1

Analytical and Experimental Comparisons of HTPB and

ABS as Hybrid Rocket Fuels

Stephen A. Whitmore*, Zachary W. Peterson † and Shannon D. Eilers‡

Utah State University, Logan, UT, 84322

Acrylonitrilebutadiene-styrene (ABS) thermoplastic is evaluated as a

potential fuel for hybrid rocket motors and compared to the widely used fuel

material Hydroxyl-Terminated Polybutadiene (HTPB). For this project Nitrous

Oxide (N2O) is used as the matching oxidizer. ABS is widely mass-produced for

non-combustion applications including household plumbing and structural

materials. Analytical predictions for enthalpies of formation, products of

combustion, and end-to-end motor performance are compared against static-

burn test results for ABS fuel grain manufactured using Direct-Digital

Manufacturing (DDM). The chemistry and performances of the ABS fuel grains

are compared against fuel grains cast from HTPB. ABS grains were DDM-

fabricated from existing stock materials composed of 50:43:7 butadiene,

acrylonitrile, and styrene mole fractions. The grains were tested and compared

against HTPB grains of equal physical volume and dimension. Test results

demonstrate a higher burn-to-burn consistency for ABS fuel grains, but slightly

* Associate Professor, Mechanical and Aerospace Engineering Department, 4130 Old Main Hill, UMC 4130, Associate Fellow, AIAA. † Graduate Research Assistant, Mechanical and Aerospace Engineering Department, 4130 Old Main Hill, UMC 4130, Student Member, AIAA. ‡ Graduate Research Assistant, Mechanical and Aerospace Engineering Department, 4130 Old Main Hill, UMC 4130, Student Member, AIAA.

47th AIAA/ASME/SAE/ASEE Joint Propulsion Conference & Exhibit31 July - 03 August 2011, San Diego, California

AIAA 2011-5909

Copyright © 2011 by Utah State University. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission.

2

reduced overall performance. The reduced performance is primarily attributed

to overall lower combustion efficiency for the ABS grains. The combustion

effectiveness of relative monomer compositions for the ABS fuel grains were

investigated; and equilibrium chemistry calculations conclude that for a given

oxidizer to fuel ratio, varying the butadiene mole-fraction in the ABS

formulation has a significant effect on the propellant performance.

Nomenclature

A =free energy of formation curve fit intercept parameter

ABS = acrylonitrilebutadiene-styrene

Aburn = fuel grain surface burn area, cm2

Ac = fuel chamber cross sectional area , cm2

Aexit = nozzle exit area , cm2

Aox = effective total oxidizer injector area, cm2

A* = nozzle throat area, sonic choke area, cm2

B = free energy of formation curve fit slope parameter

C = carbon atom

Cd ox = effective oxidizer injector discharge coefficient

CF = thrust coefficient

cp = specific heat of solid propellant fuel grain, J/kg-K

= propellant characteristic velocity, m/sec c*

3

F = steady state thrust, N

g0 = standard acceleration of gravity, 9.8067 m/sec2

H =hydrogen atom

HTPB = hydroxyl-terminated Polybutadiene

hν = heat of ablation /gasification, MJ/kg

Isp = specific impulse, s

Isp effective = effective specific impulse, s

Itotal = total impulse, N-s

Mfuel = total fuel consumed during motor burn, kg

Mox = oxidizerfuel consumed during motor burn, kg

MW = molecular weight of combustion products

= solid fuel grain mass flow rate, kg/sec

= oxidizer mass flow rate, kg/sec

N = degree of polymerization

N2 = gaseous nitrogen

N2O = nitrous oxide

n = generic mole fraction

na = acrylonitrile mole fraction

nb, = butadiene mole fraction

ns = styrene mole fraction

!mfuel

!mox

4

O/F = oxidizer to fuel mass flow ratio

OH = hydroxyl radical

Pr = Prandtl number

Pox = oxidizer injector upstream feed pressure pressure, kPa

P0 = combustion chamber pressure, kPa

Rg = gas-specific constant, J/kg-K

Ru = universal gas constant, 8314.4126 J/kg-K

= fuel grain linear regression rate, cm/sec

T = combustion temperature, K

T0 = adiabatic flame temperature, K

Vc = combustion chamber internal volume, cm3

ΔG = Gibbs free energy of formation, kJ/mol

∆G◦gg = Gibbs polymerization free energy for gaseous state, kJ/mol

∆G◦mono = Gibbs free energy for monomer, kJ/mol

= Gibbs polymerization free energy, amorphous solid from liquid monomers, kJ/mol

∆G◦poly = Gibbs free energy for polymer, kJ/mol

ΔH = mole specific enthalpy change, kJ/mol

ΔHf0

= standard enthalpy of formation, kJ/mol

ΔQc = heat of combustion, kJ/mol

ΔS =mole specific entropy change, kJ/mol-K

!r

!G0la

5

= combustion efficiency

γ = ratio of specific heats or combustion products

µ = dynamic viscosity, Nt-s/m2

µ = sample mean

ρfuel = solid fuel grain material density, kg/m3

ρox = liquid oxidizer density, kg/m3

σ = sample standard deviation

I. Introduction

ESS risky launch vehicles must be developed to operational readiness for commercial

spaceflight to become a viable market. A key to commercial spaceflight success is the cost-

efficient and safe operation of large rocket motors in civilian environments. Because of the

extreme risk and potentially negative environmental impact, large solid-propelled rocket motors

simply cannot be operated at civilian airports. Approximately 70% of catastrophic space-launch

failures are attributable to the vehicle’s power plant.1 NASA estimates that the Space Shuttle’s

liquid fueled main engines will fail catastrophically once every 1530 sorties per engine and its

solid rocket boosters will fail catastrophically once every 1,550 sorties per motor. While this

level of risk is acceptable for experimental and government-operated vehicles, it is unacceptable

for a potential commercial spaceflight operator.

Hybrid motors have the potential to fulfill this low-risk flight requirement. The inherent safety

of hybrid rocket motors greatly reduces the operational risk to a launch vehicle or any hybrid

rocket propelled missile or spacecraft. According to the U.S. Department of Transportation,

!*

L

6

hybrid motors can be stored and operated without possibility of explosion or detonation.2 Other

advantages of hybrid rockets include the ability to be stopped, restarted, and throttled, easy (and

hence potentially cheaper) ground handling, and relative insusceptibility to grain flaws.

A. Current Hybrid-Motor Manufacturing Limitations

Hybrid rocket motors have been in development for more than two decades, but have seen

little commercial application mainly due to difficulty associated with of mass-producing fuel

grains. Motor-to-motor performance consistency has also been a commercial disincentive. To

date, the majority of mainstream hybrid motor development has focused on performance

characterization and improvement. Very little development effort has been directed towards

manufacturing methods to support volume production. The current production method involves

casting viscous or non-polymerized material “by hand” in a casing mold to form the fuel grain,

to be later manually assembled with other motor components (i.e., igniter, oxidizer tank, oxidizer

valve/injectors, post-combustion chamber, and rocket nozzle). This low-tech approach results in

market prohibitive production costs and a high degree of performance variability. Motor–to-

motor response times, peak thrust, and total impulse levels can vary significantly. This high

degree of motor-to-motor variability may be acceptable for experimental or government vehicles

operated on a restricted-test range, but will not secure FAA certification for non-experimental,

commercial spaceflight operations. Finally, the current low-technology, labor intensive

fabrication processes used to cast and de-gas hybrid-motor fuel grains simply cannot produce the

numbers and varieties of motors required to support the ambitious launch rates necessary to

support what is expected to be a fast-growing commercial space industry.

