HTPB composite propellant

Combustion and Flame 161 (2014) 363–369

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Combustion and Flame j o u r n a l h o m e p a g e : w w w . e l s e v i e r . c o m / l o c a t e / c o m b u s t fl a m e

Humidity induced burning rate degradation of an iron oxide catalyzed ammonium perchlorate/HTPB composite propellant Brian A. McDonald ⇑, Jeremy R. Rice, Mark W. Kirkham AMRDEC Weapons Development and Integration Directorate, Redstone Arsenal, AL 35898, United States

a r t i c l e

i n f o

Article history: Received 25 June 2013 Received in revised form 7 August 2013 Accepted 20 August 2013 Available online 16 September 2013 Keywords: Ammonium perchlorate Burning rate Crystallization Humidity Propellant

a b s t r a c t Burn rate degradation of ammonium perchlorate based solid propellants can occur when moisture diffuses in and out of the material while exposed to fluctuating ambient humidity conditions. For high burn rate propellants with significant mass fractions of ammonium perchlorate particles less than 10 lm, small changes in particle diameter can significantly alter the total oxidizer surface area resulting in propellant burning rate reductions. Ammonium perchlorate propellant samples are aged at various relative humidity and constant temperature. The samples are subsequently dried to equal moisture content and examined by SEM and optical microscopy. The propellant samples are combusted in a closed combustion bomb to measure the burning rate of the aged samples. The results show a clear correlation of the burning rate degradation and the level of humidity exposure. Evidence indicates that the degradation is a result of ammonium perchlorate crystal size growth and surface morphology changes reducing the available surface area. The changes are shown to be correlated with the moisture content of the aging environment. Published by Elsevier Inc. on behalf of The Combustion Institute.

1. Introduction Ammonium perchlorate (AP) is the most predominantly used oxidizer in composite solid propellants and is often used in application where ambient humidity cannot be tightly controlled. Weather seals in rocket motor nozzles slow moisture diffusion but cannot completely prevent the intrusion and thus in time, the propellant may be exposed to significant water vapor. With AP based composite propellants, the burning rate can be formulated to be between approximately 0.64 and 3.3 cm/s at 6.89 MPa by simply varying the AP particle size [1]. Typical AP particle sizes range from 3 lm (ultrafine) up to 600 lm (course) and are usually blended in bimodal and trimodal mass ratios of course–medium– fine combinations. A moderate burning rate composite propellant with a hydroxyl terminated polybutadiene (HTPB) binder may contain 400 lm and 50 lm with a bimodal ratio of 70/30, while a high rate propellant may contain 90 lm and 3 lm with a bimodal ratio of 60/40. Numerous researchers have investigated the effect of AP particle size on the burning rate of composite propellants. Beckstead, Derr and Price [2,3] postulated a flame structure for AP oxidized solid propellants which demonstrates the relationship between the propellant’s burning rate and the AP particle size. At high pressure the flame structure consists of ammonium perchlorate ⇑ Corresponding author. Fax: +1 256 842 2079. E-mail address: [email protected] (B.A. McDonald).

condensed phase reactions at the solid interface, a premixed monopropellant flame consisting of reactions between AP decomposition products, a primary diffusion flame at the initial interface of the binder gasses and the AP monopropellant flame, and the final diffusion flame between binder gasses and AP decomposition products. The burn rate controlling mechanism is considered to be the primary diffusion flame that occurs on the edges of the AP monopropellant flame. These edge flamelets are of intense heat release and are in close proximity to the pyrolyzing surface and thus account for the majority of the heat transfer back to the surface. Given that the primary rate controlling flame is a diffusion flame, it is diffusion length dependent with the diffusion length scale being the AP particle diameter or distance between the monopropellant edges thus tying the burning rate to the AP particle size. Additionally, since the monopropellant flame structure is strongly pressure dependent, Beckstead shows with the model that as pressure falls the flame lifts from the surface and eventually cannot be established at low pressure with only the primary diffusion flame remaining. More recent multiphase numerical models with finite rate chemistry have been developed [4,5] and have demonstrated the burning rate dependency on particle size. Several experimental studies have examined the burning rate and AP particle size relationship. For composite propellants containing no burning rate catalyst, several factors can alter or affect the burning rate in addition to the particle size, to include the AP packing fraction and the oxidizer to fuel mass ratio. For multimodal particle size formulations, the burning rate has been shown

