Progress in Modeling Ignition in a Solid Propellant Charge 

for Telescoped Ammunition 

by Michael J. Nusca and Albert W. Horst 

ARL-TR-3673 November 2005 

Approved for public release; distribution is unlimited. 
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Army Research Laboratory 

Aberdeen Proving Ground, MD 21005-5066 

ARL-TR-3673 November 2005 

Progress in Modeling Ignition in a Solid Propellant Charge 

for Telescoped Ammunition 

Michael J. Nusca and Albert W. Horst 

Weapons and Materials Research Directorate, ARL 

Approved for public release; distribution is unlimited. 
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Progress in Modeling Ignition in a Solid Propellant Charge for Telescoped 

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13. SUPPLEMENTARY NOTES 

14. ABSTRACT 

One of the many challenges facing weapon developers is the requirement for a highly lethal, lightweight, and compact largecaliber gun system. A promising concept recently investigated by the U.S. Army is that of a swing-chamber gun, necessitating 

the use of telescoped ammunition. Such ammunition not only reduces the volume available for the propellant charge, but also 

places severe geometric constraints on both the distribution of the propellant and the location and functionality of the ignition 

system. Results of an earlier study highlighted the fact that lumped-parameter interior ballistic codes cannot capture the 

influence of these configurational complexities on the processes of flamespreading and the ensuing formation of pressure 

waves. Application of a one-dimensional, two-phase flow code to this problem revealed the likelihood of such waves and 

raised concern over possible damage to the projectile. Subsequent use of a state-of-the-art, multidimensional interior ballistic 

code provided quantitative predictions of the flow in the annular region between the sidewall of the telescoped projectile and 

the cartridge case, detailed the formation of pressure waves, and furthered concern about transient projectile loads. The present 

report extends this effort, providing results applicable both to comparison with companion gun simulator experiments and 

appropriate for coupling to projectile/gun structural dynamics codes. 

15. SUBJECT TERMS

gun charges, solid propellant, multiphase flow, computer simulation 

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19a. NAME OF RESPONSIBLE PERSON

Michael J. Nusca 

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iii

Contents 

List of Figures iv

1. Introduction 1

2. Current Approach 5

3. Results of Calculations 8

3.1 IBHVG2 Simulations ......................................................................................................8 

3.2 XKTC Simulations..........................................................................................................8

3.3 NGEN3 Simulations......................................................................................................11 

4. Conclusions and Future Efforts 16

5. References 19

Distribution List 21 
iv

List of Figures 

Figure 1. Schematic of simplified telescoped ammunition configuration (1). ...............................1 

Figure 2. Unstructured chamber (XKTC) (1). ................................................................................2 

Figure 3. Base charge (XKTC) (1). ................................................................................................2 

Figure 4. Annular charge (XKTC) (1). ...........................................................................................2 

Figure 5. Computational regions (NGEN3) (3). .............................................................................3 

Figure 6. Full charge simulation (NGEN3) (3)...............................................................................3 

Figure 7. Predicted gas pressure contours (blue to red: 0.1–350 MPa) and velocity vectors 

for full charge and 1.2–3 ms (NGEN3) (3)................................................................................4 

Figure 8. Predicted reduction in pressure waves for full charge of figure 6 with null axial 

interphase drag in the annular region of propellant (NGEN3) (8).............................................4 

Figure 9. Typical effective stress in projectile wall predicted by DYNA3D using NGEN3 

pressure data (8).........................................................................................................................5 

Figure 10. Schematic of simulator used in companion experiments (13).......................................5 

Figure 11. (a) Rear propellant loading; (b) shell propellant loading; and (c) full chamber 

propellant loading. .....................................................................................................................6 

Figure 12. (a) Rear or shell loading (IBHVG2) and (b) full chamber loading (IBHVG2).............8 

Figure 13. (a) Rear loading (XKTC); (b) shell loading (XKTC); and (c) full chamber loading 

