HOT AIR BALLOON PROGRAM PROGRESS REPORT CONTRACT NONR 1598 (05) - TASK "B"

• Document Type: CREST
• Collection: General CIA Records
• Document Number (FOIA) /ESDN (CREST): CIA-RDP78-03642A002500050001-8
• Release Decision: RIPPUB
• Original Classification: S
• Document Page Count: 67
• Document Creation Date: December 27, 2016
• Document Release Date: May 25, 2011
• Sequence Number: 1
• Publication Date: January 4, 1957
• Content Type: REPORT

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Engineering Research and Development Depart-m-
MIIN~MAPOLIS, MINNESOTA
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.ii%.Ri
Mechanical Division of
GENERAL MILLS, INC.
Engineering, Research & Development
2003 East Hennepin Avenue
Minneapolis 13, Minnesota
'c1, . , ly
ments.
HOT AIR BALLOON PROGRAM
Progress Report
Prepared by
W. C. Borgeson, H. E. Henjum, D. N. Rittenhouse,
V. H. Stone and G. R. Whitnah
Contract Nonr 1589(05) - Task "B"
Office of Naval Research
Department of the Navy
Washington 25, D. C.
4
Submitted byg `/ dla~
H. E. Froehlich
Approved by
,/ ~ Head, Geophysics Section
Otmar M. Stuetzer, Manager
Physics & Chemistry Researc
Report No. 1647
Date : January J4.
Project ~"Tj~ -1t~ ---55M11
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TABLE OF CONTENTS
I. INTRODUCTION
II. LIFT OF HOT AIR BALLOONS
III. FUELS AND COMBIJSTION
PMe
1
A.
B.
C.
Fuel Studies
Molecular Weight and Dew Point of Products
Combustion Efficiency
TV. HEAT LOSS
A. Analysis
B. Experiments
V. FUEL STORAGE
tTI. INFLATION
36
A. General Analysis
36
B. Experimental Work
39
VII. PROTOTYPE BURNER ASSEMBLY
!,4
A.
Inflation Fan and Motor
44
B.
Fuel Tank (Propane)
47
C.
Ignition System
47
D.
Main R.irner
47
E.
F.
Tank Heater
Relief and Pressure Regulating Valve
V111. CONTROL OF HOT AIR BALLOf^I^
A. Valving
ti. Reversing Inflation F?n
C. Modulating Fuel Input
IX. CONCLI JS TONS
X. RFFFRENCES
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LIST OF ILLUSTRATIONS
Figure
Page
2.1
Lift Per Unit Volume Vs Various Temperatures
4
2.2
Altitude Correction Curve
5
2.3
Required Internal Temperature
6
3.1
Molecular Weight of Combustion Products
3.2
Dew Point of Combustion Products
3.3
Excess Air Vs CO2 Concentration
16
3.4
Ventilation Heat Losses
16
4.1
Balloon Air Temperature Required for Constant Lift
20
4.2
Total Heat Required to Heat Initial Inflation Air and
Maintain Temperatures of Figure 4.1 During Expansion
on Ascent
20
4.3
Heat Loss for Single-Wall Balloon at Temperatures of
Figure 4.1
22
4.4
Heat Loss for Double-Wall Balloon at Temperatures of
Figure 4.1
22
4.5
Heat Loss for Triple-Wall Balloon at Temperatures of
Figure 4.1
22
4.6
Wall Temperatures for Double-Wall Balloon at Internal
Temperatures of Figure 4.1
23
4.7
Wall Temperatures for Triple-Wall Balloon at Internal
Temperatures of Figure 4.1
23
4.8
Test Configuration for Lift Measurements
26
4.9
Balloon Temperature Distributions
27
5.1
Weight of Spherical Tanks with Safety Factor of 2.0
34
5.2
Efficiency of Spherical Tanks
34
5.3
Energy Storage in Spherical Fuel Tanks
35
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Figure
LIST OF ILLUSTRATIONS (CONT.)
Page
6.1
Inflation Fan Test Configuration
40
6.2
Battery Drainage Tests - Six Batteries (Burgess F4BP)
Open Circuit Voltage, 18V
42
6.3
Battery Drainage Tests - Six Batteries (Burgess F4BP)
Open Circuit Voltage, 12V
42
6.4
Battery Drainage Tests - Four Batteries (Burgess ABP)
Open Circuit Voltage, 12V
43
7.1
Burner Assembly - Top View
45
7.2
Burner Assembly - Bottom View
46
7.3
Burner Controls with Ignition Control Box
48
7.4
Burner Ignition Control Box
49
7.5
Tank Heater Burner Assembly with Baffle Removed
51
7.6
Outline Drawing of Prototype Assembl
y
53
7.7
Circuit Diagram - Electric Ignition System
54
8.1
Lift Loss Rates from Circular Valves
Table
LIST OF TABLES
57
Page
3.1
Hydrocarbon Fuels
8
4.1
Summary of Unit Heat Loss Values - Experiments with
Hot Air Balloons 1
25
4.2
Results of Heat Loss Tests of 7-ft Single-Wall Balloon
28
4.3
Results of Heat Loss Tests of 7-ft Double-Wall Poly-
ethylene Balloon
28
4.4
Results of Heat Loss Tests of 7-ft Double-Wall Mylar
Balloon
30
6.1
Power Required for Inflation
38
6.2
Propellor Performance
39
7.1
Calibration of Main Burner
52
iv
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I. INTRODUCTION
Research and development work on hot air balloons is being carried out by
General Mills, Inc. under contract with the Office of Naval Research as one
phase of the program covering Low Altitude Controlled Flights.
The use of hot air as a lifting gas appears desirable for some low alti-
tude balloon applications due to the low weight and volume of the ground in-
flation equipment utilized. The inflation of balloons with hydrogen or helium
requires a supply of gas from high pressure cylinders or a field gas generator,
which are of considerable size and weight. In some cases, the application of
balloons to military tasks is limited by the requirement for inflation gas
and the associated shipping and handling problems. For example, the infla-
tion of a balloon with helium for a gross lift of 350 lb requires approximately
3,500 lb of steel gas cylinders. In comparison, results of work on the present
project indicate that the above load could be lifted to 5,000 ft and sustained
for two hours with a total equipment weight of only 150 lb. This weight in-
cludes the fuel, tank, burner, inflation fan and controls. Longer durations
could be attained by carrying more fuel and might also result from future
improvements in hot air balloon design.
Although the use of hot air balloons has been considered largely for
manned flight applications, non-manned, fully-automatic balloon systems may
find considerable application. The use of this vehicle in short-range delivery
systems could prove advantageous over systems that have been conceived in the
past.
The following sections of this report cover, in detail, the research
and development work which has been accomplished.
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II. LIFT OF HOT AIR BALLOONS
The lift obtained in a hot air balloon is a result of the decreased
density inside the envelope caused by elevated temperature. This lift is
calculable by the equations:
Lift ^V((Oa-
P-
Pa Ma
a R Ta
P - Pb Mb
b R Tb
Where V is-the balloon volume, ft3
ra is the density of ambient air, lb/ft3
Pb is the density of air inside the balloon
Pa is the ambient pressure, lb/ft2
Ma is the molecular weight of the ambient air
Ta is the temperature of the ambient air, degrees Rankine
R is the universal gas constant, 15+4 ft/?R
Pb is the pressure inside,the balloon, lb/ft2
Mb is the average molecular weight inside the balloon
Tb is the average temperature inside the balloon, ?R.