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to be dependent on the particle size mass fraction ratios even at a constant mass weighted particle diameter. Rodic and Bajlovski [6] investigated AP-composite propellant trimodal mixture formulations showing the dependency of the propellant mechanical properties, density, and burning rate on the AP mixture ratios. Banerjee and Chakravarthy [7] have examined bimodal AP formulations containing various coarse to fine particle size ratios and have demonstrated their affect on plateau burning rate trends. Charkravarthy et al. [8] examined intermittent burning and its contribution to plateau burning for various AP particle size blends as well. All three investigations, including the work by Al-Harthi and Williams [9], reinforce the strong dependency of burning rate on particle size. Of particular interest are two investigations by Kohga [10,11] who did extensive experimental work on the burning rate characteristics of monomodal and bimodal AP-composite propellants. He has investigated the burning rate dependency for constant mass weighted diameter formulations with a range of bimodal particle size combination and packing fractions. He further investigated the burning characteristics for bimodal blends using fine porous, fine hollow and spherical AP. Each formulation is at a constant AP mass fraction of 72%. SEM and the BET (Brunauer–Emmett–Teller) method [12,13] are used to determine the mass weighted particle diameter and mass weighted particle surface area for the non-spherical AP. As expected, the resulting burning rates of the formulations exhibited strong dependencies on the mass weighted diameters and surface areas. However, the results show that the strength of this dependency is not consistent over all ranges of particle sizes. At constant pressure and constant oxidizer mass fraction he finds a critical AP particle diameter of 19 lm below which the burning rate is strongly dependent on the particle size, and above which the rate is mildly dependent or even invariant with particle size. A similar functionality is seen with the particle specific surface area as opposed to weighted diameter. AP is a hygroscopic inorganic salt exhibiting solubility in water of 11.6 g/100 ml. Particle size analysis shows that ground AP tends to agglomerate if exposed to high humidity levels, with the sample mean particle diameter increasing with time. Several researchers have examined the crystal growth of AP precipitated from a supersaturated AP-water mixture using various crystal growth and morphology modifiers. In most cases, the objective of these investigations was to produce a substantial quantity of AP by crystal growth with mean particle diameters less than 20 lm, which is approximately the limit of standard grinding methods such as the ball mill. Kohga and Tsuzuki [14] used ethylene glycol as a crystal habit modifier and were able to reduce the mean crystal size produced from the supersaturated solution from 213 lm down to 142 lm. Tanrikulu et al. [15] used salting-out crystallization but found that particles less than 300 lm were difficult to produce. Superfine AP has been produced by Ma et al. [16] via ceramic membrane anti-solvent crystallization. Their research produced crystals down to approximately 3 lm with morphology of polyhedral, quadrate, and rod-shaped depending on the solution feed pressure and temperature. Vargeese et al. [17] state that AP crystals grown from its saturated solution tend to be needle-like. Using poly(vinyl alcohol) as a crystal modifier, they produced crystals in the size range of 80–500 lm with morphologies of rectangular and wedge shaped prisms. A commonality of these investigations indicate that in the absence of crystal growth modifiers or specialized methods, AP crystals formed from its saturated solution are of irregular morphology and tend toward mean particle diameters much greater than 20 lm. The burn rate of typical composite solid propellants oxidized with AP have historically been viewed as practically invariant with moderate exposure to relative humidity levels of less than 50%. These propellants, in general, will contain AP particles of 20 lm