Leveraging the recent rapid capability growth in factory automation and robotics, Direct-

Digital Manufacturing (DDM)3 offers the potential to revolutionize methods used to fabricate

7

hybrid rocket fuel grains. DDM technology can support high production rates with a much

greater degree of motor-to-motor consistency than is possible using traditional “one-off” motor

casting methods. If matured and commercialized, this technology will have a transformational

effect on hybrid rocket motor production by improving quality, consistency, and performance,

while reducing development and production costs.

II. Chemical Analysis of Hydroxyl-Terminated Polybutadiene (HTPB) and Acrylonitrilebutadiene-

Styrene (ABS) as Hybrid Rocket Fuel Grains

Many modern thermosetting polymers and thermoplastics can be mass-produced with a

uniform consistency, but have not been evaluated as hybrid rocket fuels. The most commonly

used hybrid rocket fuel, Hydroxyl-terminated polybutadiene (HTPB) is a legacy thermosetting

polymer material frequently used as binder for solid rocket propellant grains. It remains the

hybrid fuel grain material of choice primarily because of industry familiarity with its chemical

and structural properties. HTPB is a solid homopolymer made by combining two components: 1)

a hydroxyl-terminated polymer of butadiene and 2) a cross-linking agent, either isocyanate or

methylene diphenyl diisocyanate (MDI), used to polymerize and set the material. No industry

standard value for the enthalpy of formation exists. The enthalpies of formation of HTPB have

been published, but vary widely -- from 64.88 kJ/g-mol4, and 363.17 kJ/kg5, to -31.57 kJ/g-mol6 -

- as the relative degrees of polymerization and curative treatments for the mix are modified.

HTPB, as a thermosetting polymeric material, cannot be shaped and manufactured using

DDM methods. HTPB must be mixed from its liquid base-components, degassed under vacuum,

and then cast and cured in a fuel grain mold. HTPB does not melt in the presence of heat; but

instead chars and ablates. HTPB burn properties can vary dramatically depending on the degree

8

of residual gas seeding in the cast material, and the length of time the material has cured. Typical

cure times exceed 10 days to two weeks.

Alternatively, Acrylonitrilebutadiene-styrene (ABS) is an inexpensive thermoplastic that can

be easily produced in wide variety of shapes. More than 1.4 billion kilograms of ABS material

were produced chemical and petrochemical industries world wide in 20107. ABS plastics are

readily shaped into complex geometries using DDM techniques. If ABS can be demonstrated to

be thermodynamically and structurally competitive with HTPB, it would represent a very

attractive option as a hybrid rocket fuel grain material. This material is widely mass-produced for

a variety of non-combustion applications including household plumbing, and structural materials.

To date, however, no detailed analysis of the performance of ABS as a rocket fuel has been

performed. This paper details the rocket-based combustion of ABS polymer variations with

Nitrous Oxide (N2O), and compares its performance against HTPB as a baseline. Both analytical

predictions8,9 and experimental verifications will be presented. Analytical predictions include

shifting-equilibrium chemistry effects, fuel regression rates, and unsteady thrust and impulse

predictions. For this initial study, simple cylindrical port grain configurations were modeled and

tested.

Because ABS is widely used as an industrial and domestic construction material, the heat of

combustion ΔQc, in the presence of oxygen has been documented for safety considerations.10

However, because ABS has not previously been purposely considered as a potential rocket fuel

grain marterial, there has been little to no work quantifying the material’s standard enthalpy of

formation, ΔHf0. The enthalpy of formation is required to calculate the properties of the

combustion products when ABS is burned in the presence of nitrous oxide at various mixture

ratios and combustion pressures. As with HTPB it is expected that the ABS enthalpy of

9

formation will vary widely depending on the ratio of the various monomers that make up the

ABS polymer.

B. Group Addition Method for Calculating the ΔHf0 of Polymers

Because there are no reliable, industry standard values for the enthalpies of formation for both

HTPB and ABS; this study employs a systematic approach for calculating ΔHf0 using the “Group

Addition” methods developed by Van Krevelen and Chermin11,12. Van Krevelen’s method

calculates the Gibbs free energy of formation ΔGf, of a polymer in a gaseous state by summing

individual group contributions of common polymer-building molecular groups. The group

contribution to ΔGf is assumed as a linear function of temperature, with the form

. (1)

In Eq (1), A and B are linear-fit constants, and T is the temperature of formation. Using the

fundamental definition for Gibbs free energy,

. (2)

Inspecting Eqs. (1) and (2) shows that the equation intercept is identical to the standard enthalpy

of formation, and the slope corresponds to the standard entropy of formation. In Eq. (2) ΔS is the

entropy change, and ΔH is the enthalpy of formation. Because the Gibbs free energy of a

polymer can be estimated as the summed contributions of the individual molecular groups

enthalpies of formation minus the contributions of entropy / temperature product; it follows that

the enthalpy of formation of the polymer is simply the summation of the contributions of the

individual groups to the enthalpy of formation.

!Gf0group

= A + B "T

!G = !H " T # !S

10

The Gibbs free energy values given by Ref (11) are expressed for materials in a gaseous state.

Because polymers do not exist in a gaseous state, a correction is performed to account for

condensed states. Reference 12 applies a correction to the Gibbs free energy of polymerization of

the gaseous state, where the Gibbs free energy of polymerization is given by

(3)

The Gibbs free energies of a polymer and its monomer equivalent, and are

obtained by summing the group contributions as described above. A correction term is then

added to the free energy of polymerization to account for the formation of an amorphous (non-

crystalline) solid or liquid polymer from the gaseous polymer. The correction for the Gibbs free

energy of an amorphous solid polymer created from monomers in a liquid state is

. (4)

Inspecting equations (1), (2), and (4) shows that the Gibbs free-energy correction only affects the

entropy of formation and not the enthalpy or formation. Thus, it can be concluded that the

enthalpies of formation of a substance in an amorphous state is approximately the same as that

substance when in a gaseous state. Thus the enthalpies of formation can be calculated directly

from the presented data in Ref (11).

C. Calculating ΔHf0 for HTPB

To verify applicability of the group addition process for rocket combustion, the method

described in the previous paragraph was used to calculate the heat of formation of HTPB. Since

HTPB with Nitrous Oxide as the oxidizing agent is the most commonly used hybrid rocket

propellant combination; a large data based of performance data exists and can be used to verify

!Gogg = !Go

poly " !Gomonomer

!Gopoly !Go

monomer

!G0la = !G0

gg " 40 #T

11

the additive group calculations. As an introduction to the group addition method, a detailed

calculation for HTB will be presented here.

A common polymer formulation of HTPB (ARCO R-45M®13) has 50 repeating butadiene

units with a hydroxyl radical at each end14, thus the moniker “hydroxyl terminated.” Figure 1

depicts the molecular structure of the bonded HTPB chain showing the polymerized trans/cis 1,4

butadiene units, and the hydroxyl radicals attached to each end of the string. The cis/trans

structures have slightly different enthalpies of formation; however these differences are

considered negligible compared to the overall heat of polymerization, and were neglected for this

analysis. The corresponding chemical formula is C200H302O2, and the mean molecular weight is

2734 kg/kg-mol.