(XKTC)......................................................................................................................................9 

Figure 14. Expanded low pressure regions for figure 13 data (XKTC)........................................10 

Figure 15. (a) Rear loading (NGEN3); (b) shell loading (NGEN3); and (c) full chamber 

loading (NGEN3).....................................................................................................................12 

Figure 16. (a) Low pressure region for figure 15a; (b) low pressure region for figure 15b; 

and (c) low pressure region for figure 15c...............................................................................13 

Figure 17. Predicted gas pressure contours (blue to red: minimum to maximum custom scale 

for each figure) and velocity vectors, rear charge (corresponding to figures 15a and 16a): 

0.5–4.7 ms (NGEN3). ..............................................................................................................14 

Figure 18. Predicted gas pressure contours (blue to red: minimum to maximum custom scale 

for each figure) and velocity vectors, shell charge (corresponding to figures 15b and 16b): 

0.5–5.5 ms (NGEN3). ..............................................................................................................14 

Figure 19. Predicted gas pressure contours (blue to red: minimum to maximum custom scale 

for each figure) and velocity vectors, full loading (corresponding to figures 15c and 16c): 

0.5–4.7 ms (NGEN3). ..............................................................................................................14 

Figure 20. (a–b) Low pressure regions for simulation with granular drag; rear loading (left) 

and shell loading (right) (NGEN3); and (c) low pressure region for granular drag; full 

chamber loading (NGEN3)......................................................................................................17
1

1. Introduction 

U.S. Army interest in compact and lightweight future combat systems with maximum 

capabilities in terms of both lethality and survivability has led to consideration of a number of 

nonstandard weapon systems and armor approaches. One of the weapon concepts considered has 

been a compact high-performance, large-caliber gun employing telescoped ammunition as a 

means of reducing ammunition, and hence overall weapon size and weight. Challenges, in terms 

of ignition, flamespreading, and attendant pressure-wave formation, as well as overall 

performance, have motivated efforts at the U.S. Army Research Laboratory (ARL) to model the 

interior ballistic process for such configurations in increasing detail. 

An initial effort (1) employed an existing one-dimensional (with area change), two-phase flow 

interior ballistic code known as XKTC (2) to investigated the influence of a nominal, but 

complex, telescoped ammunition design on flamespreading and pressurization profiles obtained 

with propellant loaded in either base or base and annular regions of the chamber (see figure 1). 

A very brief review of results from this preliminary study displays the influence of chamber 

configuration and charge location on the nature of pressurization within the gun chamber, within 

the limits of this quasi-one-dimensional representation. Figure 2 presents computational results 

for an unstructured chamber configuration (lumped-parameter analogy) revealing the expected, 

well-behaved pressurization at rear (solid line) and forward (dashed line) ends of the chamber. 

Figure 3 shows the resulting pressure-time curves with propellant loaded only in the region to the 

rear of the annulus. Figure 4 provides results when propellant is loaded in both rear and annular 

regions. Figure 3 results were attributed to the change in cross-sectional flow area at the base of 

the telescoped projectile, leading to rapid increases in local pressures, with subsequent stagnation 

a flow reversal when the pressure front reached the front boundary of the annulus. This 

undesirable situation was only exacerbated when propellant was loaded in the annulus as well, 

the local feedback of locally high pressures into locally high burn rates and hence even higher 

pressures leading to a rapid over pressurization of the gun chamber, as indicated in figure 4. 

Figure 1. Schematic of simplified telescoped ammunition 

configuration (1).
2

Figure 2. Unstructured chamber (XKTC) (1). 

Figure 3. Base charge (XKTC) (1). 

Figure 4. Annular charge (XKTC) (1).