or larger. Several recent rocket motor tests of high burning rate propellant containing a significant mass fraction of superfine AP (3 lm) exhibited substantial burning rate reductions, extended action times, and reduced thrust and pressure. An investigation revealed that the environmental protection system of these motors failed, allowing for exposure to ambient humidties from 30% to 75% for several months. Based on a review of research in AP crystal growth and the known tendency for agglomeration, the burning rate reduction is believed to be caused by an increase in the mean AP particle diameter via re-crystallization, and thus a reduction in available AP surface area. Presented herein are the results of an investigation whereby propellant samples are exposed to a range of relative humidity levels, examined by optical microscopy and SEM and burned in a closed combustion bomb to measure the post-aged burning rate. The results show a clear correlation of burning rate degradation with the aging humidity when dried and burned at equal residual moisture.

2. Propellant aging experiment The propellant chosen for this investigation is an aluminized high burning rate composite propellant with 88% total solids loading using a hydroxyl terminated polybutadiene binder. The mass ratios of the AP, the aluminum, and the binder are 0.74, 0.12, and 0.095 respectively. Plasticizer and minor ingredients constitute the remainder of the propellant formulation. The AP particle size blend consists of a 60/40 mass ratio of 90 lm and 3 lm respectively. All propellant used in the investigation is of equal age and prior to the accelerated aging experiment is stored in identical conditions of approximately 30–35% relative humidity and 21 °C. To facilitate the surface regression calculations for the closed combustion bomb tests, the aging samples are prepared as right circular cylinders with nominal diameters and heights of 8 mm and 3.5 mm respectively. The propellant samples are placed in glass desiccators on top of a perforated ceramic plate. The humidity in each container is controlled by a saturated salt solution [18] or a drying agent. Table 1 shows the selected aging salts/agents and the corresponding controlled relative humidities. The desiccators are placed in an oven and aged for 10 days at 70 °C. All samples are then removed from the humidity control desiccators and held in a container for 10 days in the presence of a 4 Å molecular sieve to reduce the samples to equal residual moisture content. Table 2 shows the sample masses prior to and after aging in the humidity controlled containers. The mass after aging shows very little change in all cases with the exception of those samples aged at 100% RH. This indicates that the mass of water retained in the sample at equilibrium is very small compared to the sample mass. Those samples aged at 100% RH show a substantial mass loss at the end of aging. The lost mass corresponds closely to the initial mass fraction of AP in the propellant formulation indicating that the AP leached from the propellant sample leaving predominantly binder and aluminum. Figure 1 shows a 30 optical microscopy image of the propellant surface prior to aging. This level of magnification minimally distinguishes the 90 lm AP particles on the cut plane. Figures 2–5 show 30 microscopy images of the propellant surfaces after 10 days Table 1 Humidity control salts and drying agents. Salt/agent

20 °C

30 °C

40 °C

60 °C

70 °C

Molecular sieve (4 Å) MgCl2–6H2O NaCl H2O

<2 33.1 75.5 100

<2 32.4 75.1 100

<2 31.6 74.7 100

<2 29.3 74.5 100

<2 27.8 75.1 100

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Table 2 Propellant sample masses before and after humidity aging. Sample number

Humidity

Initial mass (g)

Mass after aging (g)

1 2 3 4 5 6 7 8 9 10 11 12

0. 0. 0. 27.8 27.8 27.8 75.1 75.1 75.1 100.0 100.0 100.0

0.156 0.154 0.158 0.155 0.232 0.218 0.157 0.166 0.158 0.154 0.159 0.155

0.155 0.153 0.159 0.156 0.232 0.218 0.157 0.165 – 0.051 0.056 0.054

1000 µm

Fig. 4. 30-Optical microscopy of post aging propellant surface – aged at 75.1% RH.

Rod Shaped Growth

AP Crystal

1000 µm 1000 µm

Fig. 5. 30-Optical microscopy of post aging propellant surface – aged at 100% RH.