Figure 1. Chemical Structure of N=50 Polymerization of HTPB.

The major constituent of HTPB, 1,4 butadiene has the chemical formula,

.

The heat of formation for each butadiene molecule is calculated by

!CH2 ! CH = CH ! CH2 !( )

12

. (5)

The heat of formation contribution of each hydroxyl group is

.

Each component’s mole fraction within the N=50 polymerization is

, (7)

and the molecular heat of formation is

(7)

The “reduced” chemical formula corresponding to the calculated value ΔH0f HTPB is

(8)

Using Eq. (8) the corresponding heat of formation in mass units is calculated at 456.5 kJ/kg -- a

value that is approximately 25.6% higher than that value reported by Ref. (5). The ΔH0f

presented by Eq. (7) neglects the chemical contributions of the curing agent used to polymerize

and set the grain material. However, the curing agent represents a small fraction of the cured

2 ! "CH2 "( )2 ! "CH =( ) +

2 ! "44 kJg"mol

#

$%

&

'(

2 ! ( 76 kJg"mol

)

________Total )H f butadiene

32 kJg"mol

!H f0

OH " #176 kJg#mol

nbutadienenOH

!

"#

$

%& =

50 / 522 / 52

!

"#

$

%& =

96.15%3.85%

!

"#

$

%&

!H 0f HTPB

= nbutadiene " !H0fbutadiene

+ nOH " !H 0fOH

= 0.9615( ) " 32 kJg-mol

#

$%

&

'( + 0.0385( ) " -176 kJ

g-mol

#

$%

&

'(

= 23.99 kJg-mol

0.9615( ) ! C4H6( )+ 0.0385( ) ! OH( ) = C3.8462H5.8077O0.0385

13

material – approximately than 12% by mass15 -- and for the purposes of this discussion the

effects on ΔH0f are considered to negligible.

D. Thermodynamic and Transport Properties of HTPB/Nitrous Oxide Combustion

The value for ΔH0f HTPB calculated by Eq. (7), and the “reduced” molecular formula given by

Eq. (8) were directly input into the NASA program “Chemical Equilibrium with Applications,”

(CEA)16 with HTPB as 100% of the fuel component of the propellant§. The CEA program was

configured to calculate the thermodynamic and transport properties of the combustion products

as a function of combustion pressure P0, and the mass flow oxidizer-to-fuel (O/F) ratio.

Calculated parameters include adiabatic flame temperature T0, ratio of specific heats γ, molecular

weight MW, characteristic velocity c*, dynamic viscosity µ, and Prandtl number Pr. Figure 2 plots

these thermodynamic and transport properties as a function of O/F ratio and the combustion

chamber pressure. Depending on the chamber pressure, the peak c* performance occurs at an O/F

ratio between 5.0 and 6.0. At 7000 kPa chamber pressure (580 psia) the corresponding

interpolated peak c* value is 1630 m/sec. This value agrees closely with the “textbook” peak

value of 1625 m/sec for hybrid HTPB/N2O combustion at 1000 psi chamber pressure (Ref. 4).

This 0.3% level of agreement supports the accuracy of the group addition ΔH0f calculation

methods.

§ The CEA code has extensive internal libraries for reactant thermodynamic and transport properties including standard and nonstandard temperature and pressure conditions. Unfortunately, the CEA thermochemical database does not have an entry for and of the formulations of HTPB. Consequently, HTPB fuel-grain properties, including the atomic mole-fraction formula and enthalpy of formation, must be externally input to the program.

14

Figure 2. Thermodynamic and Transport Properties of N2O/HTPB Combustion Products.

E. Calculating the ΔHf0 of Various ABS Polymer Formulations

For this preliminary analysis the ABS formulations were varied substantially to evaluate the

best performing propellant formulations for the material. The monomers that make up any ABS

thermoplastic are 1) acrylonitrile, 2) butadiene, and 3) styrene. The corresponding chemical

structures for these monomers are17

(9)

acrylonitrilebutadienestyrene

!

"

###

$

%

&&&=

CH2 = CH ' C ( NC4H6

C6H5CH = CH2

!

"

###

$

%

&&&

15

Figure 3 shows the representative chemical structure of the ABS polymer for various monomer

constituencies, where nb, na, and ns are the respective mole fractions of Acrylonitrile, butadiene,

and styrene.

Figure 3. ABS Chemical Structure.

Readily available industrial formulations of ABS tend to consist of approximately 1/2

butadiene; with a typical formulation consisting of 50% (mole fraction) butadiene, 43%

acrylonitrile and 7% styrene. While the authors recognize that the monomer ratio may not

produce a propellant with optimal energy content, the ready availability of this stock product

favored this formulation for analysis and testing during this phase of the project. Table II shows

monomer ΔHf0, heats of polymerization ΔQpolymerization, the corresponding polymer ΔHf

0, mole

fractions, and enthalpy contributions of the three ABS constituent monomers using the 50:43:7

monomer-ratios and the group addition method. The enthalpy of formation is calculated as 62.63

kJ/g-mol, and the “reduced” chemical formula corresponding to the calculated value ΔH0f ABS is

(C3.85H4.85N0.43.) (10)

The equivalent mass-based enthalpy of formation is 1097.4 kJ/kg. This value is roughly 140%

higher than the corresponding value calculated form HTPB. The higher calculated enthalpy of

formation suggests that this ABS formulation will likely not burn as hot as will HTPB.

16

Table 1. Enthalpy of Formation Contributions of Acrylonitrile, Butadiene, and Styrene Co-

polymers in a 50%/43%/07% ABS Formulation

Monomer ΔH f0

monomerKJ / g − mol

ΔQ polymerization KJ / g − mol

ΔH f

0 polymerKJ / g − mol

ABS Mole

fraction

EnthalpyContributionKJ / g − mol

Acrylonitrile 172.6218 74.3119 98.31 0.43 42.27

Butadiene 104.10(Ref.

11)

72.10 20 32.00 0.50 16.00

Styrene 146.9121 84.60(Ref. 19) 63.31 0.07 4.36

ABS Total 62.63

F. Thermodynamic and Transport Properties for ABS/N2O Combustion

The value of ΔH0f ABS calculated in Table 1 and corresponding chemical formula Eq. (10),

were input to CEA to calculate the thermodynamic and transport properties of the combustion by

products as a function of combustion pressure and O/F ratio. Calculated parameters include T0, γ,

MW, c*, µ, and Pr. Figure 4 plots these ABS/N2O Combustion properties as function of O/F ratio

and combustion pressure. When these graphs are compared against the corresponding HTPB-

based figures (Figure 2), two key features stand out: 1) the predicted c* is lower by slightly less

than 1%, and 2) peak c*values tend to occur at O/F ratios between 4.0 and 5.5. This “optimal

point” is significantly lower than the corresponding HTPB O/F optimal values that lie between

5.0-6.0. Slightly lower ABS-motor oxidizer flow levels should produce equivalent performance

when compared to the HTPB-motor.

17

Figure 4. Thermodynamic and Transport Properties of N2O/ABS Combustion Products.

III. End to End Motor Performance Modeling

The motor models to be described here will be compared against motor burn data from static-

fire tests to be described later in the Experimental Results and Discussion section. The motor

modeling equations are based on the enthalpy-balance regression rate model developed by Ref.