Time (ms) 

 Time (ms) 

Time (ms) 
3

A second effort (3) extended the analysis though application of the NGEN3 multidimensional, 

two-phase flow interior ballistic code (4–6) to this problem. This code was used to provide a full 

two-dimensional (2-D), axisymmetric representation of a nominal, degenerate telescoped 

ammunition configuration (figure 5) with four distinct regions of charge, including two for 

ignition materials (regions I and II) at the extreme rear of the chamber and two for main charge 

propellant (region III behind the projectile and region IV in the annular region adjacent to the 

sidewall of the telescoped portion of the projectile). Figure 6 confirms the prediction of strong 

longitudinal pressure waves for such a configuration, while figure 7, for the first time, provides 

details of the pressure fields associated with the development and evolution of these waves. 

Figure 8 demonstrates a substantial reduction in expected pressure waves accompanying the use 

of a low drag propellant configuration (concentric wraps) in the annular region. 

Figure 5. Computational regions (NGEN3) (3). 

Figure 6. Full charge simulation (NGEN3) (3). Pressure (kpsi) 

Pressure (MPa) 

Time (ms) 
4

Figure 7. Predicted gas pressure contours (blue to red: 0.1–350 MPa) and velocity vectors for full charge 

and 1.2–3 ms (NGEN3) (3). 

Figure 8. Predicted reduction in pressure waves for full charge of figure 6 with null 

axial interphase drag in the annular region of propellant (NGEN3) (8). 

Subsequent studies (7, 8) extended the investigation to probe the impact of such highly dynamic 

pressure environments on the correspondingly dynamic response of associated projectile 

structures through the coupling of output from NGEN3 simulations of various telescoped 

ammunition configurations to Lawrence Livermore National Laboratory’s DYNA3D structural 

mechanics code (9) (see figure 9). The negative impact of the presence of strong pressure waves 

and locally high transient pressures on projectile material stress levels and design requirements 

was clearly demonstrated, confirming the need to address the design of telescoped ammunition 

charges and the projectile designs in concert. Pressure (kpsi) 

Pressure (MPa) 

Time (ms) 
5

Figure 9. Typical effective stress in projectile wall predicted by DYNA3D using NGEN3 pressure 

data (8). 

2. Current Approach 

Recent efforts at ARL have focused on exploitation of coordinated modeling and 

experimentation efforts to probe charge/projectile interactions associated with the telescoped 

ammunition configuration. Past efforts employing state-of-the-art interior ballistic codes and 

plastic-chambered gun chamber simulators (10–12) have facilitated the detailed comparison of 

simulated and measured results, with emphasis on identifying causes and controls for 

flamespreading anomalies and ensuing deleterious pressure waves in large-caliber guns. 

A companion paper (13) describes design and initial testing in a large-caliber, telescoped 

ammunition gun simulator, shown schematically in figure 10 (note locations of strain, S, and 

pressure, P, gauges). Of specific interest are two test charge configurations, each employing the 

same ~3-kg charge of 19-perforation, partially-cut JA2 stick propellant, but located in distinctly 

different portions of the chamber: either all in the base region behind the projectile afterbody or 

as a shell of propellant filling the annular region adjacent to the wall of the telescoped portion of 

the projectile and extending to the base of the chamber. Importantly, the projectile has been 

positioned in the chamber at a location to provide the same propellant loading density 

(~0.9 g/cm3

) in the regions of the base charge and the shell charge. Ignition is provided by a 

M123 primer and an accompanying 180-g black powder donut basepad. Moreover, the projectile 

has been locked in position, focusing the experimental study on flamespreading events prior to 

projectile motion. 

Figure 10. Schematic of simulator used in companion experiments (13).
6

In this report, we describe companion calculations from a hierarchy of interior ballistic codes for 

direct comparison to experimental results obtained using the previously described apparatus, and 

optionally, to provide additional input for projectile mechanical response studies employing 

various codes (14). The computational study addresses three loading conditions, as displayed in 

figures 11a–c: rear loading, shell loading, and full chamber. The calculations assume an 

~6-l test chamber affixed to a standard 120-mm tank gun barrel, but with a shot start pressure of 

21 MPa, sufficient to ensure test chamber failure prior to motion of a 22.3-kg projectile, 

essentially allowing simulation of both behavior of flamespreading in the test simulator and the 

full interior ballistic cycle of a nominal 120-mm tank gun. 

