ELECTRON DYNAMICS DIVISION e f*/7/177 

NASA CONTRACTOR REPORT 

INTERIM TECHNICAL REPORT 

ADVANCED RADIATOR CONCEPTS 

UTILIZING HONEYCOMB PANEL 

HEAT PIPES (STAINLESS STEEL) 

(BASA-CB-171977) ADYAIPCEC RACZAIiCR COBCEPTS 

UTXLIZIIJG HCIYETCCtB PAPEL BEAl EXPBS 1 387-16755 

I (SIPAIELESS STBE$) Interim Trcbnical Begort 

I (Eughes Aircraft Co.) 70 c CSCL, 200 Unclas 

63/34 313731 

- 

CONTRACT NAS 9-16581 

W-30746 
NASA CONTRACTOR REPORT 

INTERIM TECHNICAL REPORT 

ADVANCED RADIATOR CONCEPTS 

UTILIZING HONEYCOMB PANEL 

HEAT PIPES (STAINLESS STEEL) 

CONTRACT NAS 9-16581 

W-30746 

PREPARED FOR 

NATIONAL AERONAUTICS AND SPACE ADMINISTRATION 

LYNDON B. JOHNSON SPACE CENTER 

HOUSTON, TEXAS 77058 

0. L. FLEISCHMAN 

H. J. TANZER 

AUGUST 1985 

HUGHES AIRCRAFT COMPANY 

ELECTRON DYNAMICS DIVISION 

3100 WEST LOMITA BOULEVARD 

P.O. BOX 2999 

TORRANCE, CA 90509-2999 
NOMENCLATURE 

W 

A 

D 

g 

K 

L 

Leff 

Qmax 

Qe 

QS 

QW 

TFV 

TFQ 

WFA P 

APT 

n 

Ap C 

Apt 

ApV 

AP 

g 

r 

rV 

e 

x 

C 

p!2 

Wick cross-sectional flow area 

Honeycomb core depth 

Acceleration due to gravity 

Permeability of wick 

Length 

Effective length of heat pipe 

Quantity of heat, maximum 

Entrainment limit 

Sonic limit 

Wicking limit 

Liquid tortuosity factor 

Vapor tortuosity factor 

Localized pressure drop magnification 

Total pressure drop 

Pressure drop of n f 1,2,3. . . Capillary head pressure difference 

Pressure drop due to gravity 

Pressure drop in the liquid 

Pressure drop in the vapor 

Effective capillary radius 

Hydraulic radius of vapor space 

Contact angle 

Latent heat of vaporization 

Dynamic viscosity of liquid 

Dynamic viscosity of vapor 

Density of liquid 

Density of vapor 

Surf ace tens ion 

iii 
FOREWORD 

This interim report was prepared by the Hughes Aircraft Company, Electron 

Dynamics Division, for the NASA Johnson Space Center. 

The purpose of this program was to determine the feasibility of using honeycomb panel heat pipes as reliable, lightweight and highly efficient radiators 

for future space station applications. The scope of this program includes the 

design, fabrication, testing and delivery of two prototype heat pipe panels, 

which are 24 inches (0.61 m> wide by 120 inches (3.0 m> long. This interim 

report describes the first unit, which was fabricated from stainless steel. 

The second panel shall be fabricated from aluminum. Upon completion of the 

aluminum unit testing a final technical report will be prepared. 

This program is being conducted in accordance with the Statement of Work in 

NASA Contract NAS9-16581. Mr. G. L. Fleischman is the Hughes, Electron 

Dynamics Division, Project Manager while Mr. A. Basiulis serves as both administrative and technical adviser at Hughes. Mr. H. J. Tanzer was responsible 

for the stainless steel honeycomb panel heat pipe design, analysis and fabrication. Mr. J. T. Burdette was responsible for heat pipe processing and 

testing. Technical direction was provided by Mr. J. G. Rankin, Technical 

Representative, NASA Johnson Space Center. 

V 
TABLE OF CONTENTS 

Sect ion Page 

1.0 SUMMARY 1 

2.0 INTRODUCTION 3 

3.0 DESIGN EVALUATION 5 

3.2 The Honeycomb Panel Heat Pipe Fin 5 

3.2.1 Concept 

3.2.2 Manufacturing Technology Status 

3.2.3 Design Details and Constraints 

5 

8 

11 

3.3 Analysis 

3.4 Final Design 

15 

18 

4.0 FABRICATION 25 

4.1 Pieceparts 25 

4.1.1 Diffusion Welded Material 

4.1.2 Core Ribbon Fabrication 

4.1.3 Panel Facesheet Joining 

25 

25 

27 

4.2 Honeycomb Panel 29 

5.0 'TEST PERFORMANCE 33 

5.1 Test Setup 

5.2 Processing Procedures 

5.3 Test Results 

33 

37 

38 

5.3.1 Ambient Air Testing 

5.3.2 Performance Testing Over the Temperature 

Range -4 to 149OF (-20 to 65OC) 

5.3.3 Liquid Fill Test 

5.3.4 Tilt Test 

5.3.5 Burst Pressure Test 

38 

42 

46 

48 

52 

6.0 DATA CORRELATION AND CONCLUSIONS 57 

6.1 Data Correlation 

6.2 Conclusions 

57 

65 

vii 

~ 
LIST OF ILLUSTRATIONS 

Figure 

1 

2 

3 

4 

5 

6 

7 

8 

9 

10 

11 

12 

13 

14 

15 

16 

17 

18 

19 

20 

21 

22 

23 

Sketch of space radiator element. 

Original requirements and test conditions. 

Honeycomb heat-pipe panel concept. 

Core ribbon details 

Sketch of heat-pipe honeycomb panel. 

Panel manufacture constraints. 

Liquid flow path schematic diagram: 1 cell. 

Flange section liquid flow model. 

Performance limits vs temperature for 1 inch 

honeycomb cell panel. 

Performance limits vs temperature for 0.25 i 

honeycomb cell panel. 

thick 

ch thi k 

Performance limits vs temperature for 0.25 inch thick 

honeycomb channel panel. 

Panel performance vs panel thickness. 

Comparison between initial and final panel designs. 

Cross section of facesheet/sintered wick interface (250x1. 

Core ribbon material after crimping. 

Butt welded face sheet sample showing sintered screen wick. 

Butt welded (GTAW) face sheets for heat pipe. 

Close-up of internal honeycomb structure. 

Completed honeycomb panel. 

Performance test set-up. 

Schematic diagram of test station. 

Honeycomb panel performance test instrumentation. 

Performance test temperature distribution at 180 watts 

heat input. 

ix 

Page 

6 

6 

7 

12 

14 

14 

17 

19 

20 

21 

22 

23 

24 

26 

26 

28 

28 

30 

31 

34 

35 

36 

39 
LIST OF ILLUSTRATIONS (CONTINUED) 

Figure 

24 

25 

26 

27 

28 

29 

30 

31 

32 

33 

34 

35 

36 

37 

38 

39 

40 

Strip chart temperature recording at 400 watts. 

Temperature recording at 400 watts after removal of 

trapped liquid in corner (6, 24). 

Temperature recording at 100 watts (Cost Test). 

Temperature recording at 500 watts (Elevated Temperature). 

Temperature recording at 400 watts (110% fill). 

Maximum heat transport capacity test (130% fill). 

Verification of excess liquid in corner identified as 

6, 24 (130% fill). 

Verification of excess liquid in corner identified as 

17, 29 (130% fill). 

Honeycomb panel heat pipe tilt test performance. 

Burst pressure test. 

Honeycomb panel shown as comprised to "unit cells". 

Predicted performance limits upgrade for stainless 

steel/methanol honeycomb heat pipe. 

Data correlation: thermal transport limitation. 

Data correlation: tilt test performance. 

Summary of cell hole pattern effect on vapor flow direction. 

Detail view of panel section having mixed direction 

channels. 

Predictions for as-designed and increased capacity designed 

using correlated model. 

Page 

40 

41 

44 

45 

47 

49 

50 

51 

53 

54 

59 

62 

63 

64 

67 

68 

71 

X 
LIST OF TABLES 

Table 

1 

2 

3 

4 

Manufacturing Limit 

Preliminary Design Summary; Basis for Analytical Modeling 

Honeycomb Cell Trade-off Parameters 

Details of Wicking Limitations: 

Balance Relationship 

Differential Pressure 

Honeycomb Panel Temperature Distribution at 100 WATTS 

(Cold Test) 

Honeycomb Panel Temperature Distribution at 500 Watts 

(Elevated Temperature) 

Summary of Burst Pressure Data 

Summary of Analysis Upgrading 

Summary of Predicted Pressure Losses 

Page 

10 

10 

16 

16 

43 

43 

55 

58 

65 

xi 
1.0 SUMMARY 

The feasibility of fabricating and processing moderate temperature range heat 

pipes in a low mass honeycomb sandwich panel configuration for highly efficient 

radiator fins for the NASA space station was investigated. A variety of honeycomb panel facesheet and core-ribbon wick concepts were evaluated within constraints dictated by existing manufacturing technology and equipment. Concepts evaluated include: type of material, material and panel thicknesses, 

wick type and manufacturability, liquid and vapor communication among honeycomb cells, and liquid flow return from condenser to evaporator facesheet areas. 

In addition, the overall performance of the honeycomb panel heat pipe was evaluated analytically. 

A thin-wall (0.018 inch), all-welded stainless steel design was used for the 

initial prototype unit described in this report. Stainless steel has excellent weldability and forming properties. Methanol was selected as the working 

fluid because of its favorable thermal characteristics and vapor pressure over 

the -4 to 149'F (-20 to 65OC) temperature range, and compatibility with the 

stainless steel envelope material. The design goal was to build a 12 inch 

wide by 240 inch long (0.30 m x 6.10 m) panel that could dissipate 1000 watts 

over the -4 to 149'F temperature range at lg test conditions. 

In the interest of minimizing weight, the thinnest core depth of 0.25 inch 

(6.35 mm) and the largest hexagonal cell size within manufacturing limits of 

0.50 inch (12.70 mm) were chosen for construction. The hexagonal cell was 

chosen over channels because of prior experience in the fabrication of this 

type of heat pipe and the ability to operate the panel in either the longitudinal or transverse direction. A single layer of 120 mesh stainless steel 

screen was diffusion bonded to the facesheet for good thermal conductance. A 

porous sintered screen core, perforated to allow vapor communication, was 

selected for improved transport capacity over the screenffoil core material. 