8, adjusted for non-unity combustion product Prandtl number. Assuming the nozzle throat

chokes immediately, as the propellant burns the generated gases cannot escape as fast as they are

produced and pressure within the fuel chamber builds. A balance between the gases coming into

the fuel port and the gases leaving through the choked throat calculates the time response of this

chamber pressure growth. Here the equation that describes the time evolution of the chamber

pressure is

18

. (11)

The oxidizer mass flow rate is modeled by the incompressible discharge coefficient formula

. (12)

Equation (12) is reasonably accurate as long as the motor is burned using a “top pressure” that is

higher than the saturation pressure of the N2O at the injector temperature. For “blow down”

systems that use only the natural vapor pressure of the propellant, a more complicated two-phase

model is required to accurately model the injector mass flow22. The fuel grain mass flow is

modeled as

, (13)

where is the linear fuel regression rate and ρfuel is the material density of the solid fuel grain.

The O/F ratio of the burning propellant, which varies significantly throughout the burn, is given

by

. (14)

Modeling the mixture ratio in this form allows for calculation of combustion product properties

at every time step using a table lookup of the equilibrium exhaust gas properties depicted in

Figure 2 and Figure 4. The crux of the hybrid motor model used here is the fuel grain regression

rate equation, corrected by Ref. (22) for non-unity Prandtl number

!P0

!t=

Aburn !rVc

"fuel RgT0 # P0$% &' # P0

A*

Vc

( RgT0

2( +1

)*+

,-.

( +1( #1

$

%

///

&

'

000+

RgT0

Vc

AoxCdox2"ox (Pox # P0 )

!mox = AoxCdox

2!ox ( pox " p0 )

!mfuel = Aburn ! " fuel ! !r

r•

O / F =

!mox

!mfuel

=AoxCdox

2!ox ( pox " p0 )

Aburn # ! fuel # !r

19

. (15)

In Eq. (15) the parameters µ and Pr refer to the combustion product gas properties, the

parameters Pox and ρox refer to the oxidizer liquid properties upstream of the injector, and cp, ρfuel,

Tfuel, and hv refer to the properties of the solid fuel grain.

Equations (11) through (15) are continuously integrated to calculate the chamber conditions,

and at each time frame the chamber properties are used with standard one-dimensional one-

dimensional De-Laval flow formulae23 to calculate the end-to-end rocket performance

parameters. Calculate parameters include thrust, exit properties, specific impulse, and choking

mass flow. Adjustments are made for non-zero nozzle exit angles, and heat loss within the

combustor. Nozzle throat dimensions are allowed to erode at prescribed rates.

IV. Experimental Setup for HTPB and ABS Fuel Grain Static Fire Tests

A primary emphasis of this project was the experimental evaluation of ABS as a hybrid fuel

grain, and comparison of its performance relative to geometrically identical fuel grains cast from

HTPB. Multiple, identically-cast HTPB, and DDM-manufactured ABS fuel grains were static

fired in a test cell housed on the Utah State University campus. This section details the hybrid

motor case design and assembly, fuel grain fabrication methods, test cell instrumentation, and

test procedures. Tests results and model comparisons will be presented later in the results and

discussion section.

G. Motor Case, Injector, Nozzle, and Ignitor Assembly

!r = 0.047

!fuel " Pr0.153

cp T0 # Tfuel$% &'hv

(

)*

+

,-

0.23AoxCdox

Ac

2!ox (Pox # P0 )$

%..

&

'//

45 µ

L()*

+,-

15

20

For these tests a commercially available 98-mm solid-rocket motor24 was modified by

replacing the original ejection charge on the motor cap with a single port oxidizer injector. A

threaded pressure transducer was also installed in the modified motor injector cap. To reduce

run-to-run variability due to nozzle erosion, graphite nozzles fabricated from a single piece of

graphite replaced the original manufacturer-supplied phenolic nozzle. The nozzle was designed

to have a 4.2:1 expansion ratio and had a design throat diameter of 1.7 cm. Figure 5 shows the

original Cesaroni solid-rocket 98-mm motor case adapted for these hybrid motor tests.

Additional advantages of this configuration are a ready-made “flight-weight” motor and the

ability to rapidly reload between motor tests. Table 2 summarizes the mean motor and fuel grain

dimensions and mass properties.

Figure 5: Cesaroni Pro-98 Motor Case Adapted for Hybrid-Motor Tests.

Two small Estes-class 10-gram solid rocket ignition motors were inserted into the injector

cap. Electronic matches burned by a 12 volt DC signal ignited these small motors. The motors

and e-matches were replaced after each test firing. Figure 6 shows photographic images of the

nozzle assembly, injector/motor cap assembly, and the injector port, ignition motors, and

electronic-matches.

21

Table 2. Motor Case and Fuel Grain Dimensions.

Length,

mm

Diameter,

mm

Nozzle

Throat

Area, A*,

cm2

Nozzle

Throat

Area, Aexit

cm2

Nozzle

Expansion

Ratio**

Injector

Exit

Area,

cm2

Nozzle

Discharge

Coefficient,

Cd ox

Motor

Case

98 702 2.21 9.27 4.2 0.115 0.40

Fuel

Grain

Length,

cm

Diameter,

cm

Initial

Port

Area, A*,

cm2

Post

Combustion

Chamber

Volume,

cm3

Heat of

Ablation /

Gasification,

hν, MJ/kg

Initial

Fuel

Grain

Mass, kg

Fuel Grain

Material

Density,

kg/m3

HTPB

571.5 82.6 5.07 31.95 1.825 2.55 930.0

ABS

571.5 82.6 5.07 31.95 2.326 2.67 975.0

** Based on the expected chamber pressure, this expansion ratio was optimized for the altitude of the test site, Logan Utah. Logan has an altitude of approximately 1.47 km above mean sea level.

22

Figure 6. Nozzle and Motor Cap Assembly.

H. ABS Fuel Grain Fabrication and Test Grain Geometry

Multiple cylindrical ABS fuel grains were fabricated using a Stratasys Fortus®27 Fused-

Deposition Modeling (FDM) system. Typically, these systems are used for 3-dimensional

modeling and rapid-prototyping. The Stratsys FDM systems use production-grade thermoplastic

materials that are sufficiently rugged for functional testing and end-use parts. Real production

thermoplastics are stable and have no appreciable warping, shrinkage, or moisture absorption.

The ABS stock material, Stratasys ABS-M30® thermoplastic, used for these tests approximately

the 50:43:7 chemical formulation presented earlier in Section II.F. The mean density of the ABS

23

stock material used for these test was approximately 975 kg/m3. This density was considered to

be sufficient to insure structural integrity of the fuel grain during the static test firings.

The fuel grain was fabricated to fit snugly into the motor case described in the previous

section. A post combustion chamber was pre-manufactured into the fuel grain. Each fuel grain

was approximately 57.15 cm in length, 8.26 cm in diameter, the initial fuel port diameter was

2.54 cm, and post combustion chambers were 5.66 cm in diameter and 1.27 cm deep. The mean

fuel grain weight was approximately 2.67 kg. To reduce the chances of initiating erosive

burning, the injector port was designed to protrude approximately 1 cm into the top of the fuel

grain port. Each fuel grain was bonded into a cardboard sleeve using high temperature silicone

adhesive. This sleeve is then inserted into a phonelic insulating liner, and then into the aluminum

motor casing. The nozzle and motor cap assemblies are seated with O-rings. The aluminum

motor case acts as the pressure vessel for the motor. Figure 7 presents the fuel images showing

the grain features with the major dimensions labeled.