Fig. 1. 30-Optical microscopy of pre aging propellant surface.

1000 µm

aging at 70 °C at their respected controlled humidities. No obvious distinction can be made in the images up to the humidity level of 75%. The image of the surface aged at 100% shows very clear recrystallization of the AP where it is growing rod-shaped features from multiple nucleation sites. SEM images of the post-aged surface of the 100% RH sample are shown in Figs. 6–9 with magnifications of 40 up to 3000. The 40 image shows extensive area that has no particulate on the surface but web-like structures that are consistent with the optical image in Fig. 5. The 150 image seen in Fig. 7 shows the appearance of a range of particle sizes to include several that appear close to the nominal large particle size of 90 lm. However, the image also indicates a spread of sizes that seem to fall between the 3 lm and the

Fig. 2. 30-Optical microscopy of post aging propellant surface – aged dry.

1000 µm Fig. 3. 30-Optical microscopy of post aging propellant surface – aged at 27.8% RH.

Fig. 6. 40-SEM of propellant surface aged at 100% RH.

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3. Propellant burning rate evaluation 3.1. Closed combustion bomb experiment

Fig. 7. 150-SEM of propellant surface aged at 100% RH.

Fig. 8. 600-SEM of propellant surface aged at 100% RH.

The burning rate of the aged propellant samples is measured in a 10 cc closed combustion bomb. The samples are held in a desiccated container until each individual sample is combusted such that the residual moisture in the propellant is held to a constant and equal level. Each sample is ignited using a SARC-1000 squib of approximately 80 mg in mass. Pressure data is collected at a sampling rate of 50 kHz and is shown for all samples in Fig. 10. In closed bomb data the peak pressure is primarily determined by the balance of energy generation and heat loss and is thus mostly affected by the propellant mass for samples of like formulation. Several key features are seen in the data of Fig. 10 to include the obvious squib pressure spike of approximately 1400 psi which is seen in all firings. Of note also is the pressure history for the sample aged at 100% RH. As expected, there is insufficient oxidizer remaining in the binder to support combustion and thus the pressure trace decays from the squib spike at the rate determined by the heat loss. Also noted in the figure is the point of web burn out depicted as the peak of the pressure trace. The rate of heat generation from combustion is assumed to far exceed the heat loss rate and thus the pressure will continue to rise as long as mass addition from combustion is occurring. When the trace reaches its peak and begins to decay the propellant chip is assumed to be burned out. The validity of this assumption depends on the surface history of the burning solid. Geometries that progress to burnout with only moderate decreases in surface area should fall reasonably within the bounds of this assumption. However, for geometries that slowly progress to zero surface area or form slivers of surface area may fail to produce sufficient energy to counter heat loss and maintain an increase in the pressure trace for the full duration of combustion. 3.2. Burning rate calculations The burning rate of each sample is derived from the pressure data by casting the conservation equations and the equation of state into a form that allows for the calculation of the instantaneous mass generation rate. The mass generation rate can then be transformed into the instantaneous burning rate provided that the surface regression history for the propellant sample can be

100 75.1 75.1 75.1 27.8 27.8 27.8 Dry Dry Dry

Web Burnout 17,000

Pressure (kPa)

13,600

10,200

6,800

Fig. 9. 3000-SEM of propellant surface aged at 100% RH.

90 lm nominal sizes. Some particles may be partially imbedded in the propellant binder and appear smaller. Of note also is the irregular shape of the crystals. Figures 8 and 9 show the irregular surfaces of the particles to include the eroded surfaces of the large particles. Aluminum particulate with a mean particle diameter of 20 lm is also present in the formulation and will account for some of the variation in size as witnessed in the SEM image.

3,400

0 0.10

0.11

0.12

0.13

0.14

0.15

0.16

0.17

Time (sec) Fig. 10. Closed combustion bomb data for all samples.

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tracked. Eqs. (1) and (2) show the rate forms of the ideal gas equation of state and the energy equation respectively.