 

Figure 11. (a) Rear propellant loading; (b) shell propellant loading; and (c) full chamber propellant loading. 

Three levels of interior ballistic models are applied to the problem: The IBHVG2 lumpedparameter interior ballistic code; the XKTC one-dimensional (with area change), two-phase flow 

interior ballistic code; and the NGEN3 multidimensional, two-phase flow interior ballistic code. 

Briefly, IBHVG2 (15) provides a simple but useful lumped-parameter representation of the 

interior ballistic cycle, embodying such assumptions as uniform and simultaneous ignition of the 

entire propellant charge, with combustion assumed to take place in a smoothly-varying, wellstirred mixture, the burning rate being determined by the instantaneous, space-mean chamber 

pressure. An assumed longitudinal pressure gradient is superimposed on the solution at each 

instant in time to appropriately reduce the pressure on the base of the projectile. An excellent 

tool for estimating overall performance of a gun, study of ignition-induced pressure waves 

(a major concern of this study) is clearly outside the physical scope of this model. Furthermore, 

loading configurations shown in figures 11a and 11b cannot be differentiated in this 

representation. 

Next, the XKTC code (2) provides a quasi-one-dimensional, macroscopic (with respect to 

individual grains), two-phase description of flow in the gun chamber, with the conservation laws 

formulated to neglect the effects of viscosity and heat conduction in the gas phase. Most 

importantly, however, gas and solid phases are coupled through heat transfer, combustion, and 

interphase drag, these processes being modeled using empirical correlations that relate the 

microphenomena to the average flow properties described by the governing equations. The 

igniter is either modeled explicitly or treated as a predetermined mass injection profile, and 

flamespreading follows primarily according to convection, until the ignition temperature is 

reached and combustion follows at a rate determined by the local pressure. Formulated as a 

(a) (b) (c) 
7

one-dimensional with area change representation, XKTC provides a first-level capability for 

treating the dynamics of the axial pressure field and its potential for causing potentially 

damaging overpressures. 

Finally, and of primary interest in the current study, the NGEN3 code (4–6) is a 

multidimensional, multiphase computational fluid dynamics (CFD) code that incorporates threedimensional continuum equations along with auxiliary relations into a modular code structure. 

Since accurate charge modeling involves flowfield components of both a continuous and discrete 

nature, a coupled Eulerian-Lagrangian approach is utilized. On a sufficiently small scale of 

resolution in both space and time, the components of the flow are represented by the balance 

equations for a multi-component reacting mixture describing the conservation of mass, 

momentum, and energy. A macroscopic representation of the flow is adopted using these 

equations derived by a formal averaging technique applied to the microscopic flow. These 

equations require a number of constitutive laws for closure including state equations, 

intergranular stresses, and interphase transfer. The numerical representation of these equations, 

as well as the numerical solution thereof, is based on a finite-volume discretization and highorder accurate, conservative numerical solution schemes. The spatial values of the dependent 

variables at each time step are determined by a numerical integration method denoted the 

continuum flow solver (CFS), which treats the continuous phase and certain of the discrete 

phases in an Eulerian fashion. The Flux-Corrected Transport scheme (16) is a suitable basis for 

the CFS since the method is explicit and has been shown to adapt easily to massively parallel 

computer systems. The discrete phases are treated by a Lagrangian formulation, denoted the 

large particle integrator (LPI), which tracks the particles explicitly and smoothes discontinuities 

associated with boundaries between propellants yielding a continuous distribution of porosity 

over the entire domain. The manner of coupling between the CFS and the LPI is through the 

attribution of properties (e.g., porosity and mass generation) at points in the flow. The size of the 

grid, as well as the number of Lagrangian particles, is user prescribed. The NGEN3 code takes a 

macroscopic approach to solid propellant configuration representation. Solid propellant media 

are modeled using Lagrange particles that regress, produce combustion product gases, and 

respond to gas-dynamic and physical forces. Individual grains, sticks, slab, and wrap layers are 

not resolved; rather, each medium is distributed within a specified region in the gun chamber. 