Figure 7. ABS Fuel Grain Images with Dimensions.

I. HTPB Fuel Grain Fabrication and Test Geometry

24

As mentioned earlier in section II.C HTPB fuel grains were cast using the commercially

available Arco R45M® polybutadiene resin and PAPI 94® MDI curative. Arco R45M is

polybutadiene diol manufactured by the Sinclair Petrochemicals Arco Division. The resin has a

polymerization factor of approximately 50 and a molecular weight of 2734 kg/kg-mol. PAPI 94

is a polymethylene polyphenylisocyanate produced by Dow® Plastics Inc. The formulation

contains methylene diphenylene diisocyanate (MDI) in proprietary proportions. The curative has

an average molecular weight of 290 kg/kg-mol. The nitrogen, carbon, oxygen (N-C-O) bonds in

the MDI react with the hydroxyl (OH) terminations in the polybutadiene resin to cure the fuel

grain. For these tests activated charcoal was added to the mixture to insure opaqueness and

prevent radiative heating of the fuel grain and motor case liners. HTPB/MDI/charcoal mass

proportions were set at 87%/12.5%/0.5%, respectively. Past experience has determined that these

proportions assure adequate fuel grain cure and material hardness.28

The resin and curative were mixed in a commercial paint mixer that was sealed and fitted so

that the fuel mixture could be placed under a vacuum during the mixing process. A commercial

H-VAC vacuum pump was used to remove gas bubbles created in the fuel grain during the

mixing process. The de-gassed mixture was cast in the same cardboard sleeves used for the ABS

fuel grains with a 2.54 cm (1”) polyvinyl chloride (PVC) pipe used as mandrel. Before casting

the mandrel was coated with a mold release agent to insure proper release after the fuel grain

cured. The HTPB fuel grain dimensions, including the post combustion chamber were identical

to the previously described ABS fuel grains. The mean density of the ABS stock material used

for these test was approximately 930 kg/m3, and the cast fuel grains had a mean mass of 2.55 kg.

After casting the fuel grain were heat cured and the hardness was continuously monitored

through out the cure process. The HTPB fuel grains took as much as 15 days to reach their final

25

hardness levels before static testing. Figure 8 presents “cure hardness” data for a traditionally

cast HTPB†† fuel grain to support this assertion. Here measured Shore–durometer hardness29

units are plotted in curing days. After 3 1/2 days, the material has reached only 50% of its final

hardness; and even after 12 days, the grain is still only 90% of its fully-cured hardness. Burning

a polymeric fuel grain before it fully cures can produce unpredictable ablation rates, and may

cause significant erosive burning. Erosive burning produces chaotic pitting and channeling along

the length of the fuel grain, and poses a significant potential hazard to the structural integrity of

the motor.

Figure 8. HTPB Fuel Grain Hardness as a Function of Cure Time.

J. Test Cell Apparatus and Instrumentation

As mentioned earlier, an existing test cell at Utah State University was used to perform the

motor characterization tests. Measurements obtained include chamber pressure, thrust, total

impulse, motor case temperatures, specific impulse, mass flow rate, consumed propellant mass,

and propellant regression rate. Following each motor test, the fuel grain were dissected and

visually inspected for erosive burning and structural failure. Representative fuel grain images

†† Arco Poly BD R-45 HT resin (87%), 12.5% Dow PAPI 94 polyethylene polyphenylisocyanate curative, 0.5% activated charcoal powder. Data collected April 5-25 2011.

26

will be presented later in the Results and Discussion section. All tests results will be compared to

analytical predictions for the specific motor fuel grain materials in that section.

1. Hybrid Test Stand Oxidizer Delivery System

To allow sufficient mass flow rates with minimal line losses, a predetermined mass of N2O

Oxidizer was delivered to a closely coupled “run tank” from a series of “K” sized industrial

pressure cylinders. The run tank was pressurized by gaseous nitrogen (N2) to insure a constant

injector pressure during the entire length of the burn. The N2 “top pressure” was set by a manual

regulator; and was typically maintained near 5200 kPa (800 psi) for these tests. The top pressure

kept the N2O above saturation pressure for the entire run and insured a single-phase liquid flow

through the injector. The design motor chamber pressure was approximately 3440 kPa (500 psi),

The pneumatic run valve was triggered by an electronic relay, and was automatically controlled

by the instrumentation software. Oxidizer mass flow was sensed by vertical load cells mounted

on the run tank, and by an inline Venturi flow meter mounted in the oxidizer feed-line just ahead

of the injector. Figure 9 shows the piping and instrumentation diagram (P&ID) for the hybrid

motor test arrangement.

27

Figure 9. Oxidizer Piping and Instrumentation Diagram (P&ID) for Hybrid Motor Tests.

2. Hybrid Test Stand Thrust Balance

The 6-degree of freedom (6-DOF) thrust stand developed and modified for this project

provided real-time vertical (motor mass) and axial load (thrust) measurements. Originally a total

of 6 load cells were installed on ball-joint mounts to reduce axial cross coupling. The 6-load cells

allowed all load paths to be identified. For this series of tests only the axial and vertical loads

were of primary interest. To reduce vibration during the static load tests, the lateral load cells

were replaced with support rods mounted on ball shafts. The axial load was sensed by an

Omegadyne® LCCD-500 (2225 Nt) load cell, and the vertical loads were sensed by two

Omegadyne® LCCA-25 (110 Nt) load cells. The output response for these sensors is 3 mV/Volt,

and the sensors were excited using a 12-volt DC power source. Chamber pressure was sensed

using an MSI®-600 (0-6900 kPa) absolute threaded pressure transducer mounted in the motor

28

cap (Figure 6). Omegadyne® Type-K thermocouples were mounted at the aft-end of the motor

case to sense burn temperature and thermal soak-back following the end of the burn. Relevant

manufacturer's specifications including operating range and accuracy for each of these

instruments is listed in Table 2.

Table 3. Manufacturer Specifications for Static Thrust Stand Instruments

Instrument Model Operating Range Accuracy

LCCA-25

(Lateral Loads)

+25 lbf (110 N) +0.037% of Full Scale

LCCD-500

(Axial Loads)

+500 lbf (2225 N) +0.25% of Full Scale

MSI-600

(Chamber Pressure)

0-1000 Psia (0-6900 kPa) +0.1% of Full Scale

Two National Instruments data acquisition and control devices managed motor fire control,

and logged test data. An NI-compact DAQ® 4-slot bus controller with multiple analog input (16-

bit), analog output, digital output, and thermocouple modules (24-bit) bus-cards managed the

majority of the measurements and valve control. The digital outputs from a separate NI USB-

6009® module were used to trigger the relays that fired the ignitor e-matches. Operators and

experimenters were remotely located in a secure control room separated from the test area.

Communications to the test stand were managed by an operator-controlled laptop via universal

serial bus (USB) using amplified extension cables. All control and measurement functions were

controlled by a LABview® program hosted on the control laptop. Figure 10 shows a schematic

of the test stand, and an image of the motor being fired.

29

Figure 10. Static Test Mechanical and Instrumentation Schematic.