(2.2)
(2.3)
Since the differential pressure from the inside to the outside of a
balloon is extremely small compared with the absolute pressure, it is valid
to let Pa = Pb. It has been demonstrated (see Section III) that the average
molecular weight of the products of combustion is very nearly that of pure
air, thereby allowing the simplification of letting Ma _ Mb.
eb)
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The lift equation then becomes:
Lift = V Pa Ma 1 _ 1 (2.4)
R [Fa Tb
The unit lift of hot air (in lb/ft3) has been calculated from Equation
2.4 and is illustrated as a function of Ta and Tb in Figure 2.1 for the sea
level case. A correction factor for altitude variation is given in Figure
2.2.
An illustrative case has been taken from the above analysis, for which
the operating temperatures required in three different balloons of different
diameters have been plotted against gross lift at 10,000 ft MSL in the standard
atmosphere. This information is shown in Figure 2.30
Experimental measurements of lift obtained have been made during this
project and satisfactorily confirm the information of Figures 2.1, 2.2 and
2.3. This work is described in Section N.
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GENERAL MILLS INC., ENGINEERING RESEARCH a DEVELOPMENT DEPT., MINNEAPOLIS, MINNESOTA
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III. FUELS AND COMBUSTION
A. Fuel Studies
Maintenance of the required internal temperature in a hot air balloon
requires a continuous input of heat. The combustion of hydrocarbon fuels
with atmospheric air appears to be the most satisfactory method of supplying
this heat. Since the fuel must be carried on the balloon, thus reducing
the payload capability, the weight of fuel required is an important considera-
tion.
Table 3.1 lists four hydrocarbon fuels which cover the volatility range
considered applicable to a hot air balloon system. No. 1 fuel oil, which
is slightly less volatile than kerosene, has been taken as the lower limit
for volatility. The use of No. 2 or heavier grades would only increase the
complexity of the combustion equipment and also reduce reliability. Propane
has been taken as the most volatile hydrocarbon fuel which would be practical
to contain as a liquid in storage tanks. It has a vapor pressure of 286 psi
at 130?F. The saturated hydrocarbon of next higher volatility, ethane (C2H6),
has a critical temperature of 90?F and cannot be stored as a liquid at normal
summer temperatures.
It is seen from Table 3.1 that these hydrocarbon fuels are quite
similar in heating value. Heating value increases with hydrogen content,
but the differences are not large. All of the fuels of Table 3.1 are con-
sidered applicable to a hot air balloon system. The high volatility of propane,
which allows it to be vaporized at low temperatures in the storage tank,
makes it the most desirable fuel from the standpoint of burner design. The
boiling point of propane at atmospheric pressure is -31?F and that of Butane
is +150F. Gasoline and No. 1 fuel oil vaporize at higher temperatures and
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are not readily adaptable to systems where the fuel is vaporized in the
storage tank.
Propane requires a heat input for vaporization of slightly less than
one per cent of its heating value. In some cases, this heat is obtained by
natural heat transfer through the walls of the tank from the atmosphere. If
the demand for vaporized fuel is high and the surface area is small, additional
heat must be added to vaporize the liquid. In the propane burner described
in Section VII, a small burner provides heat for this purpose.
Storage tanks for these fuels are discussed in Section V. Figure 5.3
illustrates the weight of the tank and fuel required for various quantities
of energy storage.
TABLE 30l
HYDROCARBON FUELS
Per Cent
H~
Higher
Heating Value
Formula
Propane
18.2
21,560 BTU/lb
C3H8
Butane
17.25
21,180
C4HlO
Gasoline
15.8
20,500
C8H18
No. 1 Fuel Oil
14.0
19,750
B. Molecular Weight and Dew Point of Products
Consideration has been given to selection of the most advantageous
method of transferring the heat to the air inside the balloon. The two
alternatives are (1) allowing the products of combustion to enter the balloon
directly, or, (2) making use of a heat exchanger which transfers heat from
the products to the air in the balloon, thereby allowing discharge of the
products directly overboard. Further examination shows that the first of
-8-
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these alternatives is the most simple and results in the lowest equipment
weight without any penalty in lifting capability. Calculations have been
made of the average molecular weight of the products of combustion of hydro-
carbon fuels with various amounts of air. These point out that, if the steam
formed by combustion of the hydrogen is retained in the superheated state,
the total mixture of air, nitrogen, carbon dioxide and steam is slightly
buoyant with respect to air at the same temperature. The average molecular
weights resulting from these calculations are shown in Figure 3.1.
A sample calculation illustrating the basis of Figure 3.1 is given
The products' formed by the combustion of one pound of propane with
100 per cent excess air are:
Item Weight (lb)
Coe 3.00
H2O 1.64
N2 12.02
Air 15.65
Total 32031
The volumes occupied by each item at the arbitrary standard sea level
condition, with a temperature of 590F and a pressure of 2116 psf, are deter-
mined from the perfect gas law:
V: W 1544 T= W W44)(519) W
M P M 2116 = 379 M (3.1)
These volumes are:
for C02, V = 3-79(3-00) = 25.8 ft3
44-
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for H2O, V = 379(1.64)
= 134.5 ft3
for N2, V = 379 28.02 _ 162.5 ft3
for Air, v m 379 x 15.65 = 204,5 ft3
28.97
Total Volume - 427.3 ft3
Average Molecular Weight = MW 79272331
28.66
This value, which is lower than the molecular weight of air (28.97),
is valid only if the water vapor remains in the superheated form. The dew
point of the combustion products may be determined by a method similar to
that above, in which the water vapor specific volume is calculated, thereby
fixing the saturation temperature. For the case illustrated above, the water
vapor specific volume is 260 ft3/lb and the dew point is 111?F. Dew point
values for these fuels at different air mixtures are given in Figure 3.2.
Dew point values in Figure 3.2 fall below 100OF at 200 per cent excess air.
Calculations of balloon wall temperature in Section IV indicate that it will
be practical to maintain wall temperatures above 100oF, thus assuring freedom
from condensation inside the balloon.
C. Combustion Efficiency
Fuel supplied to the burner must release heat which will pass through
the balloon walls at a rate adequate to maintain the required temperature
within the balloon. There are two possibilities for inefficiency in this
process:
1. Incomplete combustion - If all of the hydrogen is not burned
to water and all the carbon is not burned to C02, a loss of energy is in-
volved.
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SECfE I
MOLECULAR WEIGHT OF COMBUSTION PRODUCTS
28.30
100
PERCENT EXCESS AIR
Figure 3.1
NX
80o
100
PERCENT EXCESS AIR
Figure 3.2
NO.I FUEL OIL
GASOLINE
PROPANE
GASOLINE
NO.I FUEL OIL
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2. Ventilation loss - If some of the heat released by the fuel is
not transferred through the balloon film but leaves in a stream of hot gases
passing through a ventilation port, a loss of energy is involved.