Originally the test panel was intended to be 12 inches wide by 240 inches long 

(0.30 m x 6.10 rn) with a 0.50 inch (12.70 mm) wide heater running the entire 

length along one edge of the panel. However, these requirements were modified, 

1 
as a result of the design evaluation. 

currently machine limited to a maximum uninterrupted length of 120 inches 

(3.05 m). Therefore, the panel length was reduced to 120 inches (3.05 m), and 

the width was doubled to 24 inches (0.61 m). The heater was relocated to a 

1 inch (2.54 cm) wide strip along the centerline of the top surface. This 

revised test condition maintained the same heat input flux of 8.33 W/in.2 

(1.29 W/cm2), as before, and more realistically simulated an actual space radiator fin application. In this configuration the panel delivered a maximum 

power of 600 watts at an operating temperature of 122'F (50OC). 

value was 500 watts. The panel was isothermal throughout the entire surface 

to within - +3.6OF (+2OC), - except for a 6 by 6 inch (0.15 x 0.15 m) zone in one 

corner. As discussed further in this report, this particular corner was thermally isolated from the active panel. It is felt that this was related to the 

fact that vapor holes were placed only in alternate crimps of the core ribbon 

material rather than every crimp. 

The honeycomb panel manufacturer* is 

The predicted 

Additional tilt testing was performed by relocating the heater along one edge 

of the panel. Maximum powers of 70 watts and 50 watts were demonstrated at a 

panel elevation of 1/2 inch (12.70 mm) and 1.0 inch (2.54 mm), respectively. 

The corresponding predicted values were 77 watts and 59 watts. 

The stainless steel honeycomb panel described in this report weighed 1.9 lbs/ft2 

(9.2 kg/m*). 

panel was subjected to a burst pressure test. The core material failed in 

tension at an internal pressure of 250 psi (17.2 x lo5 N/rn2). 

in the wires of the core wick material adjacent to the spotwelds. This shows 

that honeycomb panels of even higher internal pressure capability can be 

designed and constructed in the future. Other recommendations for improved 

designs are given in the Conclusions, Section 6.0, of this report. 

A 5 1/2 inch by 5 1/2 inch (0.14 x 0.14 m) subscale honeycomb 

The failure was 

* AstechR, Division of the TRE Corp., Santa Ana, California 

2 
2.0 INTRODUCTION 

Future space systems will have power and heat rejection requirements that existing technology cannot effectively handle. 

completely new radiator system approach was suggested.' 

of a multikilowatt power module comprised of many high-capacity (~2 kw), large 

('b50 ft. long x 2 ft. wide) heat pipe radiator elements, each independently 

coupled to a centralized fluid heat source. 

approach would allow for construction of any desired size radiator system, and 

for easy on-orbit maintenance and replacement. NASA has initiated development 

of this approach, which became known as the Space Constructible Radiator 

(SCR). In an initial study on manned platforms2 and an extension which 

focused on unmanned platforms ,3 development of constructible radiator 

technology was judged to have significant potential for heat rejection system 

level improvements. This includes high-capacity heat pipes, efficient ''plugin" contact heat exchangers, and lightweight efficient radiator fins. 

From a heat rejection viewpoint, a 

The concept consists 

This modular building-block 

4 

As the performance of large heat pipes continues to increase, a corresponding 

improvement must be made in the efficiency of longer radiating fins in order 

to be able to make use of the heat pipe's maximum transport capabilities and 

therefore minimize the total number of heat pipes required for a given total 

system size. 

using the honeycomb heat pipe5 as reliable, lightweight and highly efficient 

space radiator fins.6 

evaluate a representative segment of a full sized radiator fin. 

consisted of the following phases: 

The purpose of this program is to investigate the feasibility of 

The program objective is to design, fabricate, test and 

The program 

I - Analysis and Preliminary Design 

I1 - Final Design and Materials Procurement 

111 - Fabrication 

IV - Test 

This document describes the results of the work performed in each phase. 

3 
3.0 DESIGN EVALUATION 

3.1 REQUIREMENTS 

As the modular SCR system is presently perceived, the condenser of the monogroove high capacity heat pipe7 is imbedded in a double-sided radiator fin, as 

illustrated in Figure 1. The honeycomb fin picks up heat at the interface 

with the condenser section of the monogroove heat pipe, and then distributes 

heat over the surface of the radiating fin. 

The program requirements are to design, fabricate, test, and evaluate a representative segment of a full sized radiator fin. Original panel dimensions 

were nominally set at 240 in. (6.10 m) long by 12 in. (0.30 m) wide. Performance requirements were to dissipate 1000 W over the temperature range of -4 to 

149'F (-20 to 65OC) at 1 g test conditions. Heat input was to be along one of 

the long panel edges, and to be transported in the short in-plane direction. 

The original requirements and test conditions are shown in Figure 2. 

requirements were modified as a result of the design evaluation, and the final 

test configuration is presented in Figure 13b. 

These 

3.2 THE HONEYCOMB PANEL BEAT PIPE FIN 

3.2.1 Concept 

The heat pipe panel, shown schematically in Figure 3, consists of a wickable 

honeycomb core, internally wickable facesheets, and an appropriate working 

fluid. Evaporation of the working fluid occurs at any section of the panel 

exposed to heating. Vapor will flow to a cooler region where it condenses, 

and the condensate will return to the evaporator by means of capillary pumping 

action of the wick structure. 

to allow intercellular liquid flow along the faces, and perforated to allow 

intercellular vapor flow. 

is necessary to ensure heat pipe action both in the plane of the panel and 

through its depth. The primary mode of heat transfer can be either transverse 

(face-to-face) or longitudinal (in-plane), with the degree in either mode 

The honeycomb cells can be notched at both ends, 

The intercellular communication of liquid and vapor 

5 

PAGE, __ INTEN'C1ONALLY BLANK 
G14251 

Figure 1 Sketch of space radiator element. 

G 14252 

QIN 1000 W HEAT INPUT AREA = 2 (0.5 x 120 IN.) = 120 IN2 (774.19 CM2) 

- Q 1OOOW = - 

A 120 IN2 

8.33 W/IN2 (1.29 W/CM2) 

PAN EL DIM ENS IONS 

LENGTH (L): 240 IN. (6.10 MI 

WIDTH (W): 12 IN. (0.30 M) 

DEPTH (D): TBD 

TEST OR I ENTATION 

EVAP 0.125 IN. (3.18 MM) ABOVE COND 

Y 

xk 

D 

Lc 

r 

= 11.5 IN. (0.29 M) 

I I 

/ Le =0.5 IN. -+- (12.7 MM) ' QOUT 

QI N I 

rLFigure 2 Original requirements and test conditions. 
G14165 

HONE 

-c-- 

TOP FACESHEET 

WICKABLE HONEYCOMB: 

NOTCHED TO ALLOW 

RATED TO ALLOW 

VAPOR FLOW 

LIQUID FLOW; PERFOMETAL SCREEN SINTERED TO 

INTERNAL FACES TO ALLOW 

ALONG FACES 

IN-PLANE FLOW OF LIQUID 

YCOMB CELL 

FLANGES FOR BONDING 

HONEYCOMB TO FACESHEETS 

Figure 3 Honeycomb heat-pipe panel concept. 

7 
varying with the wickable core design. 

require a wickable core design to enhance the longitudinal heat transfer capability. Applying this concept to the space radiator will result in an across 

the panel temperature gradient of essentially zero, while avoiding warpage due 

to thermal stresses and improving fin efficiency, which will approach 1.0. In 

addition, the honeycomb design offers high reliability. 

this design can be constructed in narrow, independent segments to provide 

redundancy for meteoroid or other forms of damage. 

The present radiator application will 

A radiator system of 

3.2.2 Manufacturing Technology Status 

The technology and commercial equipment are available to construct all-welded, 

machine-assembled honeycomb panels. At present, such panels are constructed 

and formed into various shapes for use in aircraft, missile and ship frames. 

The honeycomb structure can be manufactured from any weldable material 

(excluding aluminum), up to 48 in. (1.22 m) wide in any reasonable length, 

with a minimum overall thickness of 0.25 in. (6.35 mm), in a variety of cell 

and channel sizes from 1/4 to 1/2 in. (6.35 to 12.70 mm) and shapes with 

different facesheet thicknesses. The basic panel is readily producible into 

components by cutting, stretch-forming, drawing, welding, and riveting. 

For the honeycomb panels to function properly as heat pipes, the internal surfaces must be fabricated from porous materials or have porous materials 

attached to the internal surfaces. The structure would consist of two internally wickable faces bonded to perforated, wickable honeycomb core material. 

Calculations and experiments with pieceparts led to the development of 

machine-fabricated, wickable honeycomb subscale liquid metal test panels for 

thermal stress reductions in NASA Scramjet Engines.8 

prototype panels were built to evaluate alternative construction methods, to 

perform proof-pressure and weld integrity testing, to verify heat-pipe processing techniques, and to do performance testing. Evaluation of options 

resulted in final design and fabrication of 6 in. square by 1.14 in. 

(0.15 x 0.03 m) thick test panels, constructed entirely of stainless steel 

materials. Two designs for the honeycomb core were built: a foil-gauge 

sintered screen material, and a metal screen sintered to foil-gauge sheet 

material. The former design offers increased wicking capability and the 

Various test samples and 

8 
latter provides stronger structural design. Details of final panel 

construction (welding of sidewalls), cleaning and processing procedures, and 

experimentally-determined wick parameters (capillary radius, r,., and 

permeability, K) for porous core materials are reported in reference 8. 

Table 1 outlines current manufacturing limits on the honeycomb heat-pipe panel 

design. 

Stainless steel and the lighter-weight titanium construction materials were 

both considered for panel fabrication. Stainless steel was chosen based on 

proven wickable core fabrication experience, and due to the availabilty of a 

much larger range of screen mesh sizes. The alternative of welding stainless 

steel screen core to titanium facesheets was disapproved by the panel manufacturer due to the formation of brittle welds. The facesheets are made internally wickable by sintering relatively porous 120 x 120 mesh 316 stainless 

steel square weave screen to 0.018 in. (0.46mm) thick 316 stainless steel 

sheet.’ 

thereby improving transport capacity of the heatpipe honeycomb panel. 

panel manufacturer felt that more than two layers would create excessive 

arcing during the weld operation, leading to a high probability of pinholes 

through the facesheet. Correspondingly, a minimum starting sheet thickness of 

0.018 in. (0.46 mm) was recommended for high probability leak-free panels. If 

desired , facesheets could be made thinner after final core assembly through 

chem-milling. Two types of wickable honeycomb core designs were considered: 

316 stainless steel sintered woven wire screen; and a 316 stainless steel 

screen/foil composite consisting of one layer screen sintered to 316L stainless steel foil. The sintered screen core permits liquid flow between cells 

because it is porous through its thickness. The screen/foil core relies 

solely on notches cut in the top and bottom of each honeycomb cell wall 

(Figure 3) for liquid flow between cells. Since the higher structural 

strength of the screenlfoil core is not necessary for the space radiator, the 

sintered screen core was selected to improve transport capacity. The selected 

core consists of 165 x 1400 mesh sintered screen having a thickness of 

0.006 inch (0.15 mm). 