V. Results and Discussion

As mentioned earlier in the introduction to Section IV, multiple, identically-cast HTPB, and

DDM-manufactured ABS fuel grains were static fired to compare the relative performances of

the two different fuel grain types. To date 20 successful separate static-tests including 6 ABS and

14 HTP fuel grains have been performed. The data from several of these tests has been

discounted due to known instrumentation, software, and test apparatus difficulties. Burn Profiles

from representative tests for both the HTPB and ABS fuel grains will be presented here. Post-

burn fuel grain images and total-fuel linear-regression data will be presented for all of the “good”

motor burns. Finally, test results will be compared to model predictions.

K. Thrust-Stand Time History Data

Figure 11 presents side-by-side comparisons of time-history data collected from 4-

representative HTPB and ABS static firings. The presented time history profiles include thrust,

30

accumulated impulse, injector and chamber pressures, and oxidizer mass flow (as measured by

the in-line Venturi system). Each of the HTPB fuel grain burns is approximately 10.25 seconds

in duration. For the ABS burns, both short (5.25 s) and long (10.25 s) burn profiles are presented.

Because there was a lot of initial uncertainty regarding the regression rates for the ABS fuel

grains; the initial ABS fuel burns were kept to a short duration to reduce the chances of motor

case burn through. Once visual inspection determined that the regression rates were sufficiently

slow to insure motor survival; the ABS burn times were lengthened to 10.25 seconds.

For each case the HTPB fuel grains exhibit a small, but well-defined net-advantage in overall

performance, including thrust, total impulse, and specific impulse. The mean steady state thrust

level for the HTPB grains is approximately 755.0 N, and 717.8 N for the ABS grains. This

difference represents an advantage of approximately 4.9% for the HTPB grains. The HTPB

grains have an effective ensemble mean specific impulse of approximately 204.1 s. This value is

approximately 6.6% higher than the mean for the ABS fuel grains (190.6 s). The effective

specific impulse is calculated by dividing the total impulse generated during the burn by the fuel

and oxidizer mass consumed,

. (16)

The equivalent ensemble-mean vacuum specific impulse values are 225.7 s (HTPB) and 211.6 s

(ABS). The vacuum specific impulse calculation assumes nozzle exit area of 9.27 cm2, and an

ambient test pressure of 86 kPa.

It appears that the increased thrust for the HTPB fuel grains is not driven by a higher overall

mass-flow. The approximate oxidizer and fuel mass flow rates for the HTPB burns are 0.300 kg/s

(N2O), and 0.077 kg/s (HTPB) respectively. For the ABS burns the corresponding oxidizer and

Isp effective =Itotal

g0 ! M fuel + Mox( )

31

fuel mass flow rates are 0.304 kg/s (N2O), and 0.086 kg/s (ABS) respectively. The corresponding

O/F ratios -- calculated by dividing the total consumed oxidizer mass by the total consumed fuel

mass – are 4.19 for HTPB fuel grains and 3.75 for the ABS grains. Thus, the ABS fuel grains are

actually characterized by a slightly higher fuel mass burn rate. This comparison suggests that the

HTPB fuel grains burn at a higher flame temperature when compared to the ABS grains.

Calculating the characteristic velocity for the individual burns supports this assertion. Here the

characteristic velocity is approximated by

, (17)

where P0 and F are the mean steady-state chamber pressure and thrust levels. The mean c* value

for the HTPB burns is approximately 1523.1 m/sec compared to 1488.7 m/sec for the ABS burns.

This difference is approximately 2.3%, or slightly less than 1/2 of the observed thrust difference.

The thrust and impulse data also demonstrate that the ABS fuel grain delivers a slightly

greater run-to-run consistency than do the HTPB fuel grains. The steady-state thrust standard

deviation for the ABS grains is approximately +17.2 N compared to +32.1 N for the HTPB

grains. Although not large, this improved run-to-run consistency is likely a positive benefit of

the FDM fabrication methods used to build the ABS grains. FDM is a precisely controlled by

robotic industrial process; as opposed to the labor intensive mixing, degassing, casting, and

curing process for the HTPB fuel grains. This consistency difference may also be a result of

HTPB being a thermo-setting material and ABS being a thermoplastic material. Unquestionably

this topic requires further research; but it appears that the digital manufacturing methods

demonstrate the potential for enhanced consistency in the fuel grain fabrication process. Table 4

summarizes the individual and ensemble results for the HTPB and ABS burn tests.

c* =g0 ! IspCF

= g0 ! Isp ! A* P0F

"#$

%&'

32

Figure 11. Burn Time History Comparisons for HTPB and ABS Hybrid Motors.

33

Table 4. HTPB and ABS Motor Burn Summary.

HTPB

Grain

Burns

Steady

Burn

Time, s

Total

Impulse,

Itotal, N-s

Consumed

Mass, kg

O/F Ratio

Effective

Isp, s

Mean

Steady

Thrust, N

c*, m/s

VarHTPB4 10.25 8747.4 Oxidizer:

3.614

Fuel:0.80

O/F: 4.27

211.4

787.7 1495.6

VarHTPB5 10.25 8294.6 Oxidizer:

3.392

Fuel:0.794

O/F: 4.27

202.1

766.2 1488.3

VarHTPB6 10.25 8011.1 Oxidizer:

3.299

Fuel:0.795

O/F: 4.15

199.5 732.3 1541.0

VarHTPB7 10.25 7955.2 Oxidizer:

3.211

Fuel:0.787

O/F: 4.07

203.4 725.8 1567.3

HTPB Performance Mean and Standard Deviation µ Isp: 204.1

σ Isp: 5.12

µ F: 753.0

σ F: 32.09

µ c*: 1523.1

σ c*: 37.52

34

HTPB

Grain

Burns

Steady

Burn

Time,

sec

Total

Impulse,

Itotal, N-s

Consumed

Mass, kg

O/F Ratio

Effective

Isp, s

Mean

Steady

Thrust, N

c*, m/s

VarABS2 5.25 4289.1 Oxidizer:

1.802

Fuel:0.480

O/F: 3.71

193.4

733.2 1564.1

VarABS3 5.25 7532.0 Oxidizer:

1.801

Fuel:0.811

O/F: 4.18

183.0

694.8 1497.5

VarABS4 10.25 7814.7 Oxidizer:

3.387

Fuel:0.854

O/F: 3.98

187.2

714.7 1353.8

VarABS5 10.25 4431.9 Oxidizer:

3.443

Fuel:0.549

O/F: 3.14

198.9

728.5 1539.4

ABS Performance Mean \and Standard Deviation µ Isp: 190.6

σ Isp: 6.99

µ F: 717.8

σ F: 17.2

µ c*: 1488.7

σ c*: 94.08

35

L. Comparison to Model Predictions

As presented above, a likely candidate justifying the higher overall HTPB thrust and specific

impulse appears to result from the c* efficiency of the combustion process. Defining the

combustion efficiency as

, (18)

the combustion efficiency of the hybrid model, Eqs. (11) through (15), was adjusted until the

predicted thrust and chamber pressure values matched the measured values. This procedure

determined that the HTPB fuel grains tend to an η* efficiency near 95%, whereas the ABS fuel

grains have an η* efficiency below 90%. Figure 12 presents model/static test comparisons to

support this conclusion. Here the thrust profiles from Figure 11 are overlaid against model

predictions for η*=(100% and 92%) for the HTPB motor (Fig. 12a), and η*=(100% and 86%) for

the ABS motor (Fig. 12b). The corresponding predicted c* profiles are plotted in Fig. 12c

(HTPB), and Fig. 12d (ABS). This efficiency difference is approximately 9.5%, and although

higher than the c* differences calculated earlier; still suggests that the ABS grains tested did not

burn as efficiently as the HTPB grains.