30

@ @ ðPVÞ ¼ ðmRTÞ @t @t

25

ð1Þ Burning Rate (cm/s)

@ _ g cp T f  Q_ ðmcv TÞ ¼ m @t

ð2Þ

Because there are no outlets to the closed combustion bomb, all time derivatives of mass in the bomb are uniquely equal to the _ g . Recognizing that the rate of change gas mass generation rate of m of the internal volume is related to the rate that solid mass is converted to gas, Eqs. (1) and (2) can be combined to form a single expression in terms of the unknown gas mass generation rate as is shown Eq. (3), where the dotted accents denote derivatives with respect to time.

_  mRT _ þ RQ_ þ mRT c_ v PV cv cv

cRT f  qPp

ð3Þ

In the equations, P is pressure, V is volume, m is gas mass, R is the gas constant, T is the gas temperature in the volume, Tf is the flame temperature, qp is the solid propellant density, and Q is the heat loss. The time derivatives of the gas constant and the specific heat at constant volume are evaluated from the rules of the mixture of gasses only and do not account for any gas phase reactions. The instantaneous rate of change of pressure is found by numerical differentiation of the measured pressure data. Since the propellant samples are shaped as right circular cylinders, the theoretical surface regression can be expressed as in Eq. (4) as a function of the burn distance w. From the surface area and the mass generation rate in Eq. (3), the burning rate can be found from Eq. (5). The continuity of the calculation is completed by tracking the burn distance via the integral shown in Eq. (6).

Sw ¼ 2pðr  wÞ2 þ 2pðr  wÞðh  2wÞ rðpÞ ¼

_g m

q p Sw

wðtÞ ¼

Z

rðpÞdt

ð4Þ ð5Þ

ð6Þ

The critical unknown in Eq. (3) is the heat loss to the chamber walls which can be substantial in heavy walled and uninsulated closed combustion bombs. Clearly the rate of heat loss will not be evenly distributed across the burn time with the rate highest at the onset when the walls are at ambient. For the present work, the heat loss term in Eq. (3) is treated as constant and is found by iteration for each dataset. Since the propellant sample is assumed to burn out at the peak of the pressure curve the heat loss term can be adjusted such that integral of Eq. (6) equals the full web thickness of the individual sample at the appropriate time. Since the heat loss is treated as a constant, the results are expected to produce better predictions in the middle sections of the combustion time with deviations at the beginning and end. The validity of this approach is demonstrated in Fig. 11. Unaged propellant samples are combusted in the closed bomb and analyzed with the constant heat transfer approach. Other samples are prepared into strands and are burned in a traditional propellant strand burner over a range of pressures. The figure shows that the approach produces excellent agreement to strand data with the exception of the lowest pressure levels which correspond to the beginning of combustion as expected. The data sets begin to show a trend of divergence at the highest pressures where the walls are warm and the constant heat transfer approach over predicts the heat loss. Several attempts to develop a wall temperature dependent model were made but the results were

20

15

10

5

0 20000

40000

60000

80000

100000

Pressure (kPa) Fig. 11. Comparison of strand burn rate data to calculated rates derived from bomb data.

not convincing. Refinement of the approach generally adds additional unknowns that cannot be uniquely solved for. The results of the closed combustion bomb data analysis for the aged propellant samples are shown in Fig. 12. The slope for each curve is reasonably the same, but a clear offset in magnitude as a function of the aging humidity is evident. The data for the unaged samples and those aged in a dry environment align together in the figure and have the highest burning rate, while the samples aged at 75% RH have the lowest rates. Figure 14 shows this more clearly where the burning rate at 6895 kPa for each sample is plotted versus the aging humidity condition. There is sufficient grouping of the data points at each humidity condition to validate that a correlation between the level of the degradation of burning rate and humidity exists. Data for the 100% RH samples is not shown in the figure because most of the AP leached to the surface with leaving insufficient oxidizer within the binder to support combustion. The data in Fig. 12 is assumed to obey the power law burn rate equation of r = cPn, where n is referred to as the burning rate exponent. Using regression analysis, a power law curve fit is applied to the data in Fig. 12 with the exponents of the fit shown in Fig. 13. The figure shows a linear regression through the exponent values as well as upper and lower 95% confidence bands. The linear regression is of poor fit primarily because of the outlier data point at 75% RH. This point is a result of the data anomaly seen on the bottom curve (75% RH) in Fig. 12. However, even with this point included the exponent data shows little evidence of a clear trend as a function of humidity.