1. Incomplete Combustion
The hydrocarbon fuels considered in this work are propane, butane,
gasoline and No. 1 fuel oil. These fuels have all been used in industrial
equipment and burners have been developed in which complete 'combustion is
assured. Incomplete combustion results in production of carbon monoxide.
This situation cannot be tolerated in, for example, domestic heating equip-
ment. Precautions are therefore taken in the design of these burners to
eliminate the possibility of creating carbon monoxide. The primary consid-
eration is to provide adequate excess air for combustion. The amount re-
quired depends upon burner design and fuel properties.. Gaseous fuels, such
as propane and butane, require less excess, air than liquid fuels because they
are more easily mixed with air. Burners for gaseous fuels can operate with-
out danger of incomplete combustion with excess air quantities below 100 per
cent, and those for the liquid fuels can reliably produce complete combustion
with less than 150 per cent excess air. Satisfactory burners for the gaseous
fuels are more simple than those for liquid fuels because the fuel atomiza-
tion requirement is not present.
It is concluded from the above considerations that a burner for
application to a hot air balloon must be designed to produce complete com-
bustion. Operation in the zone of incomplete combustion would introduce
additional practical problems of luminous flames in gaseous burners and smoke
production with liquid fuels.
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2. Ventilation Loss
Continuous efficient operation of a burner requires that new air
be supplied and combustion products removed at an equal rate. The flow of
combustion products from a balloon at elevated temperature represents an in-
efficiency in the system. The amount of loss maybe determined by the equation2
below:
V=WgCp4T+9Wh(1089t0.4+55LT) (3.2)
V is the ventilation loss Btu/hr
Wg is the mass flow of combustion products, lb/hr
Cp is the specific heat of the products Btu/lb-?F
&T is the temperature difference from the inside to outside
of the balloon
Wh is the weight fraction of hydrogen in the fuel, lb/lb.
The quantity of heat leaving the system through the balloon wall is:
Q _ UA 6T
(3a.3)
Q is the heat flow, Btu/hr
U is the over-all coefficient of heat transfer, Btu/hr-ft2-OF
A is the balloon surface area, ft2
A T is the temperature difference from the inside to outside of
the balloon.
The quantity of heat released by the fuel is:
H = Wf AHf
H is the heat release, Btu/hr
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(3.4)
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Wf is the mass flow rate of fuel, lb/hr
&Hf is the higher heating value of the fuel, Btu/lb.
The energy balance for the system, with nomenclature as defined
above, is the following:
H % Q+y (3.5)
The combustion efficiency, based on zero carbon monoxide, is:
C g? H H V
(3.6)
In order to evaluate the items of Equations 3.2 through 3.6,
determination must be made of Wg, the mass flow rate of the combustion
products. In experimental work this is most conveniently accomplished by
volumetric analyses of the products of combustion with an Orsat Analyzer.
The measurement of C02, 02 and CO defines/'b Wg/Wf for a given fuel of known
composition:
- 11 Co2 + 8 02 - 7 (CO + N2 )
Wf - 3 CO 2 + CO
(3.7)
where C02, 02, CO and N2 designations are whole number percentages of con-
stituents from an Orsat analysis.
The excess air may be calculated from the composition of the
combustion products and that of the fuel by Equation M.
we
oa2l + 3 (wh wo J
Per cent excess air C02 - 0.125 )
[w+3 (wh- 00125wo1
we is the weight fraction of carbon in the fuel, lb/lb
wh is the weight fraction of hydrogen in the fuel, lb/lb
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(3.8)
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Ko is the weight location of oxygen in the fuel, lb/lb
CO2 is the whole number percentage of volume from the Orsat.
Orsat readings which correspond to excess air quantities from
zero to 200 per cent are given in Figure 3.3 for propane, gasoline and Noe 1
fuel oil. Combustion efficiencies, based on zero carbon monoxide, are shown
as a function of excess air and temperature differential in Figure 3.1 These
graphs allow convenient evaluation of performance from the Orsat analysis.
It is seen from Figure 3.4+ that combustion efficiency is rela-
tively high for all excess air values below 200 per cent. Therefore it appears
that it will be advantageous to operate with approximately 200 per cent excess
air to insure complete combustion and to reduce the dew point, since the
penalty is not severe.
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EXCESS AIR VS CO2 CONCENTRATION
02
71
\
'
NN
4 6 8 10 12 14 16
PERCENT CO2 BY VOLUME IN BALLOON
Figure 3.3
VENTILATION HEAT LOSSES
16
W
J
4
I
H (,
Z
2
aI
Z =
a $ 10
T
8
z
> 0
6
z
w
U
0=
0. 40
40 80 120
TEMPERATURE DIFFERENTIAL (?F)
Figure 3.4
SECRE~I.
CODE:
Propane
- - Gasoline
----- No. I Fuel Oil
i
...........
200% EXCESS AIR
} 100% EXCESS AIR
N0 EXCESS AIR
Ambient Air Temperature, 59?F
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IV- HEAT LOSS
A. Analysis
The total heat loss from a hot air balloon is the sum of the heat
flowing through the balloon film plus that leaving the balloon by ventila-
tion. The ventilation losses are relatively small (less than 20 per cent
of the fuel input) and are readily determined by the method presented in
Section III. The heat loss through the balloon film is of major significance.
Heat is transferred from the warm gases inside the balloon to the balloon
wall by a combination of the three heat transfer modes. radiation, con-
vection and conduction. It passes through the balloon material by conduction
and is again transferred to the surrounding air by a combination of all three
modes.
This situation is similar to the case of heat transfer between any two
gases separated by a thin membrane. It has been useful, in similar problems,
to define an over-all coefficient of heat transfer, U, having the engineering
units of Btu/hr?ft2-?F, since the heat flow in this type of problem has been
found to be linear with temperature differential:
U = A ~T
Q is the total heat flow, Btu/hr
A is the total area of the film, ft2
AT is the temperature differential between the two gases, OF0
Another concept that has been applied to this problem is that of total
thermal resistance, RT, which is the reciprocal of U. The total thermal
resistance in this case is considered to consist of that of a film of air
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on each side of the membrane plus the membrane itself. In equation form:
U = 1 1
- f + K + f
a b
(4+.2)
fa is the conductance of the first "film", Btu/hr-ft2-?F
fb is the conductance of the second "film", Btu/hr-ft2-?F
X is the thickness of the membrane, ft
K is the thermal conductivity of the membrane material,
Btu -ft/hr-ft2-?F
For thin membranes the K term drops out. This is the case with balloon
materials in the range of 1 to 10 mils thick. The computed value of the
conductance for a 2-mil polyethylene film is 1,160 Btu/hr-ft 2-?F, demon-
strating a negligible thermal resistance.
An important conclusion at this point is that the heat loss from a
single-membrane balloon is independent of the material. This fact has been
confirmed experimentally in that the heat loss from 7-ft diameter polyethylene
and Mylar balloons was found to be the same.
The heat transfer across the air "film" on each side of the membrane
is therefore the major consideration. Actual values of a or fb are known
to be strongly dependent upon air velocity and surface emissivity and to a
lesser degree upon air density and temperature difference. The effects of
the more important variables, air velocity and surface emissivity, have been
determined through work in connection with the heating of buildings. For
non-metallic materials, the infrared emissivity is approximately 0.90. The
film coefficients fa and fb have a value 2b of 1.65 Btu/hr-ft2-?F in this case.