!* = c*actualc*ideal

36

Figure 12 Model/Static Test Comparisons to for HTPB and ABS Hybrid Motors.

M. Fuel Grain Regression Measurements

The results and conclusions presented on Figure 11 and Figure 12, are further supported by

post-burn fuel grain regression measurements. Both the HTPB and ABS burned grains burned in

a linear manner, and no erosive burn patterns were noted. Following the end of each static test,

the motor was quenched and then split longitudinally to expose the burned grain pattern. The

final regression dimensions were measured at multiple points along the fuel grain; and the mean

end-to-end longitudinal fuel linear regression was calculated. Figure 13 shows side-by-side

comparisons of post-burn HTPB and ABS fuel grains. The regression measurement stations are

marked on each grain. For both the HTPB and ABS grains, fossilized surface flow patterns are

visible, and the transitions from laminar to turbulent flow patterns can be clearly seen. The burn

patterns are very similar with the ABS fuel grain showing an incipient erosion pattern just

37

downstream of the injector, and a slightly asymmetrical final burn profile (Fig. 13 b). For both

fuel grains the burn time was slightly longer than 10 seconds.

Figure 13 Burned HTPB and ABS Fuel Grains.

Figure 14 compares the longitudinal-mean regression measurements for all of the HTPB and

ABS burns against model predictions. Fig. 14a plots the total longitudinally averaged linear

regression against burn time. Dividing the total regression by the steady burn time approximates

the mean linear regression rate for each fuel grain. Fig. 14b plots the linear regression rate

against the mean port diameter. For the measured regression data, the mean port diameter is

approximated by the average of the initial and final fuel port diameters. The models show

reasonable agreement with the measured data, and the HTPB fuel grains are clearly demonstrated

to exhibit a faster overall linear burn rate when compared to ABS grains. These test results

directly supports the earlier data presented in Figs. 11 and 12. Even though the ABS formulation

has a greater material density, it simply does not burn as energetically as does the HTPB mixture.

38

As described earlier lower burn rate is likely a result of lower overall combustion efficiency for

the current ABS grain furmulation.

Figure 14. Predicted and Measured Linear Regression for HTPB and ABS Grains.

N. Heat of Combustion Calorimetric Measurements for the HTPB and ABS Fuel Grains.

A final series of tests were performed to add supporting data for the conclusions reached in

the previous sections. This section presents results from oxygen bomb calorimetry measurements

performed on sections of post-burn fuel HTPB and ABS fuel grains. Combustion with oxygen in

a sealed bomb is a very effective and reliable method for releasing all heat energy obtainable

from a sample, and the authors believe that these results verify the higher overall energy content

of the HTPB grain formulation when compared with the 50:43:7 ABS formulation evaluated in

these tests.

In this series of tests, a Parr model 1241®‡‡ Oxygen Bomb Adiabatic Calorimeter was used to

burn multiple fuel samples selected at random from the burned fuel grains. The gross heat of

combustion was measured according to procedures outlined in ASTM D2382-8830. Table 5

summarizes these test results. The measured gross heats of combustion for HTPB and ABS agree

well with published values (Ref. 10). The approximately 5% higher heat of combustion for ‡‡ Parr Instrument Co., Moline, Illinois

39

HTPB agrees well with the previously established performance and regression rate advantages.

Table 5. Oxygen Bomb Calorimetry Measurements for HTPB and ABS Fuel Grains.

Sample

Fuel

Grain

Weight,

g

Wire

Length,

cm

Initial

Bomb

Temp, C

Final

Bomb

Temp, C

Consumed

Wire

Length, cm

Heat of

Combustion,

ΔQc, kJ/kg

1 HTPB 0.307 15.24 22.255 23.551 6.99 42.36

2 HTPB 0.301 15.24 22.396 23.664 5.08 42.30

3 HTPB 0.283 15.24 22.273 23.465 7.37 42.25

4 HTPB 0.28 15.24 23.03 24.22 5.08 42.67

5 HTPB 0.48 15.24 22.07 24.16 5.08 43.64

ABS Fuel Grain Combustion Statistics:

µ ΔQc = 42.64 kJ/kg

σ ΔQc =0.58 kJ/kg

Sample

Fuel

Grain

Weight,

g

Wire

Length,

cm

Initial

Bomb

Temp, C

Final

Bomb

Temp, C

Consumed

Wire

Length, cm

Heat of

Combustion,

ΔQc, kJ/kg

1 ABS 15.24 15.24 22.54 23.80 7.62 39.38

2 ABS 0.32 15.24 22.42 23.70 10.80 39.47

3 ABS 0.30 15.24 23.06 24.24 8.89 39.70

4 ABS 0.31 15.24 22.68 23.92 10.16 39.71

5 ABS 0.31 15.24 22.68 23.92 10.16 39.74

ABS Fuel Grain Combustion Statistics: µ ΔQc = 39.60 kJ/kg

40

σ ΔQc =0.16 kJ/kg

O. Effect of Dissolved gaseous Nitrogen in the Nitrous Oxidizer Storage Tanks

The authors believe that the relatively low overall Isp and c* values for both HTPB and ABS

fuel grains, as presented in

Table 4 and Figure 12, indicate a dilution of the nitrous oxide storage tanks by the gaseous N2

used to pressurize the system. It is believed that during system verification and cold flow tests,

gaseous nitrogen from the pressurizing tanks (Item 1, Figure 9) was forced into the nitrous oxide

storage tanks (Item 2, Figure 9) during the process of supercharging and filling the oxidizer run

tanks (Item 6, Figure 9). This nitrogen gas became dissolved in the stored nitrous oxide.

To support this premise, CEA-based calculations were performed assuming that the

“oxidizer” consisted of 80% N2O and 20% N2 by mass. Figure 15 presents these results. Here the

c* values for the original propellant mixture (Figs. a, c) are compared to the c* values with a 20%

N2 dilution of the oxidizer. Clearly, there exists a significant impact on the propellant c*. The

peak O/F ratio is shifted to a much “leaner” value, and the maximum c* value drops by more

than 5% for both the HTPB and ABS calculations.

41

Figure 15. Effect of 20% Dissolved N2 in N2O Oxidizer on c*.

When the nitrous oxide storage tanks were replaced with a freshly-charged set of tanks. The

overall performances for both fuel grains improved dramatically. Figure 16 presents this result.

Here the model/static thrust comparisons of Figure 12 are repeated, but with the static thrust data

obtained using the “fresh” nitrous oxide tanks replacing the original data. The mean steady thrust

level for the HTPB motor has now increased from the original mean value of 755.0 N to 823.9 N.

The corresponding test and vacuum specific impulses have increased to 222.7 s and 244.3 s,

respectively. The combustion efficiency now begins to approach 100% of the theoretical value.

In a similar manner, the mean steady thrust level for the ABS motor has now increased from

42

717.8 N to 783.5 N, and the corresponding test and vacuum specific impulses have increased to

208.1 s and 229.2 s, respectively. The effective η* for the ABS fuel grain has now increased to

approximately 95%. For the test cases presented in Figure 16, the HTPB motor still exhibits a

slight advantage compared to the ABS motor; with the new difference barely changing at 4.9%

for 6.2% for Isp, respectively.