The constitutive laws that describe interphase drag, form-function, etc., assigned to these various 

media determine preferred gas flow paths through the media and responses of the media to 

forces. Media regions can be encased in impermeable boundaries that yield to gas-dynamic flow 

after a prescribed pressure load is reached. 

Application of the NGEN3 code to solid propellant charges for direct-fire weapons of U.S. Army 

interest is well documented (3, 6, 8, 17–20). Clearly, this code is best suited for treatment of the 

multidimensional aspects of the current problem of interest, and is applied in its latest form to the 

three charge-loading configurations for the telescoped ammunition chamber simulator described 

previously.
8

3. Results of Calculations 

3.1 IBHVG2 Simulations 

As a lumped-parameter code, input data requirements are essentially limited to gun data 

(chamber volume, projectile travel, and barrel resistance profile), projectile mass, and energetic 

material parameters (igniter and main charge masses, dimensions, burning rates, and 

thermodynamic data). Predicted pressure-time data, provided primarily for reference only, are 

computed for two loading conditions only (3- and 4.8-kg main charges). The position of the 

charge in the chamber not being treated by this code, as previously mentioned. The results are 

presented in figures 12a and 12b, displaying the (necessarily) smooth pressure-time curves with 

peak pressures of 154 and 435 MPa. 

Time (ms)

0 2 4 6 8 10 12 14 16 18

Pressure (MPa)

0

20

40

60

80

100

120

140

160

180

Time (ms)

-2 0 2 4 6 8 10 12

Pressure (MPa)

-100

0

100

200

300

400

500

*

Note: Solid line for breech and dashed line for projectile base. 

Figure 12. (a) Rear or shell loading (IBHVG2) and (b) full chamber loading (IBHVG2). 

3.2 XKTC Simulations 

Considerably more interesting are the XKTC results, reflecting the influence of the specific 

positioning and geometry of propellant charges, projectile, and gun (simulator) chamber, within 

the one-dimensional-with-area-change approximation. Moreover, these results are influenced by 

the interphase drag and heat transfer characteristics, as well as bed rheology (i.e., stress fields) 

associated with the formulation and configuration of the propellant—also beyond the scope of 

the lumped-parameter representation. Computational problems prevented some of the 

calculations from going to completion; however, available results for the three configurations are 

presented in figures 13a–13c. While overall maximum chamber pressures deviate somewhat 

from the lumped-parameter results, of particular interest are the increasing wave levels as 

propellant is positioned initially in the confined, annular region adjacent to the projectile, further 

amplified when the rear region is filled as well. Results for the time and pressure regime of 

likely interest for comparison to experimental simulator results are enlarged and displayed in 

figures 14a–14c. 

(a) (b) 
9

Time (ms)

0 2 4 6 8 10 12 14 16

P (MP )

-50

0

50

100

150

200

Rear

Forward

Difference

Time (ms)

0 5 10 15 20

Pressure (MPa) -20

0

20

40

60

80

100

120

Rear

Forward

Difference

Time (ms)

0 2 4 6 8 10 12

Pressure (MPa) -100

0

100

200

300

400

500

Rear

Forward

Difference

Figure 13. (a) Rear loading (XKTC); (b) shell 

loading (XKTC); and (c) full chamber 

loading (XKTC).