Using multiple facesheet wick layers would increase liquid flow area, 

The 

Alternate core ribbon designs were investigated with regard to manufacturing 

constraints and their implication on heat pipe capacity. Core ribbons are 

9 
TABLE 1 

MANUFACTURING LIMITS* 

Facesheet 

0 Longest piece before welding: 48 in. (0.64 m) 

Honeycomb Panel 

0 Materials: stainless steel, titanium 

0 Longest section before welding: 120 in. (3.05 m) 

0 Panel width: 48 in. (0.64 m) 

0 Core depth: 0.25 to 2.00 in. (6.35 to 50.8 mm) 

0 Core ribbon thickness: 0.006 in. (0.15mm) maximum 

0 Cell or channel sizes: 0.25, 0.375 and 0.5 in. 

(6.35, 9.52 and 12.70 mm) 

0 Facesheet thickness: 0.010 to 0.030 in. 

(0.25 to 0.76 mm) 

*Dictated by materials supplier and panel manufacturer 

TABLE 2 

PRELIMINARY DESIGN SUMMARY; BASIS FOR ANALYTICAL MODELING 

Select design parameters 

Fixed 

0 All-stainless steel construction 

0 Facesheet strip thickness: 0.018 in. (0.46 mm) 

0 Wire mesh laminate for core ribbon 

0 Overall length of panel: 120 in. (3.05 m> 

0 Overall panel width: 24 in. (0.60 m> 

Vary 

0 Core ribbon depth (D): 0.25, 0.5, or 1.0 in. 

(6.35, 12.70 or 25.40 mm) 

0 Honeycomb size: 0.375 or 0.5 in. (9.52 or 12.70 mm), 

cell or channel 

0 Facesheet wick layers: 1 or 2 

Optimize performance 

10 
initially punched to create desired holes or notches, are then corrugated and 

have their top and bottom edges folded to form mini-flanges. 

cell core is formed when the core ribbons are welded to each other at face 

junctions and to the facesheets and mini-flange intersections (Figure 4). 

size, shape and location of honeycomb cell wall perforations are trade-offs 

between structural soundness and heat pipe performance. Vapor flow perforations in cell walls should be no larger than is required to keep the sonic and 

entrainment limits of the heat pipe from becoming overall performance limiters. 

Notches at the core ribbon/facesheet interface increase liquid flow due to 

unobstructed wick area, however, good weld consistency requires that the miniflange itself remain intact in order that weld resistance be constant during 

the welding operation. 

and thereby effectively decreases liquid flow area (inhibiting transport capacity by reducing the heat pipe wicking limitation). 

facturer is intent on providing high quality structural resistance to shear 

and tension through close spacing of spotwelds; however, to increase liquid 

flow area for this application, increase in spotweld spacing should be considered. Alternately, the core can be constructed with channels placed in the 

direction of heat transport, which eliminates concern both for perforations 

(not needed) and for liquid flow pressure drop due to spotwelds. 

The honeycomb 

The 

Each spot weld region destroys local wicking action 

Typically, the panel manuTable 2 lists design parameters which are considered fixed as a result of preliminary design evaluation, and those which were varied during the following 

analytical trade-off investigation. 

chosen for fabrication since this is the longest uninterrupted length which 

the panel manufacturer's present machine can produce. However, in order to 

maintain the original plan area of 20 ft2 (1.86m2), the width was doubled to 

24 inches (0.61m). A honeycomb cell size of 0.25 in. (6.35 mm) is not further 

investigated due to excessive liquid flow pressure drops at weld zones, 

resulting in the poorest thermal transport capacity of the several designs 

considered. In addition, the 0.25 in. (6.35 mm) cell size has the highest 

unit weight. 

A panel length of 120 in. (3.05m) was 

11 
G14253 

FACESHEET 

0.018 IN. (0.46 MM) 

THICK 

SCREEN, 120 x 120 MESH 

1 OR 2 LAYERS 0.0076.008 IN. (0.186.20 MM) 

THICK EACH, SINTERED TO FACESHEET 

1 \ 

CORE RIBBON 

FLANGE TO 

FACESHEET 

CORE RIBBON: 

SCREEN 0.0055 IN. (0.14 MM) THICK 

DIFFUSION - BONDED 

SPOT 

CELL 

WELD 

WALLS 

ZONE: / 

' 0 SLOTS - TWO PER CELL WALL, 

AS CLOSE TO EACH 

EDGE AS POSSIBLE h f 

0 HOLES - ONEORTWOPER 

CELL WALL, diaT CENTRALLY LOCATED 

Figure 4 Core ribbon details. 

12 
3.2.3 Design Details and Constraints 

Alleviation of face-to-face (transverse) temperature differences, hence, 

thermal stress problems in Scramjet wall structures at 1202'F (650'~) 

operating temperatures has been demonsttated.1° 

for use as a very efficient lightweight space radiator fin requires the heat 

flow to be in-plane rather than transverse. In-plane heat transfer dictates 

longer vapor flow paths through the cell walls and correspondingly longer 

liquid return paths aong the wick-covered internal surfaces. Detailed 

evaluation of wick types and configurations, honeycomb cell sizes, panel 

thicknesses, materials, and manufacturing techniques was necessary to optimize 

desired panel characteristics. 

The extension of this concept 

As a result of experimental work reported previously,8 certain design details 

and test data have been established, and are presently used as a baseline. 

The entire honeycomb panel is fabricated using an automated procedure for 

simultaneously resistance welding corrugated honeycomb core ribbons to each 

other (in the case of cells) and to both facesheets, forming 0.25, 0.375 or 

0.5 in. (6.35, 9.52 or 12.70 mm) hexagonal cell or channel configurations. A 

sketch of the heat pipe honeycomb panel and associated component parts is 

shown in Figure 5. 

The honeycomb panel manufacturer is machine limited to a maximum uninterrupted 

length of 120 inches (3.05 m). 

to produce any longer length desired. 

however, is created by weld joining of facesheet subsections. The sintered 

facesheet fabricator produces a maximum standard length of 48 in. (1.22 m), 

which dictates that honeycomb panel weld seams shall occur at least every 

48 in. (Figure 6). 

Beyond that, sections can be welded together 

The initial break in panel continuity, 

Since the heat flow is in the same direction as the weld seam, and since the 

core-ribbon crosses the welds to provide liquid communication, the detrimental 

effect on heat pipe performance is minimized. 

welding will produce a seam width of approximately 3/32 in. (2.38 mm). Electron beam welding can reduce the facesheet wick destruction zone to an absolute minimum, and is therefore the preferred method. 

Conventional plasma butt13 
G14254 

CORE RIBBONS 

RESISTANCE WELDED 

TO FACESHEETS 

CORE RIBBONS 

RESISTANCE WELDED 

TO FACESHEETS CELL SIZE 

SEAM WELD, 

0.25 ALL AROUND IN. (6.35 MM) 

MINIMUM 

PER FOR ATIONS, 

VARIOUS SHAPES 

AND SIZES 

4 

Figure 5 Sketch of heat-pipe honeycomb panel. 

G14255 

SUBSECTION WELD SECTION WELD 

48 IN. (1.22 M) 

120 IN. (3.05 M) 

Figure 6 Panel manufacture constraints. 

14 
3.3 ANALYSIS 

When the decision was made to double the panel width (Section 3.2.2), it was 

also decided to relocate the heater to a 1 inch (2.54 cm) wide strip along the 

centerline of the top surface. This revised condition maintains the same heat 

input flux of 8.33 W/in.2 (1.29 W/cm2) as before, and more realistically 

simulates the actual space radiation fin application. 

expressions were used to calculate heat pipe performance limits for 

longitudinal (24 in. direction) heat transport along a 120 in. (3.05 m) length 

of honeycomb panel. Note, however, that the panel is symmetrical with heat 

input along the centerline as described above. Therefore, the analysis was 

performed for only one-half the actual heater and panel widths. 

Standard analytical 

Critical parameters which were varied to determine the effect on maximum 

longitudinal thermal transport were: core ribbon depths of 0.25, 0.5 and 1.0 

in. (6.35, 12.70 and 25.40 turn); 0.375 in. (9.52 mm> cell and channel and 0.5 

in. (12.70 mm) cell sizes; and 1 or 2 facesheet wick layers. 

calculations are based on methanol working fluid at the radiator operating 

temperature range of -4 to 149'F (-20 to 65OC). 

and each channel in the longitudinal direction of thermal transport was 

modeled as a separate heat pipe; therefore, adjacent porous cell walls were 

dimensionally halved when determining wick cross-sectional liquid flow area. 

The multitude of liquid and vapor flow paths --the associated lengths and 

pressure drops -- were calculated and summed according to series or parallel 

path modeling. The following summary of heat pipe performance limitations 

contains details of the calculations. 

Transport 

Each row of adjacent cells 

Sonic and Entrainment 

For the parameters selected, these are relevant only at operating temperatures 

lower than -4'F (-2OOC). 

direction perpendicular to intended thermal transport. 

chosen parameters and assumptions used in calculating these limits. 

Vapor flow areas are represented by openings in the 

Table 3 summarizes 

15 
TABLE 3 

HONEYCOMB CELL TRADE-OFF PARAMETERS 

05 

7- 

0.25 in 0.065 x 0.025 in (Xz) 1/16 in (XI) 

0.5 in 0.065 x 0.025 in (Xz) 1/8 in (XI) 

1.0 in 0.065 x 0.025 in (X2) 1/8 in (X2) 

Core Depth (D) 

Total X-Sect Flow Area 

Per Face Per Cell (X2) 

0.00619 in2 0.01238 in2 

0.0 1539 0.03078 

0.02766 0.05532 

TABLE 4 

DETAILS OF WICKING LIMITATIONS: DIFFERENTIAL 

PRESSURE BALAiiCE RELATIONSHIP 

C' aPC = 2ucos8/rc 

AP =pg(LsinC + D) 

Fluid viscosity u, wetting angle 8, capillary radius r 

Lsin c = 0.125 in. minimum; tilt angle9, length L 

g D = 0.25, 0.50, or 1.0 in.; core depth 

Assume laminar flow through area given by hydraulic radius of vapor space 

Leff = 6.0 in.; effective length 

Absolute vapor viscosity pv, heat transfer Q, vapor density pv I 

"n All series flow paths: APT = API + AP2 + AP3 + ... 