Figure 16. Model/Static Test Comparisons to for HTPB and ABS Hybrid Motors with

Recharged N2O Tanks.

At the time of the writing of this report additional HTPB fuel grains have been cast and tested

using the recharged nitrous oxide tanks. The performance gains presented Figure 16 in have been

consistently observed in these follow-on tests. Additional ABS fuel grains are being fabricated

and will be tested as they become available. Test procedures to prevent the issue with nitrogen

gas dilution of the oxidizer storage tanks have been put in place to avoid this issue in future tests.

P. ABS Fuel grain Monomer Ratio Optimization

As the data presented for the earlier 50:43:7 butadiene/acrylonitrile/styrene monomer mole-

fraction demonstrate; the energy content of this ABS formulation is slightly, but consistently

lower than the R-45m HTPB formulation. This final section examines the individual ABS

43

monomer mole fraction ratios to search for an optimal formulation. The mole-fractions of each

polymer were varied from 0 to 100% of the constituents of the polymer, such that the total

makeup of each of the formulations equals 100%. For example, if the first component was 50%

of the total constituency, the other two components were varied so that the sums of their mole

fractions totaled the other 50%. The group addition method was used to calculate ΔHf0 for the

various ABS polymerizations; and this data along with the corresponding molecular formulae,

was input into the CEA program for various oxidizer-to-fuel (O/F) ratios and combustion

pressures. The resulting values for the equilibrium species concentrations, thermodynamic

properties, and transport properties were used to calculate the propellant’s characteristic velocity

for each of the possible monomer combinations. The monomer ratios producing the higher c*

values give the more desirable propellant combinations.

Figure 17 plots the characteristic velocity for ABS/N2O combustion at 2500 kPa chamber

pressure as a function of butadiene mole fraction for various remaining acrylonitrile/styrene

(A/S) ratios. Calculations for different mixture ratios – O/F=4.5 and O/F=6.0 -- are presented.

Clearly, for the lower O/F ratio (4.5), there exists an optimal butadiene content that peaks near a

30% total mole fraction and 90% A/S ratio. For lower A/S ratios this optimal point diminishes. In

contrast, for the higher O/F ratio (6.0) there is no optimal point, and increasing butadiene content

always leads to a better performing propellant. For the higher O/F ratio the propellant

performance is more or less independent of A/S ratio. For a 2500 kPa operating chamber pressure

and an O/F ratio of 6.0, 70% butadiene content results in a c* value 1618 m/sec, that exceeds the

HTPB c* value 1608 m/sec for the same operating pressure and O/F ratio (Fig. Figure 2). At this

juncture in the research, it is unclear what allowable fraction of butadiene will retain adequate

structural integrity for the fuel grain material. This analysis suggests that future preparations of

44

ABS may be better optimized for hybrid propellant performance; but also the mixture that

produces best results is closely coupled with the operating O/F ratio of the motor.

Figure 17. Effect on ABS Monomer Mole-Fractions on Combustion Performance.

VI. Conclusion

A primary objective of this work was to investigate the viability of Acrylonitrilebutadiene-

styrene (ABS) as a hybrid rocket fuel material. ABS is an inexpensive thermoplastic that can be

easily produced in wide variety of shapes. ABS plastics are readily shaped into complex

45

geometries using Direct Digital manufacturing Techniques. Leveraging the recent rapid

capability growth in factory automation and robotics, Direct-Digital Manufacturing (DDM)

offers the potential to revolutionize methods used to fabricate hybrid rocket fuel grains. DDM

technology can support high production rates with a much greater degree of motor-to-motor

consistency than is possible using traditional “one-off” motor casting methods. Once matured

and commercialized, this technology will have a transformational effect on hybrid rocket motor

production by improving quality, consistency, and performance, while reducing development

and production costs. If demonstrated to be thermodynamically competitive with Hydroxyl-

Terminated Polybutadiene (HTPB), ABS presents a very attractive hybrid rocket fuel option.

The performance of ABS thermoplastic is compared to HTPB with Nitrous Oxide (N2O) as

the matching oxidizer. Multiple ABS grains were DDM-fabricated from existing ABS stock

materials composed of 50:43:7 butadiene, acrylonitrile, and styrene mole fractions. These grains

were static-tested and compared against traditionally cast HTPB grains of equal physical volume

and dimension. Both the HTPB and ABS grains burned in a linear manner, and no erosive burn

patterns were noted. Test results demonstrate a higher burn-to-burn consistency for ABS fuel

grains, but slightly reduced performance. For the original series of tests the hand-cast HTPB

grains exhibited approximately 4.9% greater steady thrust and 6.6% higher effective specific

impulse. End-to-end linear regression rate and bomb calorimetry measurements of the heats of

combustion support the static-fire burn performance data. In all cases the ABS formulation does

not regress as rapidly or burn with as much heat as does the HTPB material. Possible sources of

this inefficiency could be the higher heat of gasification for ABS, and a higher enthalpy of

formation for the stock ABS materials tested.

46

The authors believe that the relatively low overall Isp and c* values for both HTPB and ABS

fuel grains indicate a dilution of the nitrous oxide storage tanks by the gaseous N2 used to

pressurize the system. When the nitrous oxide storage tanks were replaced with a freshly charged

set of tanks, the overall performances for both fuel grains improved dramatically. With the

recharged tanks the performance margin between the HTOB and ABS fuel grains remained

essentially unchanged. At the time of the writing of this report additional HTPB fuel grains have

been cast and tested using the recharged nitrous oxide tanks. The performance gains with

recharged oxidizer tanks been consistently observed in these follow-on tests. Additional ABS

fuel grains are being fabricated and will be tested as they become available. Test procedures to

prevent the issue with nitrogen gas dilution of the oxidizer storage tanks have been put in place

to avoid this issue in future tests. Furthermore, nitrogen has been replaced by helium as the

pressurizing fluid.

This study concludes with an examination of the effects of relative monomer compositions for

the ABS fuel grains. This final section examines the individual ABS monomer mole fraction

ratios to search for an optimal formulation. Calculations for different mixture ratios were

presented. For the lower O/F ratio (4.5) there exists an optimal butadiene content that peaks near

a 30% total mole fraction and 90% A/S ratio. For the higher O/F ratio (6.0) there is no optimal

point, and increasing butadiene content always leads to a better performing propellant. For the

higher O/F ratio, 70% butadiene content results in an ABS c* that exceeds the HTPB c* for the

same operating pressure and O/F ratio. At this juncture in the research, it is unclear what

allowable fraction of butadiene will retain adequate structural integrity for the fuel grain

material. This analysis suggests that future preparations of ABS may be better optimized for

hybrid propellant performance; but also the mixture that produces best results is closely coupled

47

with the operating O/F ratio of the motor. In all cases the viability of industrially produced ABS

fuel grains has been clearly proven. Future studies will emphasize methods for increasing the

burn efficiency of the ABS fuel material formulation.

Acknowledgments

This work was partially funded under a grant from the Utah Governor’s Office of Economic

Development Technology Commercialization and Innovation Program (TCIP). Additional

funding was received from the NASA Experimental Program for Stimulating Competitive

Research (EPSCOR). The authors want to thank teaming partner Avridyne Technologies/Rocket

Crafter’s, Inc of Layton Utah for procuring the ABS fuel grains from Stratus, Inc.

References

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