(a) 

(b)

(c) 
10

Time (ms)

0123

Pressure (MPa) -5

0

5

10

15

20

25

Rear

Forward

Difference

Time (ms)

0123

Pressure (MPa) -5

0

5

10

15

20

25

Rear

Forward

Difference

Time (ms)

-0.2 0.0 0.2 0.4 0.6 0.8 1.0

Pressure (MPa) -5

0

5

10

15

20

25

Rear

Forward

Difference

Figure 14. Expanded low pressure regions for 

figure 13 data (XKTC).

(a) 

(b)

(c) 
11

3.3 NGEN3 Simulations 

NGEN3 is the first of the codes applied to the present problem that allows a full 2-D, 

axisymmetric treatment of the three loading configurations. Further, while the code does not 

currently support the partially cut multiperforated stick propellant geometry, an excellent 

approximation can be obtained by using the 19-perforated grain option with the axial interphase 

drag nulled (actual decrease is indeed about two orders of magnitude). For comparison, 

calculations were run with both drag levels (i.e., granular and null). 

Figure 15 shows the computed pressure/time curves for the three propellant loading 

configurations, assuming null interphase drag throughout. The plot scales have been chosen to 

correspond to those used in figures 13. Pressure data was collected at rear, middle, and forward 

axial stations along the radial wall of the chamber—at 6, 28, and 50 cm from the chamber breech 

face, respectively. The pressure difference was computed by subtracting the pressure at the 

forward axial station from that of the rear axial station. Comparing figures 13 and 15, there is a 

remarkable correspondence between the XKTC-computed and the NGEN3-computed 

pressure/time curves including the pressure rise rate, the maximum pressure, and the pressure at 

shot exit. Additionally there is a certain degree of agreement between the two codes as to the 

wave dynamics in the chamber as evidenced by the oscillations during pressure rise and the 

pressure differential. Agreement between the codes for results of the amplitude and frequency of 

pressure differentials is especially remarkable for the case of the shell propellant loading. The 

NGEN3 code typically demonstrates some degree of pressure fluctuation near the maximum that 

in part can be contributed to numerical instabilities (17–20); in this case, the pressure 

fluctuations are particularly evident for the full chamber loading (figure 15c). Nevertheless, the 

degree of agreement between the one-dimensional and 2-D simulations lends credence to both. 

Results for the time and pressure regimes of interest for comparison to experimental simulator 

results are enlarged and displayed in figure 16. Again, the plot scales have been chosen to 

correspond to those used in figure 14. Here the degree of correspondence between the XKTC 

and NGEN3 results is less evident. However, certain common patterns can be noted, especially 

for the shell loading configuration (figures 14b and 16b). Recall that the XKTC code is 

modeling the solid propellant charge as partially cut multiperforated sticks while the NGEN3 

code is using a granular propellant model with null interphase drag. Some of the differences in 

pressure results for the early stages of flamespreading can be attributed to this difference, 

especially for the rear loading and the full charge loading configurations (figures 16a and 16c). 

The flamespreading process in each propellant loading configuration can be described in greater 

detail using the NGEN3 code, as demonstrated by the sequence of pressure fields depicting the 

development and evolution of pressure waves in the chamber (figures 17–19). These series of 

computed pressure contours, spanning the time from 0.5 to ~5 ms, are shown in color-scale from 

0.1 MPa (blue) to the maximum pressure for each time snapshot (red). The plots also contain
12

Figure 15. (a) Rear loading (NGEN3); (b) shell loading 

(NGEN3); and (c) full chamber loading

(NGEN3).

(a) 

(c) 

(b)
13

Figure 16. (a) Low pressure region for figure 15a; (b) low 

pressure region for figure 15b; and (c) low 

pressure region for figure 15c.

(c) 

(a) 

(b)
14

Figure 17. Predicted gas pressure contours (blue to red: minimum to maximum custom scale for each figure) and 

velocity vectors, rear charge (corresponding to figures 15a and 16a): 0.5–4.7 ms (NGEN3). 

Figure 18. Predicted gas pressure contours (blue to red: minimum to maximum custom scale for each figure) and 

velocity vectors, shell charge (corresponding to figures 15b and 16b): 0.5–5.5 ms (NGEN3). 