6.00

5.00

Burning Rate (cm/s)

_g¼ m

Strand Data Bomb Data

Unaged Uaged Dry Dry Dry 28% 28% 28% 75% 75% 75%

4.00

3.00

2.00 4000

6000

8000

10000

Pressure (kPa) Fig. 12. Propellant burning rates derived from combustion bomb data.

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0.95

Burn Rate Exponent

0.90 0.85 0.80 0.75 0.70 0.65 0.60 0

10

20

30

40

50

60

70

80

Relative Humidity Fig. 13. Burning rate exponents of aged samples.

4.19

Burning Rate (cm/s)

4.06 3.94 3.81 3.68 3.56 3.43 3.30 0

10

20

30

40

50

60

70

80

Relative Humidty (%) Fig. 14. Burning rate at 6895 kPa versus aging humidity.

Because all propellant samples are dried after aging to the same moisture content prior to burning rate testing, the witnessed degradation in burning rate represents a permanent change in the material. The test data also shows that the level of this change is proportional to the moisture content of the aging environment. Certainly the equilibrium moisture content of the samples is determined by the balance of the partial molar free energy across all phases [19]. Since all samples are aged at the same temperature, the difference in equilibrium moisture content is directly proportional to the humidity. The SEM and the optical microscopy are inconclusive for all cases with the exception of the 100% RH samples. In the 100% RH samples both particle size and morphology changes are evident. An examination of the data in Fig. 12 shows that over most of the pressure range considered in testing, the burn rate curves at different humidities do not converge. This indicates that the alteration of the burning rate is affected through the thickness and not simply a surface phenomenon. At the higher pressures the burning rate for the 75% RH samples begin to drastically increase which may indicate that the inner most particles of these samples are not fully affected by the moisture intrusion. Hydrolysis reactions between the polymer binder and the absorbed moisture could lower the heat of combustion, however rocket motors that have experience similar degradation when exposed to moisture typically show invariance in the delivered total impulse as the burning rate degrades. The propellant in this study uses iron oxide as a burning rate catalyst giving the propellant the red color. The only conceivable Fe2O3–water reaction would be to form Fe3O4 however the