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For a relative air motion of 5 mph, fa increases to 3.0 Btu/hr_ft2_oF for
normal temperature ranges. The corresponding values for a single-wall
balloon are 0082 and,1o06 Btu/hr_ft2_op. From this same work, the conductance
of a stagnant air space (between two membranes) is 1.18 Btu/hr_ft2_OF, near
room temperature.
The foregoing statements indicate a possible advantage in fabricating
a balloon out of several Payers of material, thereby creating air spaces of
significant thermal resistance. With a two-layer balloon having one air
space, the U value for zero wind is 0.485 Btu/hr_ft2_op, and, with a 5-Mph
relative wind, U equals 0.56 Btu/hr_ft2_oFa
Extending this analysis to a
three-layer balloon with two air spaces, the U values are 0034 and 0038
Btu/hr_ft2_op for the zero and 5-mph examples respectively.
Fabrication of multilayer balloons in which an effective air space is
achieved has not been accomplished to date. However, it is expected that
significant reductions in heat loss could be accomplished by this method.
Calculations of heat required have been made for a sample balloon of
27,000-cu ft volume, with a gross lift of 350 lb. The required internal
temperatures for this constant lift are shown as a function of altitude in
Figure 4.1. These values are based on the NACA standard atmosphere. The
equations used in this calculation are given in Section Ii.
Calculations have been made of the fixed quantity of heat required to
warm the initial inflation air to the operating temperature and to counter-
act the cooling tendency produced by the expansion during ascent. These
values are given in Figure 4.2o In evaluating the heating requirement caused
by the ascent, the standard dry adiabatic lapse rate has been used. The
values of Figure 4.2 represent 3 to 5 lb of fuel, an amount which is significant
19
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BALLOON AIR TEMPERATURE REQUIRED FOR CONSTANT LIFT
9
0
J
170
0 5 10 15
Balloon Volume = 27,000 cu. ft.
Gross Lift = 350 lbs.
NACA Standard Atmosphere
ALTITUDE IN THOUSANDS OF FEET
Figure 4.1
TOTAL HEAT REQUIRED TO HEAT INITIAL INFLATION AIR AND
MAINTAIN TEMPERATURES OF FIGURE 4.1 DURING EXPANSION ON ASCENT
(DOES NOT INCLUDE BALLOON HEAT LOSS)
50,0008
5 10 15
ALTITUDE IN THOUSANDS OF FEET
Figure 4.2
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sICRUT
but not large compared with the expected fuel load of 30 to 70 lb.
Values of continuous heat loss are determined as a function of altitude
for single, double and triple-layer balloons at operating temperatures as
shown in Figure 4.1. These values are given in Figures 4.3, 404 and 4,5.
They are all based on calculated U coefficients as described above.
The balloon wall temperatures expected in this example are given in
Figure 4.6 for a double-layer balloon and Figure 4.7 for a triple-layer bal-
loon. Surface temperatures for a single-layer balloon will be nearly equally
removed from the internal and external temperatures.
Referring to Section V, the energy storage capacities in Figure 5.3
can be correlated with the heat input requirement described above. In order
to provide a two-hour flight duration at 5,000 ft for a 27,000-cu ft balloon
with 350 lb gross lift, the following amounts of energy must be supplied:
Initial Inflation and Ascent (Figure 4.2) 62,000 Btu
Climb at 440 ft/min (Figure 4.3) 99,000 Btu
Floating for Two Hours at 5,000 ft, zero wind (Fig. 4.3) _860,000 Btu
Total Energy to Start of Descent 1,021,000 Btu
The fuel requirements during the descent period are dependent upon the
control method used. It is possible that no further addition of heat would
be required. Figure 5.3 indicates that a total weight of 81 lb for tank and
fuel (with propane) would provide a useful output of 1,200,000 Btu, a value
which provides a small safety factor in the above example.
B. Experiments
Several controlled experiments were performed with small balloons
(7-ft diameter) to determine the U values for hot air balloons. In addition,
estimates were made from earlier inflations of 20, 30 and 39-ft balloons.
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1,000,000
HEAT LOSS FOR SINGLE WALL BALLOON AT TEMPERATURES OF FIGURE 4.1
800,000
700,000
600,000
500,000
400,000
300,000
5 10 15
ALTITUDE IN THOUSANDS OF FEET
Figure 4.3
Balloon Volume = 27,000 cu. ft.
Gross Lift = 350 lbs.
NACA Standard Atmosphere
HEAT LOSS FOR DOUBLE WALL BALLOON AT TEMPERATURES OF FIGURE 4.1
600,000
100,000
0
400,000
300,000
200,000
5 10 15
ALTITUDE IN THOUSANDS OF FEET
Figure 4.4
HEAT LOSS FOR TRIPLE WALL BALLOON AT TEMPERATURES OF FIGURE 4.1
500,000
IOOA000
400,000
300,000
200,000
5 10 15
ALTITUDE IN THOUSANDS OF FEET
Figure 4.5
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5 MPH RELATIVE AIR MOTION
ZERO RELATIVE AIR MOTION
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DGl,RL I
WALL TEMPERATURES FOR DOUBLE WALL BALLOON
AT INTERNAL TEMPERATURES OF FIGURE 4.1
5 10 15
ALTITUDE IN THOUSANDS OF FEET
Figure 4.6
INNER WALL - ZERO RELATIVE AIR MOTION
INNER WALL - e MPH RELATIVE AIR MOTION
Balloon Volume= 27,000 cu. ft.
Gross Lift = 350 lbs.
.NACA Standard Atmosphere
WALL TEMPERATURES FOR TRIPLE WALL BALLOON
AT INTERNAL TEMPERATURES OF FIGURE 4.1
ALTITUDE IN THOUSANDS OF FEET
Figure 4.7
0 5 10 15
INNER WALL - ZERO RELATIVE AIR MOTION
INNER WALL - 5 MPH RELATIVE AIR MOTION
OUTER WALL - ZERO RELATIVE AIR MOTION
OUTER WALL - SMPH RELATIVE AIR MOTION
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SEGRET
Table 11.1 summarizes tests of single-layer balloons. The lift values agree
with the graphs of Section II. The heat transfer values are in general
agreement with the value of 0.82 used in the above calculations. Tests with
the 7-ft balloon resulted in an average coefficient U of 1.067 Btu/hr-ft2-oF,
which is considerably higher than those found in other tests. This is due
primarily to the very great temperature difference which was required in
these tests to produce measurable lifts in the small balloons.
Table 1+.2 presents more detailed information for five tests on a single-
layer Mylar balloon. This data points to an increase in the U value with
increased temperature differentials.
These tests were conducted, as shown in Figure 4+.8, utilizing propane
bunsen-type burners. Fuel consumption was measured with a calibrated orifice-
manometer combination, and lift was determined by changes in the reading of
a balance. Chemical analysis of the products of combustion was made with an
Orsat Analyzer. Figure 4+.9 presents the temperature distributions determined
by a thermocouple probe. Mean temperatures and measured lift are correlated
with theoretical lift values. Very good agreement was found.
Ventilation losses (from a 6-in."diameter vent port in the side of the
balloon) were determined from the measured internal and external tempera-
tures and the Orsat analysis by making use of the equations in Section III.