All parallel flow paths:,p; I 111 +Ap2 Ap3 1 

- + - +--+A7 

Total pressure dropApT, wall wick area Aw, permeability K, latent heat o vaporization h 

Material 

120x120 facesheet 84.9~ screen 

165x 1400 core screen 7.5 1 x 10" I 

16 
2 An evaporator heat flux of 8.33 W/in2 (1.29 Wlcm ) is lower than measured data 

of 13 to 32 W/in2 (2.02 to 4.96 W/cm 

combinations, l1 and is therefore non-limiting. 

2 for methanollstainless steel heat pipe 

Wicking 

For all cases considered, capillary pumping restricts maximum heat transport. 

Table 4 provides details of the differential pressure balance relationship 

which defines maximum heat transport. Figure 7 briefly outlines the liquid 

mass flow (A) path methodology for one honeycomb cell. 

The spot welded regions both at core ribbon flange to facesheet and at core 

ribbon cell wall interfaces were assumed to be 75 percent (25 percent porous) 

impervious to liquid flow. A visual inspection on remnant material revealed 

variation in spot weld consistency. This weld variaton made it difficult to 

establish a percentage factor for open wick cross-sectional area. The initial 

0 14256 

TOP VIEW 

SIDE VIEW 

FACESHEET WICK 

CORE R18EON 

FACESHEET WICK ---- 

LEFT SIDE 

Figure 7 Liquid flow path schematic diagram: 1 cell. 

17 
estimate of 25 percent porous zone may be optimistic. Spot weld resistances 

in the cell configuration contribute significantly to overall liquid flow 

pressure drops (APE). Figure 8 shows a detailed schematic model of spot 

welded core ribbon flange to facesheet region. Predicted performance curves 

for several of the wickable core designs are presented in Figures 9, 10, and 

11. In all cases, two layers of wick sintered to the facesheet nearly doubles 

the transport capacity of a single layer. Honeycomb channel design roughly 

triples the transport capacity of a cell design of similar size, primarily as 

a result of reduced pressure losses of liquid flow at weld zones and vapor 

flow in open channels. This analysis shows that performance is not 

significantly affected as honeycomb cell size varies. Relatively small 

increases in performance are possible as panel thickness increases from 0.25 

to 1.0 in. (Figure 12). 

3.4 FINAL DESIGN 

As previously stated, the initial design goal was a panel size of 240 inches x 

12 inches (20 ft -plan area). 

the length was reduced from 240 inches to 120 inches. Consequently, the width 

was increased from 12 inches to 24 inches in order to maintain the overall 

plan area of 20 ft . The heat input zone was also relocated to the center of 

the panel to better simulate the intended application as a fin. Figure 13 

shows a comparison between the initial and final panel design configurations. 

2 Because of manufacturing limitation, however, 

2 

The stated radiator power dissipation requirement throughout its operating 

temperature range of -4 to 149'F 

alternative core designs (Figures 9, 10 and 11) that were considered. The 

near-term objective of this program is to prove performance. A minimum transport capacity of 1000 watts (at -4OF) is predicted for a heat-pipe honeycomb 

cell radiator panel which measures 120 by 12 by 0.25 in. (3.05 m x 0.30 m x 

6.35 mm) (Figure 10). Not requiring excess capacity, and in the interest of 

minimizing weight, the 0.25 in. (6.35 mm) core depth panel with a single layer 

of facesheet wick was chosen for construction. Moreover, the 0.5 (12.70 mm) 

inch cell option was selected over channels because of the previously successful honeycomb panel heat pipe construction (5, 8). 

(-20 to 65OC) can be satisfied by any of the 

18 
G 14257 

Lune: nioouiu 

SINTERED SCREEN 

165 x 1400,316 SST, 

0.0055 IN. (0.14 MM) ’ THICK 

0.040 IN. (1.02 MM) 

FACESHEET 

316 SST 120 x 120 SCREEN 

CELL Sl?E 

0.25,0.375 AND (0.46 MM) THICK 

0.5 IN. (6.35.9.52 

AND 12.70 MM) 

SINTERED TO 0.018 IN. 

316 SST SHEET 

A. SCHEMATIC OF HONEYCOMB CELL WALL 

SPOT-WELD 

ZONE 

(-75% OF 

CROSS SECTIONAL 

FLOW AREA) 

0.0055 IN. (0.14 MM) 

SCREEN FLANGE 

i 

0.0074 IN. (0.19 MM) SINTERED ]& 120 x 120 MESH 

\ / ** 

.#\’ 

POROUS I 

ZONE 0.018 IN. (0.46 MM) 

(-25% OF FACESHEET 

CROSS SECTIONAL 

FLOW AREA) 

6. CLOSE-UP SCHEMATIC OF CROSS-SECTIONAL FLOW AREA 

Figure 8 Flange section liquid flow model. 

19 
- G 14258 

1 

v) 

II- 

- 34 

c 

a5 2 10 

a 

a 

v) 

z 

It II 

5 

t 

/,' ENTRAINMENT 

TEST VEMICLE. 

- PANEL TMlCKNESS IHI' 1 " H" 1 HONEYCOMB 

I 

I I I I I I 

-13 32 77 122 167 212 (' 

-25 0 25 50 75 100 (OC) 

. TEMPERATURE 

Figure 9 Performance limits vs temperature for 1 inch 

thick honeycomb cell panel. 

20 
G 14259 

106 F 

Le=1/2 IN. ' Q * I SONIC/// Lc = 11.5 IN. 

PANEL - THICKNESS (HI: 0.25 IN. 

--- 

METHANOL FLUID 

-13 32 77 122 167 212 (OF) 

-25 0 25 50 75 100 (OC) 

TEMPERATURE 

Figure 10 Performance limits vs temperature for 0.25 inch 

thick honeycomb cell panel. 

21 
0 14260 106 

t 

c 

a 

v) 

g 

a 

a 

z 

t IU 

W 

I 

SONIC / 

Le = 1/2 IN. 

Lc= 11.5 IN. 

PANEL THICKNESS (HI: 0.25 IN. 

- HONEYCOMB 3'8 IN' CHANNELS 

METHANOL FLUID 

- K2 

rl 

FACESHEET WICK LAYERS 

22 
G14261 

71 

4 

w 

z 

B 

0 

6 

% 

0 

0 HONEYCOMB 

CHANNELS (3/8 IN.) 

CELLS (3/8 IN.) 

- HONEYCOMB 

FACESHEET SCREEN 

LAYERS: 

HIGHER VALUE - 2 

LOWER VALUE - 1 

4 

w 

z 

B 

(INCH 

(MM) 

Figure 12 

0.25 0.5 1 .o 

6.35 12.70 25.40 

PANEL THICKNESS 

Panel performance vs panel thickness. 

23 
G14262 

QIN = 1000 W HEAT INPUT AREA = 2 (0.5 X 120 IN.) = 120 IN2 (774.19 CM2) 

- Q 1OOOW 2 - = 8.33 W/IN2 (1.29 W/CM 

A 1201N2 

Y QOUT Lc= 

r 

11.5 IN. (0.29 MI 

f -------- 

Le * 0.5 IN. (12.7 MM) 

240 IN. (6.1M) 

A. ORIGINAL REQUIREMENTS AND TEST CONDITIONS. 

G 14263 

0.5 IN. (12.70 MM) HONEYCOMB CELL, 1 LAYER FACESHEET WICK, 

ALL SST CONSTRUCTION 

120 IN. (3.05 M) 

kL 

0.25 IN. 

(6.35 MM) ‘ I 

Q/A 1000 W/(1 IN. X 120 IN.) = 8.33 W/IN2 (1.29 W/CM2) 

B. FINAL CONSTRUCTION DETAILS AND TEST CONDITIONS. 

Figure 13 Comparison between initial and final panel designs. 

24 
4.0 FABRICATION 

As the fabrication process progressed from raw materials to finished panel, 

several design parameters deviated from the original specification as a result 

of manufacturing difficulties and limitations. The following fabrication 

sequence describes events leading to the finished panel. 

4.1 PIE CEPARTS 

4.1.1 Diffusion Welded Material 

As previously described, the facesheet material consists of one layer of 

120 x 120 mesh 316 SST screen, diffusion bonded (sintered) to one layer of 

.018 in. (0.46 mm) thick 316 SST sheet. Figure 14 is a cross-section of the 

facesheetlsintered wick interface. The largest available stock sizes for this 

material are 24 x 48 in. (0.61 x 1.22 m) and 36 x 36 in. (0.91 x 0.91 m) due 

to vacuum furnace size limitations. 

24 inches (0.61 m), the panel welder requires a starting facesheet width of at 

least 28 inches (2 inches extra at both edges for "grabbing" the panel during 

the weld operation). Therefore, the 36 x 36 in. sintered stock was ordered. 

This facesheet material, as originally received, was unacceptable due to the 

presence of small, sharp depressions. The material was reworked by flattening 

in a rolling machine. However, it required trimming in size to 26 x 36 inch. 

It follows, then, that the final panel would consist of five sections joined 

together to create a total length of 120 inches (3.05 m); in other words, 

welds would occur every 26 in. rather than 48 in. as originally planned. 

To achieve a finished panel width of 

The core ribbon material consists of sintered 316 SST twilled-weave, wire mesh 

laminate (165 x 1400 mesh) which is 0.0055 in. (1.67 mm) thick. Sheets 

measuring 24 x 48 inches were prepared. 

4.1.2 Core Ribbon Fabrication 

The ,ribbon material was cut into thin strips (48 in. long), crimped, corrugated, and then folded 90° at both 48 in. (1.22 m) edges to form mini-flanges. 

25 
E4938 

Figure 14 Cross section of facesheet/sintered wick 

interface (250X). 

E4939 

Figure 15 Core ribbon material after crimping. 

26 
Holes (perforations) were punched into the ribbon with small dies. 

on hole placement are related to cracking and buckling of the ribbon during 

subsequent panel weld operations. After some experimentation, the panel 

welder settled on the inclusion of two 1/16 in. (1.59 mm) diameter holes which 

were located very near the mini-flanges at the top and bottom of every second 

cell face (vapor flow area through the honeycomb cells is the summed open area 

from two faces per cell). A photograph (Figure 15) of the finished core ribbon shows the holes punched on alternating cell walls. However, the actual 

vapor flow area was only 75 percent of that originally desired. Also, vapor 

flow was constrained to a diagonal direction through the panel, or a 

41 percent increase in travel length from the straight direction. 

Constraints 

Liquid transport is affected by the location and spacing of spot welds in its 

flow path. 

requested. However the actual extent of this increase is unclear. Upon close 

examination of a finished panel specimen, variance in spotweld consistency 

makes it difficult to establish a percentage factor for "open" wick cross sectional area. The initial estimate of 25 percent porous zone (see Section 3.3, 

under Wicking) may be optimistic. 

sensitive to this available wicking flow area. However, epot welds at core 

ribbon cell wall interfaces (see Figure 4) were eliminated entirely (as 

requested) and this should increase transport capacity contribution of the 

core wick. 