literature shows no propensity for formation of Fe3O4 by this mechanism. In addition Fe3O4 is black and would alter the appearance of the propellant. One possible explanation of the degraded burning rate is that the sensible enthalpy of the combustion products is reduced by residual moisture in the sample not removed in the drying step, or by hydrolysis reactions with the binder during aging. In either case the flame temperature would be reduced causing a reduction in the burning rate. Figure 15 show the normalized pressure integral and the normalized maximum pressure for all samples tested. The data is normalized by the sample mass and plotted as a function of the relative humidity of aging. The maximum pressure normalized by the mass is essentially the gas temperature at maximum pressure. Neither curve shows a correlation with humidity; however variation in the data is evident which is not intuitive given that the samples are from the same propellant mix. Figure 16 shows the pressure integral and the maximum pressure plotted as a function of the sample mass. A very clear correlation exists with the highest pressures seen at maximum mass as would be expected in a closed bomb and thus accounts for the variation seen in Fig. 15 given that the sample masses do not correlate with the aging humidity. The lack of a clear correlation in Fig. 15 indicates that residual water or aging related reactions that lower the heat of combustion do not account for the degradation in burning rate. Kohga [10] generated burning rate data versus mean AP particle diameter for a bimodal blend which is shown in Fig. 17. The Kohga data is taken at 7000 kPa and is presented as a percent change in burning rate from the maximum rate measured as a function of the mass weighted particle size for the bimodal blend. Assuming that the bimodal propellant used in the aging study will follow a similar relationship of percent change in burning rate versus mean particle diameter, the aging data at 6890 kPa in Fig. 14 is plotted with the Kogha data. The base burning rate of the data in Fig. 14 is slightly higher than the base rate of the Kogha data for a direct comparison. The unaged mean particle diameter of the propellant is 55.2 lm (60/40 mass ratio of 90 lm and 3 lm). Following the Kogha trend, the mean particle diameter would need to increase to approximately 79.4 lm to account for the degradation in burning rate seen at an aging relative humidity of 75%. For a 60/40 mass ratio of 90 lm and 3 lm AP, the particle count ratio of small to large is 18,000:1 with a surface area ratio of 20:1. Therefore the available surface area of the large particles for water contact is negligible as compared to the small particle. Assuming then that the mean particle diameter shift is entirely effected by a change in

10600

300

10400 250

10200 10000

200

9800 Unaged Samples

9600

150

9400 100

9200 9000

50 Pressure Integral Maximum Pressure

8800

0

Maximum Pressure per Unit Mass (kPa/g)

1.00

Pressure Integral per Unit Mass (kPa-s/g)

368

8600 0

10

20

30

40

50

60

70

80

Relative Humidity (%) Fig. 15. Normalized maximum pressure and pressure integral versus relative humidity.

B.A. McDonald et al. / Combustion and Flame 161 (2014) 363–369 16000

450

15000 14000 350 13000 300

12000 11000

250 10000

Maximum Pressure (kPa)

Pressure Integral (kPa-s)

400

200 9000

Pressure Integral Maximum Pressure

150 0.12

0.14

0.16

0.18

0.2

0.22

8000 0.24

Mass (g)

369

of humidity. The samples are dried to equal moisture content and examined with SEM and optical microscopy. The images of the surfaces show extensive alteration of the particle size and morphology of the samples aged at 100% RH, but at lower humidities the images are inconclusive. These samples are combusted in a closed combustion bomb whereby the propellant burning rate is shown to degrade as a function of the level of the humidity in the aging environment. The degradations appear to be a permanent condition that permeates the bulk of the material. Analysis of the data shows that chemical changes and or residual moisture in the propellant are unlikely explanations for the degradation of the rates. The rate change is consistent with changes that would occur if the mean particle size of the ammonium perchlorate is increased. This mean particle size increase reduces the available AP surface area with a corresponding decrease in burning rate. Acknowledgements

Fig. 16. Maximum pressure and pressure integral versus sample mass.

The authors would like to acknowledge the contribution of Dr. John V. Foreman, Weapons Science Directorate, AMRDEC for producing the SEM images.

Kogha Data (Ref 10) 28% RH Aging 75% RH Aging

0.6

Percent Change in Burning Rate

References 0.5

0.4

0.3

0.2

0.1

0.0 0

20

40

60

80

100

120

140

Mass Weighted Mean Diameter (µm) Fig. 17. Percent change in burning rate versus mean particle diameter for a bimodal blend.

the diameter of the small particles, the 40% by mass 3 lm particles would need to agglomerate to an average diameter of 63.5 lm. This level of agglomeration or re-crystallization is not inconceivable and is consistent with data for particles formed from a saturated solution without growth modifiers as presented in Refs. [14–17]. Furthermore, the high particle count of the small particulate packed in the interstitial space of the large would readily allow for the diffusion of material with substantial impedance of the binder. Alternatively, the ratio of large to small particulate may increase by deposition by evaporation of the AP dissolved in the resident moisture onto the large particulate. This would tend to grow the large particles and decrease the quantity of small particles. However, as the mass ratio of large particles to small increases the packing fraction would also be expected to change. The propellant volume may grow as the ratio of course to fine AP increases, or the volume may decrease if the precipitating AP fills the interstitial spaces. Sample density changes were not measured in this study. 4. Conclusions Propellant samples containing very fine ammonium perchlorate particle sizes are aged at constant temperature and various levels