A 7-ft diameter, double-layer, polyethylene balloon and a similar Mylar
balloon were fabricated. These were made by sealing tubular gores together
and gathering the ends to form a cylinder balloon. A small quantity of air
was placed in the tubes to separate the two walls. This method was not
entirely satisfactory in that the air collected in the top and bottom of
the gores, leaving the layers in contact with each other in the center. An
- 24+ -
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TABLE 4+.1
SUMMARY OF UNIT HEAT LOSS VALUES - EXPERIMENTS WITH
HOT AIR BALLOONS
Balloon
Size
Q T
Lift
Lb/Ft3
U
Btuhr-ft2
?F
-
7 ft
151?F
0.0160
1.067
20 ft
50?F
0.00355
0.85+
30 ft
43.7?F
0.0033
0.7+8
39 ft
75?F
0.00926
0.776
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TEST CONFIGURATION FOR LIFT MEASUREMENTS
FIGURE 4.8
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TEST
TE
MP
ER
ATU
RE
No.
A
B
C
D
E
F
G
H
I
J
K
L
M
N
O
P
Q
R
S
T
U
V
W
X
y
l
fV
N
Cw
?
m
N
2
N
N
N
1
N
8
N
1
O
N
N
N
-
M
-
0
N
(Y)
4`
4t
t~~
OD
MM
N
N
t
~
N
2
N
N
N
N
p~
N
N
N
N
M
N
N
N
!O
N
N
'D
'a
m
N
w IQ
cm Of
N
.
m
ED
d
O
O
O
-
-
-
-
Cy
(V
N
N
N
N
N
lV
N
N
N
N
N
N
N
N
N
N
N
N
M N
~
N
I: I
N
N
N
N
N
N
-
W
4 J
Dj
lj
~~ N
N
N
a-D
N
N
N
10
1
1,
N
1
0
0
N
N
WI
M)l
1
N
0
~
C\j
(v
N
=
5
N
L0
1
N
w
0
N
N
N
rol
N
w
N
N
N
N 'n1
N
0
"
m 1
1
2
1"),
8
N
N
N
a
,
N
6
N
N
N
N
M
CV
M
N
N
N
N
N
N
N
N
N
19
N
(,I
N
N
N
I
NOTE: ALL TEMPERATURES ARE FAHRENHEIT
FIGURE - 49.
BALLOON TEMPERATURE DISTRIBUTIO
2571 82 1.0173 12501
2441 80 1.0163 1 220 [~
258 1 85 1.0155 1230[1
SEC r,
ET
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2781 84 1.0183 1270f
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SECRET
estimated 30 to 50 per cent of the area was acting as a single-wall balloon.
The U values obtained from the polyethylene balloon are shown in
Table 4.3, and those pertaining to the Myrlar balloon are given in Fable 404.
These U coefficients are somewhat lower. than those of Table 4.2 for the
single-layer balloon. However, they do not reflect the expected improvement
corresponding to an effective double-layer balloon, due to the fabrication
difficulties discussed above.
The combined results presented in Section II, III and IV indicate that,
from the point of view of minimum fuel consumption, operation with relatively
low temperature differentials and large balloons will be advantageous. This
is true because the area to volume ratio decreases with increased size, and
the heat transfer coefficient, U, decreases moderately with reduced tempera-
ture differential.
A further conclusion from the investigation of heat loss is that ordinary
single-wall balloons may be expected to have U values between 0o75'and 101
Btu/hrmft2'oF over a range of temperature differences from 50?F to ~500p.
- 28
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0
41 0
O 0
ON
r
a\ N OD i 0
0 rI H 4 H
O 0 0 0 0
r- H H H -i
lH 0 \D N
(Y) LrN M
LrI\ C&
r-4 0 N N
0 c0rl 00 00
0
t`
H N N Cu M
O 0
(drON
N r-I W
0
O~ 00
r-i
r-I
to
W 4-1
0 0
1?a^ CdH
w o o x
Lr\
30
O qq
0 0 0
\ to 0 Lf\
U) C 0 r-i 0
W m
I U
29
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CO 0 O\ 0 0
uj \U ti rn
frl M - Lt\ \10"
t` - N CO
00 00 0 CD
o \0 1? 'D U \o
0 0 0 0 0
CO M O O\ O-\
H r-I -i r -q
O\ CO OD -:t [--
4 M N M M.
E-1 Lr\ Lr\ r-4 is N
Lr\ r-4 r-q
QO rHI rM-I H
Q i Lf \
Cd A r-i
9
to LrI\ to
rH r A r-i
to H l~-
O\ -~ ON
?5. ^ . q
OD O\ co
ON 0
UN
Cu
M
g4 c O
E-1
a C cc C
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C
P O\
G r-
O O
M N? 0
ti 0 N
O O O
N N 0 Cu N ..~
Am m '
00 Co CO CO
r-q r-I r-q H
p pv v
0-
\,b
tt~ V~ t`
N N Cu N
0 0 0 0
0 u` 0 tr\
w
Cu cn rj M
0 0 0 0
O' H r-! -
a Cu M
w 0 w.
VJ N N- CO
tYl N \O Cu
N
Co CO CO d Co O
O ONp
OO O 0 O O 0 0 0 0 0 0 0 0 0
U O O O O O O O 6 6 6 6 6 6 6
.f O N Co
N? N \o VJ tr\
r-I r-H r-q ri rH
r4 O rn N rn
H r-I-I r-!1
tYl uJ M
i H r-I
- O \D O tr\ 0 u\ --t N U\ co. tr\ -zt r i r 1 N 4 N 4 M M M M rh M M
.r.{ 0
Cc
O\ a O
0
OD
H
00
r-q
CD
ri
CO
H
00
r-I
co
r-q
CO
r-I
00
H
CD
H
00
H
Lr\
01\
I'D
0
1
0,
Q\
w
N
w.
0
w.
N
.
P
0
CO
-
N-
0
O:
M
M
\O
O
N
N
/-1
r-q
r-!
N
N
O
O
0
O
O
O
O
O
O
0
O
w
O
w
w
w
w
0
tr\
CD
U'\
0
0
N-
r-q
H
-4
r-I
Cu
N
Cu
Cu
tr\ tf\
0
Lr\
\O
0
0
o ti
Co
0
0
0
\0
CO
chi
N
CO
N
\O
-
r-!
H
w
w
w
w
w
.
w
w.
w
w
H
rI
N
M
M
tr\
M
tr\
\p
CO CO Co u) 0
cr\ N o
O\ 0\
O O O O O O O
O\ - N N tr\ t--
\L) tJ \,O t--
r-4 r-! H r-q r-!
G Q\ O O
r-4 r-I N Cu
O Cu
O\ r-I _4 H
r
H r4
-30?
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8
0 O U to
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V. FUEL STORAGE
Tanks for storage of fuel for continuous heat supply have-been studied.
Analyses pertaining to storage tank weight have been made, based on a spher-
ical shape. This shape has the most advantageous area to volume ratio and
can be used for a field unit.
Tanks having internal volumetric capacity of 1,000, 2,000 and 3,000 cu
inches have been used in illustrative calculations. Their respective dia-
meters are 12.5, 15.6 and 18 inches and their respective capacities in U. S.
gallons are 4.33, 8.66 and 12.99.