An increase of mini-flange to facesheet spot weld spacing was 

Thermal transport capacity predictions are 

I 

I 

4.1.3 Panel Facesheet Joining 

Electron beam welding of the facesheet sections was originally planned, but 

due to cost and time constraints, the panel fabricator elected to use buttweld (GTAW), shown in Figure 16. Four butt welds of this type were used for 

welding five facesheet sections into an overall length of 120 inches (3.05 m). 

Note that these welds are in the transverse direction parallel to the liquid 

flow path. 

liquid flow communication. 

Moreover, the core ribbon material crosses these welds to provide 

27 
€4940 

Figure 16 Butt welded face sheet sample showing 

sintered screen wick. 

E4941 

Figure 17 Butt welded (GTAW) face sheets for heat pipe. 

28 
Metallurgical cross-sections of the butt-welded face-sheet samples are shown 

in Figure 17. 

facturer, AstechR, Division of TRE Corp., Santa Ana, California. 

These cross-sections were provided by the honeycomb panel manu4.2 HONEYCOMB PANEL 

Figure 18 is a close-up photograph of the internal honeycomb structure with 

one facesheet removed. The sintered core material (ribbons), perforated holes 

for vapor and liquid communication among the cells, and facesheet wick sintered to the facesheet material can be seen in this photograph. Spotwelds for 

attaching the core ribbon to the facesheets and a facesheet butt weld are also 

vis ib le. 

After fabrication and trimming of the honeycomb panel material to size 

(24 in. x 120 in.), a rolled edge was formed, completely encircling the panel. 

The edge was butt fusion (GTAW) welded all-around to produce a leak-tight seal 

(Figure 19). 

(24 in.) edge and a 1/4 inch (6.35 mm) diameter x 12 inch (0.30 m) length fill 

tube was welded in place. 

rather than two welds plus additional piece-parts €or the edges as originally 

envisioned. The completed panel is shown in Figure 19. The final leak test 

consisted of internally pressurizing the panel to 5 psig using helium, and 

verifying leak tightness with a high speed sniffer probe attached to a VeecoR 

mas s spec t rome t e r . 

A 1/4 inch diameter hole was drilled at the center of one short 

This technique resulted in a single edge weld 

29 
€4942 

CORE RIBBON: 165 X 1400 MESH 

SINTERED LAMINATE 

FACESHEET: 1 LAYER SINTERED \\ 120 X 120 MESH 

DIRECTION 

(FLOW 

RESTRICTED 

45' TO 

DIRECTION 

OF HEAT 

TRANSPORT) 

DIRECTION -I 

OF 

INTENDED 

HEAT 

\ 

PUNCH ED 

HOLE 

FACESHEET VAPOR 

SPOT WELD 

FACESHEET 

BUTT WELD 

\ 

TRANSPORT 

Figure 18 Close-up of internal honeycomb structure. 
I€ 

E :t 
5.0 TEST PERFORMANCE 

Preliminary performance testing and checkout of the honeycomb panel heat pipe 

were conducted in laboratory ambient air, For testing over the temperature 

range of -4 to 149'F (-20°C to 65OC), however, the panel was installed in a 

special test station. The test station, methods, and heat transport 

performance test results are described in the following paragraphs. 

5.1 TEST SETUP 

Figure 20 is a photograph of the test station designed for testing the honeycomb panel performance over the temperature range -4 to 149'F (-2OOC to 65OC). 

Referring to the sketch in Figure 21, it can be seen that the heat sink is 

provided by six 8 inch (0.20 m) wide flanged aluminum extrusions with 1 inch 

(25.4 mm) diameter liquid nitrogen (LNZ) coolant passages. Three extrusions 

are placed above the test panel and three below for heat rejection from both 

sides of the panel. The test panel is centered approximately 3 inches 

(76.2 mm) from the flanged surfaces of the heat sinks, using a total of eight 

adjustable plexiglass support pegs. The test setup was enclosed in a 132 inch 

long x 30 inch wide X 24 inch (3.35 x 0.76 x 0.61 m) high plexiglass chamber. 

Note that this is not a vacuum chamber, and heat transfer from the heat pipe 

to the heat sink is by radiation, conduction, and natural convection through 

the surrounding air. 

Thirty chromel-constantan (Type E) thermocouples were spot-welded directly to 

the panel surface at the locations shown in Figure 22. Thermocouples on the 

bottom surface are placed directly underneath the top ones. Note, however, 

that there are no thermocouples underneath the four circled ones (Figure 23). 

The reason for this is that a thirty channel strip chart recorder was selected 

for recording the data. Since there are no edge effects associated with these 

locations, it is felt that the top thermocouples are sufficient. 

I 

Heat input was provided by four strips of Clayborne Labs heater tape (E-16-21 

wired in parallel. 

(25.4 mm) wide strip running the entire length of the panel (120 inches), as 

shown in Figure 22. 

For initial testing the heat input zone was a one inch 

This approach simulates the honeycomb panel being used as 

33 
E4944 

COLD WALLS (LN2/ 

ALUMINUM EXTRUSIONS) 

L LEVELLING LEGS CHAM BE R 

Figure 20 Performance test set-up. 

34 
I' 

F 

a 

w 

z 

U 

atcn 

w 

tI -1 

C 

0 

r( 

U 

Q 

U 

v) 

U 

v) 

9) 

3 

M 

35 
(D 

N 

f 

0 

\ 

t 

t 

1 

w 

2 

a 

> 

c 

U 

i 

> I 

9 

c) 

x 

z 

w 

a 

0 

e 

ww 

0 

L 

rn 

C 

rl 

U 

v) 

01 

U 

L4 

al 

a 

4 

al 

C 

(d 

a 

P 

.5 

N 

N 

al 

L.c 

3 

M 

rl 

cr, 

36 
a high efficiency radiator fin in conjunction with a high capacity axial 

transport heat pipe or coolant loop. 

5.2 PROCESS ING PROCEDURES 

The test panel was received at Hughes on November 21, 1983. The panel was 

subjected to a leak test by internally pressurizing the panel to 5 psig using 

helium, and checking for leaks with a high speed sniffer probe attached to a 

VeecoTR mass spectrometer leak detector. No leakage was detected. 

Following this leak test, the panel was vacuum baked at a temperature of 

approximately 1490F (65OC). for a period of 261 hours. Electrical resistance 

heater tapes were .bonded to the external surfaces to provide an elevated 

temperature during bakeout. 

lo-' torr during the entire bakeout period. 

The panel outlet pressure was on the order of 

The panel was then processed with 1.14 lbm (0.518 kg) of methanol by vacuum 

distilling the methanol into an evacuated stainless steel cylinder. This 

cylinder was connected to a valve attached to the previously evacuated panel, 

and the methanol was drained into the panel under vacuum conditions. The 

appropriate fill was determined experimentally by weighing a small sample of 

residual honeycomb material before and after saturating the wicks with 

methanol. This approach provided the amount of methanol per unit of panel 

area. Even though the methanol in the charging cylinder was degassed prior to 

filling, the panel was also subjected to the degassing procedure to provide 

assurance against the possibility of introducing gas during the transfer of 

methanol. 

open cell polyurethane foam underneath and on top of the panel, and cooling by 

pouring liquid nitrogen onto the top layer. 

cooled (<-40°F) the fill valve was opened to vacuum to vent the gases. 

The degassing was accomplished by placing a 3-inch thick layer of 

When the panel was sufficiently 

37 
5.3 TEST RESULTS 

5.3.1 Ambient Air Testing 

As mentioned above, preliminary checkout of the honeycomb panel heat pipe was 

conducted in laboratory ambient air. The panel was instrumented, as shown in 

Figure 22, and was level to within 2 1/8 inch (+3.18 - mm) by using a large 

carpenter's level and by adjusting the screw-thread plexiglass support legs. 

An initial power of 180 watts was applied to the heater with the panel on the 

bench. The resulting steady state temperature distribution is given in Figure 23. Note that the heater is on the top surface only. The entire active 

surface of the panel is isothermal with - +1.8'F (+ -1OC). Next the power was 

increased to 280 watts and 400 watts. Figure 24 is a strip chart recording 

showing all 30 temperature traces at 400 watts. At 570 watts there was 

evidence of hot spots (local dryout) in the area of thermocouple numbers 2 and 

19. 

In both Figure 23 and Figure 24 it can be seen that there are two cold corners 

on the panel, as indicated by thermocouple numbers 69 16, 24 and 30. These 

inactive zones were approximately 6 inches by 6 inches (0.15 x 0.15 m) at each 

corner. Since both of these corners were on the same longitudinal edge, that 

side was elevated to distinguish possible liquid slugging from gas. With 

sufficient time, the thermocouples (6, 24) on the corner opposite the process 

tube and valve fell in line with the others, leaving only one inactive area 

(16, 301, as shown in Figure 25. This indicates that the corner identified as 

6, 24 was observed to be cold due to the accumulation of liquid in that area, 

while the other corner must be either noncondensable gas or closed off from 

the remainder of the heat pipe. 

Acting on the assumption that the other corner was due to gas, the heat pipe 

was connected to the vacuum system for a second degassing attempt. 

resulted in no observable change in performance. Next, heaters were bonded in 

each corner near the process tube and insulated. These heaters were allowed 

to operate approximately eighteen hours (overnight) in an attempt to displace 

gas and disperse it to another location in the panel. The centerline heater 

was started while the corner heaters were still on. Then, when this heater 

This 

38 

~ 
In 

(D e N 

c 

0 

K 

w 

a 

w 

I 

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c 

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w 

4 

W 

c 

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v) 

u 

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s 

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a3 

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e 

0 

4 

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-0 

al 

h 

1 

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0 

wu 

ha ala @Is 

39 
v) 

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0 

0 

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MU 

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41 
was increased to 300 watts, the insulation was removed and the corner heaters 

were turned off. Thermocouple numbers 16 and 30 immediately began a steady 

drop-off to return to their original position. For these reasons it was concluded that this was a "dead corner." It behaved as though no fluid occupied 

this area. At this point, it was decided to proceed with testing over the 

temperature range -4 to 149'F (-2OOC to 65OC). 

5.3.2 Performance Testing Over the Temperature Range -4 to 149'F (-2OOC to 

The honeycomb panel heat pipe was installed in the test chamber (Figures 20 

and 211, and an initial checkout was performed. This was to verify results 

consistent with bench testing. 

