For a constant surface pressure gradient 

d

dx

cp

cp,max

⎛

⎝

⎜

⎞

⎠

⎟ = d

dx

dy

dx

⎛

⎝

⎜ ⎞

⎠

⎟

2

= a0

Integration from x0 to x yields 

y = ±

2

3a0

⎛

⎝

⎜

⎞

⎠

⎟ a0 ( ) x

3 ⎡ 2

⎣⎢ ⎤⎦⎥

after applying the boundary conditions y=y0=0 and (dy/dx)0=0 at x=x0=0. The constant is 

given by 

a0 = 1

1.8

dCp

dx

Three gradient surfaces were designed within the allowable constraints imposed by the 

original model design. The surfaces are described in the Figure A-1, along with the surface contour. 
 

105

Config

a0

dCp/dx 

(ft-1)

45

s

-.00025 -.005

46

s

0.00025 0.005

47

s

0.00050 0.011

Figure A-1.- Pressure gradient surface definition. 
REPORT DOCUMENTATION PAGE Form Approved

OMB No. 0704-0188

2. REPORT TYPE 

Technical Memorandum

 4. TITLE AND SUBTITLE

Cavity Heating Experiments Supporting Shuttle Columbia Accident 

Investigation

5a. CONTRACT NUMBER

 6. AUTHOR(S)

Everhart, Joel L.; Berger, Karen T.; Bey, Kim S.; Merski, N. Ronald; 

Wood, William A.

 7. PERFORMING ORGANIZATION NAME(S) AND ADDRESS(ES)

NASA Langley Research Center

Hampton, VA 23681-2199

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National Aeronautics and Space Administration

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14. ABSTRACT

The two-color thermographic phosphor method has been used to map the local heating augmentation of scaled idealized cavities at conditions simulating the 

windward surface of the Shuttle Orbiter Columbia during flight STS-107. Two experiments initiated in support of the Columbia Accident Investigation were 

conducted in the Langley 20-Inch Mach 6 Tunnel. Generally, the first test series evaluated open (length-to-depth less than 10) rectangular cavity geometries 

proposed as possible damage scenarios resulting from foam and ice impact during launch at several discrete locations on the vehicle windward surface, though 

some closed (length-to-depth greater than 13) geometries were briefly examined. The second test series was designed to parametrically evaluate heating 

augmentation in closed rectangular cavities. The tests were conducted under laminar cavity entry conditions over a range of local boundary layer edge-flow 

parameters typical of re-entry. Cavity design parameters were developed using laminar computational predictions, while the experimental boundary layer state 

conditions were inferred from the heating measurements. An analysis of the aeroheating caused by cavities allowed exclusion of non-breeching damage from 

the possible loss scenarios being considered during the investigation.

15. SUBJECT TERMS

Space Shuttle Columbia; Accident Investigation; Cavity; Fluids; Heating; Hypersonic; Phosphor Thermography

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106

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(443) 757-5802

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377816.06.03.03.08

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01- 03 - 2011