Four high-strength materials have been studied for this application.
They are stainless steel (No. 316), aluminum (615-Tl+), glass fiber, and the
aluminum alloy 56S-H38-
Tank weight valued have been calculated from the following calculations,
pertaining to a thin hollow sphere:
_ 2S
Pr
t is the wall thickness, inches
S is the allowable unit stress, psi
P is the design internal, psig
r is the internal radius of the sphere, inches.
r is the internal radius of the sphere, inches
V is the internal volume of the sphere, in3.
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dv = 4 IT r2 dr (5-3)
dV is the volume between the inner and outer surfaces of the
tank, in3
r is the internal radius,,inches
dr = t, the tank thickness, inches.
Wt
= dV
Wt is the weight of the tank shell, pounds
,;P is the density of the tank material, lb /in3.
where
Wf + Wt
(5.5)
t is the tank efficiency, per cent
Wf is the weight of the fuel, pounds
Wt is the weight of the tank, pounds.
The weight of tanks, based on the above equations, is directly related
to the storage pressure. Exceptions to this statement occur with low pres-
sure fuels where pressure is no longer the design criterion and minimum thick-
nesses for structural integrity are used. Figure 5.1 shows calculated weights
for 1,000, 2,000 and 3,000-cu inch tanks at working pressures to 1,000 prig.
Stresses are based on a safety factor of 2.0, referred to the yield point of
the material.
Propane is the most volatile fuel which was considered in detail. It
has a vapor pressure of 286 psia at 130?F. From Figure 5.1, the corresponding
- 32 -
100 W f
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SECRET
weight for a 3,000 in3 tank would be 5 to 12 lb, depending upon the material
selected. Tank efficiencies, based on a fuel of density equal to propane,
are shown in Figure 5.2. A tank efficiency of 82 to 92 per cent would be
expected (at a design pressure of 286 psi).
The weight of a spherical tank and the fuel for a given energy storage
capacity are shown in Figure 503. The material is No. 316 stainless steel.
A combustion efficiency of 85 per cent (as defined in Section III) has been
assumed. The quantities of energy storage.in Figure 503 are correlated with
expected duration in Section IV. Figure 5.3 illustrates the great similarity
of the several hydrocarbon fuels when compared on'.the storage weight basis.
Selection of fuels will, therefore, be based primarily on other considera-
,tions, of which controllability of combustion is the most significant.
The major conclusion from the analysis of fuel tanks is that adequate
lightweight tanks can be made which will contribute only slightly to the
weight of the over-all hot air balloon system.
m33o
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WEIGHT OF SPHERICAL TANKS WITH SAFETY FACTOR OF 2.0
20
0
0
400 ?600
WORKING PRESSURE (p.s.i)
Figure 5.1
EFFICIENCY OF SPHERICAL TANKS
1
X1000 CU. IN
CODE:
Stainless Steel (316)
- - -- Aluminum (56S-H38)
- -Aluminum (61 S - T4)
- - -Glass Fiber
0
400 600
WORKING PRESSURE (p.s.i.)
Figure .5.2
SEC E T
Fuel Density equal to Propane
Tank Volume, 3000. cu.in.
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0
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70
J
60
IL.
~c
50
Z
9
0
40
H
c2
30
W
0
1 1 1 1 1 1 I
.PROPANE
BUTANE
GASOLINE
0 200 400 600 800 1000 1200 1400
HEAT OUTPUT OF FUELS IN THOUSANDS OF BTU
(85%, Efficiency)
Assuming
Minimum tank wall thickness of 0.040? for propane and butane
Minimum tank wall thickness of 0.030' for gasoline
Spherical tanks - 0.80 volume factor applied to propane. and butane
Material, stainless steel
Figure 5.3
S
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VI. INFLATION
The operation of a hot air balloon requires an initial inflation with
air. It has been found that the natural induction of air by the operation
of the burner is not sufficient to fill the balloon. Inflation of the bal-
loon by some powered means is required. The purpose of this section of the
report is to describe the analysis and experimental work which led to the
.selection of prototype hardward to meet the above requirements, as described
in Section VII.
A. General Analysis
A balloon volume of 27,000 ft3 was taken as a nominal maximum value
and an inflation duct of 15-in. diameter was arbitrarily selected as a
practical design limit. Inflation rates consistent with filling times of
40, 30, 20 and 10 minutes were considered. The above assumptions resulted
in inflation velocity values which then made possible the calculation of
velocity head and air horsepower values. Formulas 2c,2d used in this analysis
are given below:
Q = AV (6.1)
is the volume flow rate, ft3/min
A is the free inflation duct area, ft2
V is the average velocity, ft/min.
(6.2)
by is the velocity head, inches of water
V is the average velocity, ft/min
36
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4005 is a constant, evaluated for normal sea level conditions.
P = 6350
(6.3)
P is the volume rate of flow, ft3/min
H is the total head, inches of water
6350 is a constant evaluated for normal sea level air.
It was assumed that the inflation duct would have a re-entrant entrance
and an abrupt exit. From reference 4 this combination results in the equation:
H = 1.85 by (6.4)
The magnitude of the total head, H, may be reduced by aerodynamic
design of the entrance. However, it is questionable that the additional cost
of an entrance of this type could be justified, for a semi-expendable field
Further assumptions were made for the case in which inflation could
be accomplished by an electrically-driven fan. An adiabatic fan efficiency
of 60 per cent and a motor efficiency of 50 per cent were arbitrarily set.
Over-all efficiency of the combination is 30 per cent. Results are given in
Table 6.1.
It can be seen from Table 6.1 that the power requirements are quite
low and an electrically-driven inflation fan is reasonable. Other possi-
bilities for delivering these quantities of power to a fan were also con-
sidered. Specifically considered were the use of an internal combustion
engine (as built for models) and the application of a hand crank through
gear train.
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TABLE 6.1
POWER REQUIRED, FOR INFLATION
Inflation Time
(min)
Q
Ft3 min
V
(Ft/min)
by
in.H 0
H
in.H o
Shaft HP
P 60% Eff.
(HP) (HP)
Motor HP
50% Eff.
(Hp)
40
674
548
0. oig
0.035
0.0037 o 00617
0.01234
30
900
732
0.034
0.063
o? oo89 0.0148
0.0296
20
1350
1100
0.075
0.139
,0.0295 0.0492
o.0984
10
2700
2200
0.300
0.555
0.236 0.393
0.786
Model engines capable of delivering power throughout the range of
Table 6.1 are readily available and represent a highly-developed, lightweight,
compact power source. The primary disadvantage of the model internal com-
bustion engine is its questionable starting reliability in the field. Ex-
perience indicates that this power source would be somewhat less reliable
than the electric motor when compared on the basis of starting the unit after
lengthy storage. Other special problems associated with this motive unit are
the requirement of a special fuel and the lack of good speed control. Both
of these latter difficulties could be overcome if there-`we~ future incentive
to use this motive unit.