The cold test was performed first. First only the lower cold wall was turned 

on. With 100 watts input to the panel, the temperature was lowered 

approximately 27'F (15OC) in 1 hour. When the lower cold wall was turned off 

and the upper one turned on, the panel was cooled an additional 39.6 F (22OC) 

in one-half hour. This translates into a cooling rate of three times higher 

than that observed with the lower cold wall. This indicates that natural 

convection coupling between the panel and the upper cold wall was significantly higher than for the lower cold wall, as would be expected. The steady 

state temperature results are tabulated in Table 5, and the corresponding 

strip chart recording is given in Figure 26. When the power was increased to 

120 watts, a dry-out was observed. 

0 

For elevated temperature operation, the upper and lower cold walls were not 

used. The interior air temperature of the chamber was heated by ducting the 

exhaust air from an environmental chamber into the test chamber. A small fan 

was provided inside to distribute the air uniformly. Steady state temperatures for a nominal temperature of 149'F (65OC) and a power of 500 watts are 

tabulated in Table 6. The corresponding strip chart recording is shown in 

Figure 27. A dry-out was observed at 550 watts. 
TABLE 5 

HONEYCOMB PANEL TEMPERATURE 

DISTRIBUTION AT 100 WATTS (COLD TEST) 

Thermocouple 

Thermocouple 

No.* 

Temperature The rmocoup 1 e 

1 

2 

3 

4 

5 

6 

7 

8 

9 

10 

11 

12 

13 

14 

15 

No. 

1 

2 

3 

4 

5 

6 

7 

8 

9 

10 

11 

12 

13 

14 

15 

Temperature 

OF 

149.5 

149.9 

156.4 

149.5 

149.0 

148.1 

149.5 

149.5 

149.5 

149.5 

149.4 

149.2 

149.5 

149.4 

149.4 

F 0 

4.6 

8.6 

11.8 

8.8 

8.2 

4.1 

6.1 

4.8 

7.7 

9.5 

11.6 

9.5 

11.3 

8.1 

8.6 

OC 

-15.2 

-13 .O 

-11.2 

-12.9 

-13.2 

-15.5 

-14.6 

-15.1 

-13.5 

-12.5 

-11.3 

-12.5 

-11.5 

-13.3 

-13.0 

The rmocoup le 

No.* 

16 

17 

18 

19 

20 

21 

22 

23 

24 

25 

26 

27 

28 

29 

30 

Temperature 

OF 

5.2 

8.6 

5.0 

8.6 

12.2 

9.0 

8.2 

5.9 

Open 

11.4 

9.5 

11.1 

9.3 

8.4 

6.3 

*See Figure 22 for thermocouple locations. 

TABLE 6 

HONEYCOMB PANEL TEMPERATURE 

DISTRIBUTION AT 500 WATTS (ELEVATED TEMPERATURE) 

* NO 

65.3 

65.5 

69.1 

65.3 

65.0 

64.5 

65.3 

65.3 

65.3 

65.2 

65.2 

65.1 

65.3 

65.2 

65.2 

16 

17 

18 

19 

20 

21 

22 

23 

24 

25 

26 

27 

28 

29 

30 

43 

OC ~- ~ 

-14.9 

-13.0 

-15.0 

-13.0 

-11 .o 

-12.8 

-13.2 

-14.5 

Open 

-11.4 

-12.5 

-11.6 

-12.6 

-13.1 

-14.3 

Tempe ra t ure 

OF 

103.6 

149.0 

148.8 

150.1 

149.7 

151 .O 

149.2 

149.0 

150.1 

149.5 

149.4 

149.2 

149.2 

148.6 

102.6 

OC 

39.8 

65.0 

64.9 

65.6 , 

65.4 

66.1 

65.1 

65.0 

64.6 

65.3 

65.2 

65.1 

65.1 

64.8 

39.2 
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0 

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U 

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t3 

5 

45 
5.3.3 Liquid Fill Test 

After completion of the low and high temperature performance testing the panel 

was removed from the test chamber and connected to the process station. An 

additional 50 grams of methanol were added to the panel in order to start the 

liquid fill test. This corresponds to a 10 percent increase in the liquid 

fill. 

The methanol was added to the panel in a manner identical to the initial processing (Section 5.21, except the panel was not evacuated prior to filling. 

Even though the methanol in the charging cylinder was degassed prior to 

filling, the panel was also subjected to the degassing procedure described in 

Section 5.2 to provide assurance against the possibility of introducing gas 

during the transfer of methanol. 

The honeycomb test panel was then removed from the process station and 

installed in the test chamber. Upon initial application of 100 watts to the 

heater, a dry-out condition was observed for the central portion of the panel. 

This is an indication that the working fluid does not readily redistribute 

itself inside the panel, and there may be excessive welding of the core ribbon 

to the panel facesheets. 

Figure 28 shows the results for the 110 percent fill test. Although there was 

still a local area near the center of the heater that was 8.1 F (4.5OC) warmer 

than the remainder of the active panel surface, there was no evidence of 

excess liquid. For this reason, it was decided to increase the methanol fill 

by 10 percent again. 

0 

This time the panel was tested on the bench in order to facilitate the detection of excess liquid by tilting in various directions. When the methanol 

fill was increased to 120 percent, the performance was essentially the same as 

at 100 percent and 110 percent; i.e. no evidence of excess liquid at 400 watts 

heat input. This shows that the panel was relatively insensitive to an 

underfill. The fill was then increased by 10 percent again to 130 percent. 

46 
U 

m 

3 

I I I I 

0 

0 

U 

I I 

I , I I 

U 

Q 

47 
48 

At this time it was decided to perform a maximum power test to determine if 

the additional fill had any effect on the heat transport capacity. As shown 

by the results in Figure 29, the maximum power increased to 600 watts with 

only a - +2.7'F (+1.SoC) - AT over the entire active surface of the panel. At 

700 watts the AT increased to 19.8'F (ll°C) and this was considered to be a 

dryout. However, the panel recovered when the power was reduced to 400 watts. 

Note that there is evidence of excess liquid in the corner 6, 24 in Figure 29. 

In order to check this, the opposite corner (7, 23) was lowered, as indicated 

in Figure 30. Thermocouple No. 6, which is on the upper surface, responded 

instantaneously and TC No. 24 on the bottom surface slowly recovered also. 

This test was not conclusive, however because this corner (6, 24) had been 

known to trap liquid at the lower fills also (see Figure 25). But in this 

case TC Nos. 17 and 29 slowly started to decrease in temperature with time as 

shown in Figure 31. When this corner (17, 29) was tilted upward there was a 

rapid recovery, indicating the presence of excess liquid in the panel. 

It was concluded from these tests that a fill of 125 percent represents the 

optimum fill for this panel. 

or 0.07 lbm of methanol per square foot (355g/m 1. 

This amounts to approximately 1.43 lbm (0.65 kg) 

2 

5.3.4 Tilt Test 

Next the heater was removed from the centerline of the panel and placed along 

one edge in order to perform the tilt test. The heater was bonded to the top 

surface just inboard of thermocouple Nos. 6, 10, 12 and 16 (see Figure 22). 

The purpose of this test was to demonstrate that the panel does operate properly even against a positive gravity loading. One side of the panel was 

always gravity aided, with the heater in the center, regardless of which 

direction the panel was tilted. Note, however, that a factor of four reduction in maximum heat transport is to be expected in this configuration. This 

is because all of the heat input must be transported over the full width of 

the panel. When the heater was in the center, only one-half of the heat load 
89 OS ZE 

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51 
was required to be transported over one-half of the panel. Therefore, with no 

tilt and the heater located on the edge of the panel, maximum heat transport 

should be 150 watts. 

This test was performed with the panel outside the test chamber using the room 

temperature ambient air as the heat sink. The tilt test results are plotted 

in Figure 32. In all cases, thermocouple Nos. 10 and 12 were the first to 

indicate dryout, followed by thermocouple Nos. 26 and 28. The lower data 

points represent steady state heat transport capability over a period of at 

least one hour, whereas the upper points indicate a temperature difference of 

greater than 18.0°F (lO°C) between any two temperatures on the panel; i.e., 

dryout. 

Although the maximum heat transport is lower than expected, these results 

verify the ability of the panel wicking system to successfully perform against 

an adverse gravity field. Because the curve "levels off" at low values of 

tilt, however, it is concluded that the reduced performance is due to a higher 

than expected vapor pressure drop. This will be discussed further under Data 

Correlation and Conclusions, in Section 6.0. 

5.3.5 Burst Pressure Test 

The burst pressure test was performed on a subscale 5 112 inch by 5 112 inch 

(0.14 x 0.14 m) honeycomb panel. This sample was identical in materials and 

construction to the large thermal performance test panel. It was also 

constructed at the same time as the large panel. The test sample was helium 

leak checked before pressurizing. 

Figure 338 shows the five locations where the sample thickness was measured. 

(Note that location No. 5 is not at the center, but at a point conveniently 

measured with technician's calipers.) First the thickness was measured in the 

free unpressurized state and recorded on the data sheet. The panel was then 

placed in a safety chamber and pressurized with ultra-pure nitrogen in 50 psig 

increments. After holding the pressure for ten minutes, the panel thickness 

was measured and recorded. The test data are summarized in Table 7. Between 
/ 

a 

a 

W + 

W 

I 

53 
PRESSURE PORT / 

G14267 

TYP. 

t 

2+ \-\i 

A. BURST PRESSURE TEST SAMPLE DIAGRAM SHOWING 

MEASUREMENT LOCATIONS. 

E4945 

B. BURST PRESSURE TEST RESULT 

Figure 33 Burst pressure test. 
TABLE 7 

SUMMARY OF BURST PRESSURE DATA 

Thickness (in. ) 

Pres sure 

(psig) 

0 

50 

100 

150 

200 

0 

250 

1 

0.311 

0.312 

0.311 

0.313 

0.316 

0.315 

-- 

2 

0.310 

0.307 

0.309 

0.310 

0.315 

0.312 

-- 

3 

~ 

0.311 

0.311 

0.312 

0.314 

0.318 

0.316 

-- 

4 

0.311 

0.312 

0.312 

0.314 

0.321 

0.320 

-- 

5 

0.314 

0.312 

0.314 

0.317 

0.322 

0.319 

-- 

Cormnents 

With pressure 

With pressure 

With pressure 

With pressure 

Unit failed 

each pressure point the sample was removed from the test chamber and helium 

leak checked to verify continued structural integrity. 

At 200 psig a maximum deflection of .01 inch (0.25 urn) was observed. This 

deflection relaxed to only .009 inch (0.23 nrm) when the pressure was vented to 

atmosphere, indicating permanent deformation. The sample was repressurized to 

250 psig where, after approximately five minutes, the panel failed (Figure 33B). 

through the fill tube, revealed that the sintered wire core material had 

failed adjacent to the spot welds. Upon failure of the core material the unit 

expanded, like a "pillow". A leak check revealed two small leaks; one in the 

seam weld and one adjacent to the seam weld in a corner.' 