[1] A.O. Oberth, Principles of Solid Propellant Development, CPIA Publications, 1987, p. 468. [2] M.W. Beckstead, R.L. Derr, C.F. Price, A model of composite solid propellant combustion based on multiple flames, AIAA J. 8 (1970) 12. [3] M.W. Beckstead, Solid propellant combustion mechanisms and flame structure, Pure Appl. Chem. 65 (1993) 2. [4] Weidong Cai, Piyush Thakre, Vigor Yang, A model of AP/HTPB composite propellant combustion in rocket-motor environments, Combust. Sci. Technol. 180 (2008). [5] R.P. Fitzgerald, M.Q. Brewester, Flame and surface structure of laminate propellants with coarse and fine ammonium perchlorate, Combust. Flame 136 (2004). [6] Vesna Rodic’, Miomir Bajlovski, Influence of trimodal fraction mixture of ammonium perchlorate on the characteristics of composite rocket propellants, Sci. Technol. Rev. LVI (2006) 2. [7] S. Banerjee, S.R. Chakravarthy, Ammonium perchlorate-based composite solid propellant formulation with plateau burning rate trends, Combust. Expl. Shocks 43 (2007) 4. [8] S.R. Chakravarthy, J.M. Seitzman, R. Lillard, E.W. Price, R.K. Sigman, Intermitten burning and its contribution to plateau burning of composite propellants, in: 39th Aerospace Sciences Meeting and Exhibit, Reno, AIAA, 2001. [9] A. Al-Harthi, A. Williams, Effect of fuel binder and oxidizer particle diameter on the combustion of ammonium perchlorate based propellants, Fuel 77 (1998) 13. [10] Makoto Kohga, Burning rate characteristics of ammonium perchlorate-based composite propellant using bimodal ammonium perchlorate, J Propul Power 24 (2008) 3. [11] Makoto Kohga, Burning characteristics of AP/HTTP composite propellants prepared with fine porous or fine hollow ammonium perchlorate, Propellants, Explos., Pyrotech. 31 (2006) 1. [12] S. Brunauer, P.H. Emmett, E. Teller, Adsorption of gases in multimolecular layers, J. Am. Chem. 60 (1938). [13] S. Brunauer, I. Odler, M. Yunenfreund, The new model of hardened portlant cement paste, Highway Restoration 328 (1970). [14] Makoto Kohga, Hirotatsu Tsuzuki, Crystal habit modification of ammonium perchlorate by ethylene glycol, Adv. Powder Technol. 21 (2010). [15] S.U. Tanrikulu, I. Eroglu, A.N. Bulutcu, S. Ozkar, The growth and dissolution of ammonium perchlorate crystals in the fluidized bed crystallizer, J. Cryst. Growth 194 (1998). [16] Z. Ma, C. Li, R. Wu, P. Chen, Z. Gu, Preparation and characterization of superfine ammonium perchlorate (AP) crystals through ceramic membrane anti-solvent crystallization, J. Cryst. Growth 311 (2009) 21. [17] Anuj A. Vargeese, Satyawati S. Joshi, V.N. Krishnamurthy, Role of poly(vinyl alcohol) in the crystal growth of ammonium perchlorate, Cryst. Growth Des. 8 (2007) 3. [18] Lewis Greenspan, Humidity fixed points of binary saturated aqueous solutions, J. Res. Natl. Bur. Stand. 81A (1976) 1. [19] V.M. Kazanskii, The equilibrium humidity of dispersed solids at different temperatures, Kiev Inst. Eng.-Constr. 20 (1971) 5.