Reference 2e indicates that a man can exert energy at a rate of
90,000 ft lb/hr (0.0454 HP) for a considerable length of time, the duration
being dependent upon the atmospheric conditions. At an effective temperature
of 105?F'the expected duration would be approximately one-half hour. At
lower temperatures this duration increases significantly. Referring to
Table 6.1, expenditure of energy at this rate would result in an inflation
? time slightly more than 20 minutes. In comparing this method with the use
of an electrically-driven fan, it is concluded that it could serve as an
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5LGRET
emergency method in case of electrical system failure.
B. Experimental Work
Because of the need for a relatively high-volume flow rate with low
pressure differential, an axial flow-type fan appeared most suitable. In
order than an electric motor have minimum weight, a high shaft speed was
considered a desirable feature. This combination defines a low torque re-
quirement for the motor. It was found that small wooden propellers made for
model airplane work had these desirable characteristics.
Tests were run on several of these propellers having various diameters
and pitch angles. During the tests the propellers were driven by a direct
current aircraft-type motor operating from a 12-volt wet cell. The test unit
was mounted in a 14.5-in. diameter duct, as illustrated in Figure 6.1. Motor
power input was measured by voltmeter and ammeter, speed by a stroboscopic
tachometer, and air velocity by a vane anemometer. Results of these tests
are given in Table 6.2.
TABLE 6.2
PROPELLER PERFORMANCE
(Motor Voltage - 12V)
Propeller
Diam. & Pitch
Current
a s
Speed
z? m
Air Vel.
ft min
Air Flow
(cfm)
Inflation Time
for 27,000 ft3
(min)
Over-all Eff.of
Fan & Motor
(, )
12"
5
4.17
3870
898
1100
24.5
22.8
12"
8
4.90
3200
967
1185
22.8
24.4
14"
6
4.90
3200
l046
1280
25.4
30.9
14"
8
5.75
3200
862
1060
21.2
1.4.8
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ID, Z7
INFLATION FAN TEST CONFIGURATION
0
Joy
Motor
1 Test
15 11 Propeller -
Ammeter i
Cross section of test duct
showing anemometer positions
Figure 6.1
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The performance of the 14-in. diameter propeller with 6-in. pitch
was found to be most efficient and was therefore recommended for use in the
prototype burner-inflation assembly.
Further tests were made on operation of the electric motor and pro-
peller with dry cell batteries. Battery drain tests were made with various
combinations of 6-volt dry cells (Burgess F4BP). It was found that six of
these cells connected to-supply a nominal 18 volts resulted in an inflation
time of 30 minutes for a 27,000-cu ft balloon. Six cells connected so as to
provide a nominal 12 volts resulted in an inflation time of 35 minutes, and
four cells connected to provide 12 volts resulted in an inflation time of
38 minutes. Voltage, current and speed as a function of time are shown in
Figures 6.2, 6.3 and 6.40
Since these batteries weigh approximately 1-1/4 lb each, the use of
a pack of six weighing 7-1/2 lb appears to be a very practical solution to the
inflation problem. Wet-type 12-volt batteries, weighing 15 lb and used in
light airplanes, are available which would reduce the inflation time to
approximately 25 minutes.
A preliminary performance test was run on a "McCoy 60" model aircraft
engine. This motor is rated at 1.32 HP at 17,000 rpm, which is somewhat higher
than the range of values in Table 6.1. When operating with a 14-in. propeller,
it delivered 4,000 cfm, a rate which corresponds to an inflation time for a
27,000-cu ft balloon of approximately 7 minutes. This rate of inflation is
higher than that obtained with the electrical system and points to the use of
this type of motive unit in cases where extremely rapid inflation is required.
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BATTERY DRAINAGE TESTS
SIX BATTERIES (BURGESS F4BP) OPEN CIRCUIT VOLTAGE, .I8V
TIME, MINUTES
Figure 6.2
20 10 20 30
eed
0
8
oits
4
ms
BATTERY DRAINAGE TESTS
SIX BATTERIES (BURGESS F4BP) OPEN CIRCUIT VOLTAGE, 12V
O
Its
6
S eed
Am s
F
T I
30
TIME, MINUTES
Figure 6.3
SECRE
28
26 0
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24 Sr
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20 a
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%.0 %j I!- 1
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'IN'd*H d0 SO38ONnH NI O33dS
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? VII. PROTOTYPE BURNER ASSEMBLY
A prototype burner and inflation assembly has been constructed for
this project. This unit is designed to provide a continuous, controllable
heat input to the balloon as well as incorporating the initial inflation
equipment. The assembly is shown as viewed from the top in Figure 7.1 and as
viewed from the bottom in Figure 7.2. Figure 7.6 shows the over-all dimen-
sions of the assembly. The prototype burner assembly consists of the follow-
ing components:
1. Inflation fan and motor
2. Fuel tank (Propane)
3. Ignition system
4. Main burner
5. -Tank heater
6. Relief and pressure regulating valves.
The total weight of the entire protype assembly excluding the fuel
supply is 59 lb.
A. Inflation Fan and Motor
The inflation fan and motor are mounted in the assembly on a detachable
spider frame. The motor is a direct-current compound-wound unit made by the
Joy Manufacturing Company and is the motive unit for their fan model number
X-702-29A. The fan blade is a wooden model-aircraft propeller 14-inches in
diameter with a 6-in. pitch and bears the "Rite Pitch" trade name. Power
consumption of this combination is 60 watts at 12 volts and its air flow rate
is 1,280 cu ft/min. The resultant inflation time for a 27,000-cu ft balloon
is 21 minutes.
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SECRET
3
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m
W
2
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B. 'Fuel Tank (Propane)
The fuel tank selected for the prototype unit is a standard (military)
low-pressure-breathing oxygen tank made of stainless steel with banded con-
struction. Its volumetric capacity is 2,100 cu in. with a design working
pressure of 1+00 psig. This tank has been fitted with an internal dip tube of
3/8-in. diameter copper to facilitate use of the tank in a horizontal posi-
tion. Gaseous fuel is drawn off the top without danger of liquid carryover.
The tank is equipped with a pressure relief valve set at 375 psig.
This oxygen tank was used in the prototype model because of its general
suitability, having approximately the desired capacity and an ample safety
factor for use with propane. The tank is less efficient weightwise than those
discussed in Section V because of its elongated shape (cylinder with hemi-
spheric ends) and its high design working pressure. This tank holds approxi-
mately 35 lb of propane.
C. Ignition System
An,'electrical spark-type ignition system is incorporated in the pro-
totype model to.provide remote ignition and a reliable method of relighting
the burner during flight. This ignition system utilizes two model sizes of
spark plugs which are provided with intermittent high voltage through separate
ignition coils energized by a relay. The electrical diagram for this system
is shown in Figure 7.7. Photographs of the completed unit are shown in
Figures 7.3 and 7.1+. The ignition system provides instantaneous lighting of
the main burner unit.
D. Main Burner
The main burner unit is designed to operate on high pressure propane
throughout a pressure range from 5 to 100 lb psig. The ring for the burner
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BURNER CONTROLS WITH IGNITION CONTROL BOX
FIGURE 7.3
(ZP1 P T
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W
w
M
0
I'
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has 60 separate combustion heads, each of which induces small quantities
of air. The burner heads are a commercial item (No. 1352-BU) manufactured
by Otto Bernz, Inc. Mounting nipples of brass tubing are silver-soldered
into a hard copper manifold. This design was selected in order to minimize
manifold size and weight and yield efficient operation of a wide range of
fuel input rates.