Subsequent examination of the panel, with a borescope inserted 

55 
6.0 DATA CORRELATION AND CONCLUSIONS 

6.1 Data Correlation 

Subsequent to thermal performance testing, refinements were made in the honeycomb heat pipe thermal modeling, and software was created for the IBM PC-XT 

computer. 

to reflect as-built panel parameters described in Section 4.0. 

performance models for the honeycomb heat pipe were discussed in Section 3.0. 

Theoretical predictions aided the selection of design parameters to meet performance requirements. 

parameters differed between the as-designed and the as-built hardware. 

Table 8 compares preliminary design to actual parameters and summarizes their 

influence on predicted heat pipe transport capability. The thermal model was 

therefore upgraded, not only to incorporate actual parameters, but also to 

more accurately describe the complex physical geometry of the honeycomb panel 

and to expand its prediction capability. 

The thermal transport model for the honeycomb fin was also upgraded 

Basic analytic 

However, as previously mentioned, several important 

Unit Cell Approach 

A typical honeycomb panel consists of individual and repeatable cells. 

liquid flow resistance of one such cell can be estimated by defining its 

physical constituents and using parallel and series flow path modeling. 

the analysis, the approximately hexagonal cell shape is assumed to be square. 

The basic honeycomb cell has an "effective" flow resistance parameter 

described by: 

The 

For 

- L.l , defined as C1, 

AwlKl 

where: 

L1 = length (in liquid mass flow direction) 

K1 = permeability 

Awl = WIH; W1 = width, H = height 

57 
Par ame t e r 

Vapor flow area 

Vapor path length 

(along one 

"channel" 

Spotweld spacing: 

Mini flange to 

f aceshee t 

Cell interfaces 

Resulting in: 

Unit cell flow 

resistance 

parameter (C1) 

Fluid tortuosity 

factors 

Effective capillary 

pore radius 

TABLE 8 

SUMMARY OF ANALYSIS UPGRADING 

Value 

Pre-Fab. 

2 0.012 in 

(7.92 mm2) 

12 in 

(0.30 m) 

25 

25 

Actual 

2 0.0031 in 

2 (1.98 mm 

16.9 in 

(0.43 m) 

Percent Open : 

%lo 

100 

13 m-3 3.58 x 10 

TF, = 1.0 

TFII = 1.0 

= 23 x loe6 m 

(165 x 1400 

core wick) 

'C 

13 m-3 4.95 x 10 

1.0 >TF, > 0 

1.0 ZTF~ > o 

23 x m <rc 

584.9 x m 

(120 x 120 

faces hee t 

wick) 

58 

Effect 

Decrease Q Sonic Decrease Q Entrainment 

Increase APv; 

Decrease Q Wicking 

Increase APv: 

Decrease QWicking 

Increase AP ; 

Decrease QWicking II 

Decrease AP, ; 

L Increase Q Wicking 

Increase AP,, 

Increase AP2; 

Both decrease 

QW i c king 

Reduce AP,; 

Decrease Q Wic king 
Using this "unit cell" parameter in a building block fashion (see Figure 341, 

the total flow resistance of a complete honeycomb panel can be calculated. 

Parallel to flow direction: 

n additional cells will multiply; 

C1n (added flow length) 

Perpendicular to flow direction: 

n additional cells will divide; 

(flow splitting) '1 /n 

G 14268 

PANEL 

MULTIPLE (n) 

CELLS IN X AND 

Y DIRECTIONS, 

1 CELL DEEP. 

3 

L 

0 

Figure 34 Honeycomb panel shown as comprised of "unit cells". 

59 
= c (A)($) L1 

The unit cell approach was introduced into the wicking transport equations of 

the computer model. It permits easy interchange of variable panel sizes, 

evaporator and condenser locations, and direction of heat flow. 

Tortuosity Factors 

The effect of fluid flow “tortuosity“ in its complex travel path through each 

honeycomb cell has not been fully included in the analytic model. 

sudden fluid flow bending, enlargement, contraction, etc., will add to the 

total pressure drop and thus act to reduce the wicking transport capability. 

The computer model permits independent selection of liquid (TFQ) or vapor 

(TFV) tortuosity factors. Values ranging from unity (no flow obstruction) to 

zero (complete flow obstruction) can be used to correlate localized pressure 

drop magnification (I@ A PI: 

In reality, 

TF = l/MFAP 

Composite Wick Considerations 

The honeycomb core consists of a composite wick structure, i.e., the Eacesheet 

and core ribbon wicks have different pumping pore radii (r and permeability 

(K) characteristics. Specific r and K values have been experimentally 

determined for a previous honeycomb panel (Ref. 10) having identical composite 

wicks as used in this panel. Although the analytic model accounts for the 

varied permeability between alternate liquid flow paths, it is not certain 

C 

C 

60 
which pore radius governs the capillary pumping. Since the heat input is into 

the facesheet, it seems that this is the pore most likely to govern pumping. 

An upgraded performance prediction for the -4 to 149'F (-20 to 65OC) temperature range is shown in Figure 35. Approximately a four to one variance in the 

capillary limit is possible as a result of selecting the high or low pore 

radius. Performance curves are also shown which assume independent existence 

of facesheet wick only or of core ribbon wick only; these are useful in 

establishing absolute minimum thermal transport capacities. 

Table 9 shows pressure drops predicted by analytic performance model for three 

temperature levels. The relative significance of vapor pressure drop varies 

significantly over the radiator operating temperature range; from being the 

largest contributor at low temperatures to relatively insignificant at higher 

temperatures. Maximum thermal transport results from the various tests are 

shown as a function of fluid temperature in Figure 36. The performance uses 

the conservative pumping pore radii-case 2 (Ref. Figure 35) and this is plotted as baseline. The independent influences of liquid and vapor tortuosity 

factors are also plotted. At the 75 percent level (implying a localized pressure drop increase of 33 percent), the vapor tortuosity factor correlates to 

test data at low tempeatures, while the liquid tortuosity conservatively correlates at higher temperatures. 

Tilt test results are shown as a function of fluid temperature in Figure 37. 

Test results correlate to a combined (vapor plus liquid) tortuosity factor of 

about 0.8 for the larger tilt heights. However, the actual performance levels 

off at low values of tilt. This may be explained by vapor flow effect, not 

included in the model. 

61 
t 

0 

@ 

@ 

I A - 

I- 120" -I g 

/ TFQ=TF,=l.O 

23 10-6 4.95 10~3 COMPOSITE. OPTOMISTIC 

84.9 10-6 4.95 8 1013 COMPOSITE. CONSERVATIVE 

84.9 I 6.54 x lot3 FACE SHEET ONLY 

I @ 1231 I 9.3. 1014 I CORE RlmONONLV 

10 1 I 1 I I I I 1 

-4 32 68 104 140 176 212 

TEMPERATURE (OF) I I I I I I 

-20 0 20 40 60 80 100 

TEMPERATURE (OC) 

Figure 35 Predicted perfornance limits upgrade for stainless steel/methanol honeycomb heat pipe. 

62 
GI4270 

lo3 c 

- 

*--- 

TFp = 0.75 

-' / 

CAPILLARY @ 

TFp TFV 1.0 '/ 

Le=l.OO" Lc=11.5" (EACH SIDE) 

rc (FACESHEET WICK) = 134.9 x ( + ) ciLL = 4.95 x 10 13 m- 3 

- THEORETICAL PERFORMANCE 

BENCHPERF. LOCAL DRYOUT 

POWER HELD (100% FILL) 

FILL TESTS LOCAL DRYOUT 

(130% FILL) POWER HELD 

TESTS I 

I I I I 1 I 

32 68 104 140 176 2 

101 I 

(OF) 4 

FLUID 

2 

Figure 36 Data correlation : thermal transport limitat ion. 

63 
G14271 

180 

160 - 

v) 

c 

c 

- 140 

X 

s 

U 

= 

v) 

e- 120 

2 

z 

a 

a 100 

Ic 

a 

w 

I 

80 f 

260 

E 

X 

40 

20 

0 

HEATER 

---,- THEORETICAL PERFORMANCE - 

LOCAL DRYOUT 

POWER HELD 

TILT TESTS 

(130% FILL) 

PE RAT" R E : 75.2-80.6' F (24-27'C 1 

- TFp TF, 5 0.8 

0 0.4 0.8 1.2 

TILT HEIGHT (IN) 

I I I 1 

8 10 20 30 

(MM 1 

Figure 37 Data correlation: tilt test performance. 

64 
TABLE 9 

SUMMARY OF PREDICTED PRESSURE LOSSES 

AP, N/m2 

50.6 

253.0 

287.0 

590.6 

LOSS TERM 

%TOTAL AP, NIm2 %TOTAL AP, NIm2 %TOTAL 

8.6 48.3 9.5 45.9 10.7 

42.8 357.0 70.5 346.0 81.1 

48.6 101.0 20.0 34.8 8.2 

100.0 506.3 100.0. 426.7 100.0 

GRAVITY 

HONEYCOMB 

CORE LIQUID 

HONEYCOMB 

CORE VAPOR 

TOTAL 

G14272 

r=3o.z0F I Q=246W IT=IOOAOF( Q=557W 1T=170.6°F1 Q=780W 

6.2 Conclusions 

The major objective of this honeycomb heat pipe development program was to 

demonstrate performance for the space radiator fin application. The feasibility of fabricating and processing a stainless steel heat pipe in a honeycomb configuration for the fin application has been successfully established. 

A heat transport capacity of 600 watts at 122'F (5OoC) or 60 watts per foot 

(200 watts per meter) of fin length was demonstrated. 

design goal of 1,000 watts, primarily because the vapor holes were punched by 

the vendor only in every other crimp of the core wick material, rather than 

every crimp as originally specified (see Figures 15 and 18). 

This falls short of the 

A close look at the test panel interior reveals constraints placed on its performance and helps explain test results. 

which shows the several possible vapor flows resulting from alternate placement of holes in the core ribbon. With holes in every honeycomb cell face, as 

originally designed, the vapor can communicate between all cells and can work 

This is illustrated in Figure 38 

65 
its way from any evaporator location to a point directly opposite it on a 

condenser edge (see Figure 38a). As mentioned previously, the vapor flow area 

and length were adversely affected due to the inclusion of vapor holes in 

every other crimp only, rather than in every crimp (Ref. Figure 18). If the 

holes of adjacent core ribbon strips are in phase as in Figure 38b, the vapor 

will flow diagonally across the entire panel. If however, the adjacent strip 

of core material is shifted one cell as shown in Figure 38c, vapor will be 

forced to take a 90' turn in a following cell, depending on how adjacent 

strips are lined up. 

Holes punched into the core ribbon constrain the vapor to flow along paths 

which can be viewed as "channels." The test panel actually consists of three 

sections which differ in arrangement of those vapor channels. The left and 

right sections contain single direction channels, but are oriented 90' out of 

phase with each 'other. The middle section is comprised of a mix of channel 

sections creating very complicated vapor flow. 

material is permeable to vapor - if given sufficient time to pass through - it 

does severely diminish its ability to freely and rapidly communicate with all 

areas of the panel. As a result, the full radiator area is not utilized 

effectively, as shown in the detail schematic of the middle core section 

(Figure 39). The areas above the top shaded channel (for center heater) and 

the area above the bottom shaded channel (for edge heater) are cut off from 

"free" vapor flow and therefore become non utilized or poorly utilized 

sections of the heat pipe condenser. Similarly, the panel corners at TCs 16, 

30 and 6, 24 are isolated from vapor by continuous channel walls - if the 

heater is centrally located as during baseline, performance, and fill tests. 

Since these corners are thermally isolated, they become subcooled relative to 

the rest of the panel; differences in vapor pressure will thus cause liquid to 

fill them. Under conditions that create cold corners, excess working fluid 

will be required to prevent premature panel dry-out. When the heater was 

relocated to the panel edge during tilt testing, these cold corners did not 

reappear. 

Although the core ribbon wick 

Severe localized effects are possible within fluid channels of the middle 

panel section. These are: the same channel can cross under the heater several 

times and thus receive multiple heat loads; vapor pressure drops will increase 

66 
GI4273 

mw 

A. ORIGINALLY DESIGNED 

B. AS BUILT; IN-PHASE SECTION 

C. AS BUILT: OUT-OF-PHASE SECTION 

SINGLE 

CORE RIBBON w SPOT WELD 

ACTUAL CELL SHAPE 

RIBBON JOINING SHOWN 1 

(OUT-OF-PHASE CORE 

NOTE: TOP PANEL VIEW SHOWN 

A 

TRANSVERSE 

FLOW 

Figure 38 Summary of cell hole pattern effect on vapor flow direction. 

67 
e 

c 

J 

v) 

I 

i2 

I3 

2 

L 

LL 

z 

4 

> 

Io 

W 

-I 

W 

v) 

z 

I, 

v) 

w 

z 

z 

o 

w 

a 

r 

n 

n 

a 

I, 

.. 

W 

k0 

z 

W 

cu 

cu- -F 

c 

v) 

d 

a, 

C 

C 

(d 

c 

U 

v) 

C 

0 

4 

U 

U 

a, 

&I 

.r( 

a 

a 

-I4 

2 

E 

brJ 

C 

rl 

> 

(d 

c 

C 

0 

4 

U 

U 

a 

v) 

rl 

a 

C 

(d 

a 

%I 

0 

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7.1 

> 

d 

rl 

(d 

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a, 

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m 

Q) 

&I 

3 

M 

.r( 

lk 

68 
due to numerous changes in flow direction and from longer flow paths; and 

entrainment and sonic Limits will be depressed by effective reductions in 

vapor flow area. In most cases, initial dry-out was observed at evaporator 

thermocouples located nearest the troublesome middle panel section (TC for 

central heater and TC's 10,12 for edge heater); the region where multiple heat 

loading exists. 

The current analytic model incorporates single-directional channel characteristics shown in Figure 38b only, in other words, the complications of the 

middle channel section are not included. Although greater prediction accuracy 

can.be developed, it would be a time-consuming task. The preferred approach 

is to ascertain that future honeycomb panels have adequate vapor communications and thus utilize the existing correlated model to predict performance. 

This is done for the two additional cases, as shown in Figure 40. The upper 

curve shows the increase in as-designed panel performance, which incorporates 

a four-fold increase in vapor flow area and 50 percent open liquid flow area 

at facesheet spotwelds. Leveling of the curve is a result of the relative 

significance of vapor as compared to liquid pressure drops at lower temperatures. However, to meet the design goal of 1000 watts over the entire operating temperature range shown in the upper curve, it is necessary to incorporate the lower flow resistance of honeycomb channels. Both cases use 

combined tortuosity factors of 0.8, which is very conservative for the channel 

design. 

The conclusions based on results of this development program are summarized 

below: 

Acceptable fabrication and processing techniques of stainless 

steel/methanol heat pipes in a honeycomb configuration for the space 

radiator fin application have been developed. 

0 A heat transport capacity of 600 watts or 60 watts per foot 

(197 watts/meter) of fin length was demonstrated. This falls short 

of the 1000 watt design goal, primarily because vapor holes were 

punched in every other crimp of the core ribbon only, rather than 

every crimp as originally designed. 

69 
due to numerous changes in flow direction and from longer flow paths; and 

entrainment and sonic limits will be depressed by effective reductions in vapor 

flow area. In most cases, initial dry-out was observed at evaporator thermocouples located nearest the troublesome middle panel section (TC for central 

heater and TC's 10,12 for edge heater); the region where multiple heat loading 

exists. 

The current analytic model incorporates single-directional channel characteristics shown in Figure 38b only, in other words, the complications of the 

middle channel section are not included. Although greater prediction accuracy 

can be developed, it would be a time-consuming task. The preferred appr6ach 

is to ascertain that future honeycomb panels have adequate vapor communications and thus utilize the existing correlated model to predict performance. 

This is done for the two additional cases, as shown in Figure 40. The upper 

curve shows the increase in as-designed panel performance, which incorporates 

a four-fold increase in vapor flow area and 50 percent open liquid flow area 

at facesheet spotwelds. Leveling of the curve is a result of the relative 

significance of vapor as compared to liquid pressure drops at lower temperatures. However, to meet the design goal of 1000 watts over the entire operating temperature range shown in the upper curve, it is necessary to incorporate the lower flow resistance of honeycomb channels. Both cases use 

combined tortuosity factors of 0.8, which is very conservative for the channel 

design. 

below: 

0 

0 

The conclusions based on results of this development program are summarized 

Acceptable fabrication and processing techniques of stainless 

steel/methanol heat pipes in a honeycomb configuration for the space 

radiator fin application have been developed. 

A heat transport capacity of 600 watts or 60 watts per foot 

(197 watts/meter) of fin length was demonstrated. This falls short 

of the 1000 watt design goal, primarily because vapor holes were 

punched in every other crimp of the core ribbon only, rather than 

every crimp as originally designed. 

70 
L p- 120" -4 

TFp = TF, = 0.8 rc * 84.9 x 106M 

*HONEYCOMB CELLS (.5" SIZE) 

- L 13 3 ~3.22~10 M' 

L **HONEYCOMB CHANNELS (.5" SPACING) 

13 3 =1.54x10 M- .. 

I I 1 I I 

32 68 104 140 176 212 (OF) -4 

FLUID 

TEMPERATURE 

I I I I I I I 

(OC) -20 0 20 40 60 80 100 

Figure 40 Predictions for as-designed and increased 

capacity designed using correlated model. 
0 A temperature differential of - +2.7'F (1.5OC) (at 600 watts) over the 

entire active surface of the honeycomb panel was demonstrated. The 

existence of two small cold corners is due to vapor isolation created 

by core ribbons without vapor holes. 

0 The inclusion of vapor holes in every core ribbon crimp will ensure 

the necessary vapor communication throughout all areas of the 

honeycomb panel and will result in the panel exceeding all 

performance requirements. 

0 The correlated analytic model can be used with subscale panel segment 

tests to predict performance of future panel hardware. As-designed 

panel performance is predicted to be from 500 to 1000 watts over the 

operating temperature ranges; an open channel design is predicted to 

achieve greater than 1000 watts over the entire temperature range. 

0 The panel can operate over a wide temperature range without liquid 

blockage effects due to liquid expansion. The honeycomb core 

material inherently provides excess liquid reservoir capacity 

distributed throughout the panel. 

The technology and commercial equipment are currently available to construct 

all-welded, machine-assembled heat pipe panels using stainless steel and other 

ferrous materials, nickel alloys, titanium, and titanium alloys. Aluminum 

panels will require further development of these techniques and/or alternative 

methods of fabrication. 

An improved lightweight heat pipe radiator fin will evolve from additional 

efforts to reduce weight and maintain the panel within structural allowables. 

Fabrication processes for lightweight aluminum, low density wickable core 

honeycomb panels require further evaluation. Channel and strut designs could 

provide improved performance. Modular design trade-offs must be evaluated to 

essentially isolate panel compartments, and provide high reliability to micro 

meteroroid and space debris damage. 

72 
REFE RE NCES 

1. 

2. 

3. 

4. 

5. 

6. 

7. 

8. 

9. 

10. 

11. 

Ellis, W.E., and Rankin, J.G., "Heat Rejection Development Requirements 

for Orbital Power Systems", Paper No. 799295, 14th Intersociety Energy 

Conversion Engineering Conference, August 1979. 

Oren, J.A., "Study of Thermal Management for Space Platform 

Applications, NASA CR-165238, December 1980. 

Oren, J.A., "Study of Thermal Management for Space Platform Applications: 

Unmanned Module Thermal Management and Radiator Technologies," NASA CR165307, May 1981. 

Rankin, J.G., and Marshall, P.F., "Thermal Management System Technology 

Development for Space Station Applications," Paper No. 831097, 13th 

Intersociety Conference on Environmental Systems, July 1983. 

Basiulis, A., and Camarda, C.J., "Design, Fabrication and Test of Liquid 

Metal Heat-Pipe Sandwich Panels," AIAA Paper No. 82-0903, AIAAIASME 3rd 

Joint Thermophysics, Fluids, Plasma & Heat Transfer Conference, June 1982. 

Tanzer, H.J., "High Capacity Honeycomb Panel Heat Pipes for Space 

Radiators," AIAA Paper No. 83-1430, June 1983. 

Alario, J., Haslett, R., and Kosson, R., "The Monogroove High Performance 

Heat Pipe", AIAA Publication 81-1156, June 1981. 

Tanzer, H.J., "Fabrication and Development of Several Heat Pipe Honeycomb 

Sandwich Panel Concepts", NASA CR 165962, June 1982. 

Fleischman, G.L., Scollon, T.R., and Loose, J.D., "Vapor Chambers for an 

Atmospheric Cloud Physics Laboratory," AIAA Paper No. 82-0903, July 1980. 

Camarda, C.J., and Basiulis, A., "Radiant Heating Tests of Several Liquid 

Metal Heat-Pipe Sandwich Panels," AIAA Paper No. 83-0319, January 1983. 

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