It was found experimentally that a stainless steel collector ring is
necessary to insure instantaneous ignition of all of the burner heads. This
unit provides a communicating channel for the gas at the time of ignition.
The burner is calibrated for fuel consumption rate versus manifold
pressure by measurement of fuel weight loss over known time intervals.
These fuel consumption values are given in Table 7.1.
The burner is mounted in a housing made of aluminum. This housing
serves as a windshield and mounting structure for the inflation fan and motor.
The burner is insulated from the housing by sheets of corrugated asbestos.
Rings are provided on the upper surface of the frame for attaching the burner
to the balloon.
E. Tank Heater
The continuous evaporation of the liquid propane requires an input of
heat which is greater than that obtainable from the surrounding air. This
need is met by a small propane burner tank heater containing two high-pressure
propane burners of the same type as the main burner, mounted in a shield or
enclosure as shown in Figure 7.5. Heat from the products of combustion is
transferred to the tank by natural convection. A stainless steel radiation
shield is provided by means of the burner and tank to prevent local over-
heating.
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TANK HEATER BURNER ASSEMBLY WITH BAFFLE REMOVED
FIGURE 7.5
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Since, the requirement for heat input to the fuel varies directly with
the consumption of the main burner, it is desirable to connect the tank
heater in parallel with the main burner. By this method, linear modulation
of the tank heater with demand is accomplished.
F. Relief and Pressure Regulating Valve
The tank relief valve is a Superior No. 1032B-X-1 set at 375 psig.
This valve is a combination filling valve and pressure release.
A manually-adjustable pressure regulator has been selected to allow
modulation of the manifold pressure between 3 and 130 psig. This unit is
a "M-B Model R-G" automatic pressure-reducing and regulating valve which
maintains constant downstream pressure with variable upstream pressure.
TABLE 7.1
CALIBRATION OF MAIN BURNER
Manifold Pressure
Psig
Heating Rate
BTU/hr
10
85,000
20
155,000
30
225,000
40
280, 000
50
345,000
60
410, 000
70
470,000
80
530A00
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SEC
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3M.ALT
VIII. CONTROL OF HOT AIR BALLOONS
Control of the lift in a hot air balloon will be required, particularly
in manned operations. Three possibilities have been considered:
1. Valving of hot air from the top of the balloon
2. Drawing air from the bottom of the balloon by reversing the in-
flation fan
3. Modulating the fuel input to the burner..
A. Valving
Calculations have been made of the valving areas required for lift
loss rates up to 20 lb/min. These are shown in Figure 8.1. Valving areas
equivalent to circular openings 6 to 18 inches in diameter result.
Figure 8.1 is based on the assumption that a 27,000-ft3 balloon
operates at a constant lift of 350 lb at altitudes from sea level to 15,000
ft by maintaining the required internal temperature. At every altitude the
density difference between inside and outside is 0.01296 lb/ft3. The required
rate of air flow for a lift loss rate of 20 lb/min is 25.7 ft3/sec. Low
lift loss rates require proportionally lower air flow rates.
Flow in this case is incompressible, and the equation below is valid:
CA V r,0__
911
Q is the flow rate, ft3/sec
C is the orifice coefficient (assumed 0.60)
A is the free flow area, ft2
g is the acceleration of gravity ft/sec2
h is the motive head, ft.
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(8.1)
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0
The motive head1 h~ can be-determined from the height of the valve above
the zero pressure level as follows:
h Pa Pb) L
Pb
(8.2)
e a is the density outside the balloon, lb/ft3
e b is the density inside the balloon, lb/ft3
L is the height of the valve above the zero pressure level, ft.
B. Reversing Inflation Fan
The inflation fan used in the prototype unit handles 1,280 cfm. This
corresponds to a lift loss rate of 16.6 lb/min following the analysis above.
The possibility has occurred of reversing this fan so as to draw air from the
balloon for control purposes,. thereby eliminating need for a valve in the
top.
C. Modulating Fuel Input
Modulation of the fuel input to the burner appears to have merit as a
control technique. The prototype assembly described in Section VII is cap-
able of operating over a range of 10 per cent to 100 per cent of its full
capacity, thereby allowing control of the lift. The dynamic response of
the system to this type of control is not reliably calculable, however, and
must be determined by field experiment.
56
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E
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HIM
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IX. CONCLUSIONS
From the work covered in this report, several conclusions are drawn.
These are given below:
1. The lift produced in a hot air balloon is consistent with the
classic theory of buoyancy. The graphical solutions of Section II provide
accurate determination of lift and required balloon size.
2. Maximum surface temperature occurs at the apex of the balloon and
is approximately equal to the average internal air temperature as shown in
Figure 4.9. This means that, even though the general surface temperature is
considerably lower than the internal air temperature, the balloon must be
capable of withstanding an actual film temperature equal to the design average
internal air temperature. Mylar is a more satisfactory balloon material than
polyethylene because of its higher maximum operating temperature. Mylar
balloons should be satisfactory for temperatures up to 250?F and polyethylene
should be limited to a temperature of 150?F.
3. Heat loss through the balloon film can be closely predicted by
the method outlined in Section IV. Fuel input must be greater than the heat
loss through the film plus ventilation heat loss, as evaluated in Section III.
4'The most satisfactory fuel for a hot air balloon is Propane, which
can be contained as a liquid and burned in the gaseous form. This fuel allows
broad modulation of the fuel input with relatively simple, lightweight equip-
ment.
5? Combustion mixtures involving approximately 200 per cent excess
air are recommended to insure complete combustion and to maintain a low dew
point within the balloon.
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OCURC I
6. Electric ignition systems can successfully provide remote-controlled
re-lighting of the burner with a low-weight dry battery power source.
7. Initial installation of a man-carrying hot air balloon can suc-
cessfully be accomplished in approximately 30 minutes with an electrically-
driven inflation fan, similar to the type used in the prototype, with a dry
battery power-pack weighing 7.5 lb.
8. A duration of two hours at 5,000 feet is considered compatible with
a 27,000-ft manned-balloon system utilizing a propane fuel system, as dis-
cussed. in Section IV*
9. The field of multilayer' balloons holds promise of reducing the
fuel consumption from that of a single-layer balloon. Double-wall balloons
were built (see Section N) which showed moderate improvements. Reductions
in heat loss more nearly consistent with the theoretical analysis should be
obtainable if a successful method of separating the layers is found.
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X. REFERENCES
to Marks, L. S., Mechanical Engineers' Handbook, Fifth Edition,
McGraw-Hill, New York,1951, (a) P- 311.3, (b) P. 348.
2. Heating Ventilating Air Conditioning Guide, Volume 33, American
Society of Heating and Air Conditioning Engineers Inc., New York, 1955,
(a) P? 356, (b) p. 174
179, (c) P? 735, (d) P. 759, (e) p. 117.
3. Timoshenko, S., Strength of Materials, Part I, Second Edition,
D. Van Nostrand Co.., New York, 1940, p. 42.
1+. Carrier, W. H., Cherne, R. E., and Grant, W. A., Modern Air
Conditioning, Heating and Ventilating, Pitman, New York, 1940, p. 235.
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