Figure 19. Predicted gas pressure contours (blue to red: minimum to maximum custom scale for each figure) and 

velocity vectors, full loading (corresponding to figures 15c and 16c): 0.5–4.7 ms (NGEN3). 
15

velocity vectors (shown in black) which are overlaid so as to depict both the direction and 

velocity magnitude of gas movement in the chamber (the scaling for the vectors are the same for 

all plots). For each time, the chamber and projectile afterbody are shown from the common 

centerline to the radial wall of the chamber. The locations of the three pressure collection points 

(R, M, F) used in figures 15 and 16 are also shown along the chamber wall for reference. 

The time sequence displayed in figure 17 was computed using the NGEN3 code for the rear 

loading configuration. High pressure gas is generated along the chamber breech face by the 

igniter as gas flow is driven forward through the propellant bed located behind the projectile base 

(0.5 ms). The velocity vectors through this region, between X = 5 and 20 cm, are indicative of 

the low level of interphase drag. Convective heat transfer causes ignition of the propellant and 

the pressure rises behind the projectile base due to propellant burning and flow stagnation on the 

base (0.5 ms). Gas flow readily enters the empty annular region from X = 22–55 cm, stagnates 

at the chambrage (1.5 and 3.5 ms) and returns to the rear of the chamber. As the projectile 

moves (3.5–4.7 ms), flow is drawn into its wake, the propellant bed is completely ignited, and 

the pressure field is nearly uniform. Consulting figure 16a for this time period, there are no 

significant negative pressure differentials in the chamber. 

The time sequence displayed in figure 18 was computed using the NGEN3 code for the shell 

loading configuration. High pressure gas is generated along the chamber breech face (0.5 ms) by 

the igniter as gas flow is driven forward through the propellant bed located along the radial 

chamber wall (above Y = 6 cm) and into the empty region behind the projectile base (below 

Y = 6 cm). The velocity vectors through this annular region are indicative of the low level of 

interphase drag. Gas flow stagnates on the projectile base and a region of recirculating flow is 

established in the void behind the base (0.5–1.5 ms). Convective heat transfer causes ignition of 

the propellant in the annular region around the projectile afterbody as flow stagnation at the 

chambrage builds pressure in this area and returns flow back toward the breech (1.0 and 2.5 ms). 

As the projectile moves (4.0–5.5 ms), the propellant bed is completely ignited and the pressure 

field is nearly uniform. Consulting figure 16b for this time period, there are significant negative 

pressure differentials generated before projectile movement and a series of pressure waves travel 

between the chamber breech and forward chambrage. 

The time sequence displayed in figure 19 was computed using the NGEN3 code for the full 

chamber loading configuration. High pressure gas is generated along the chamber breech face 

by the igniter as gas flow is driven forward through the propellant bed located behind the 

projectile base and in the annular region around the projectile afterbody (0.5 ms); the velocity 

vectors through the chamber are indicative of the low level of interphase drag. Convective heat 

transfer causes ignition of the propellant bed and the pressure rises behind the projectile base due 

to propellant ignition and flow stagnation on the base (0.5 ms). The propellant in the annular 

region around the afterbody ignites as flow stagnation at the chambrage builds pressure in this 

area and returns flow back toward the breech (1.0 and 1.5 ms). The full chamber loading of 
16

propellant is quickly ignited and the pressure field is nearly uniform by 2.4 ms as the projectile is 

accelerated forward (3.5–4.7 ms). Consulting figure 16c for this time period, there are 

significant negative pressure differentials between 1.0 and 1.5 ms and around 2.4 ms. The 

pressure contours displayed below also indicate that the pressure is momentarily higher near the 

chambrage for certain time periods. 

As indicated previously, the NGEN3 code was also run for each of the propellant loading 

configurations using the granular interphase drag option (figure 20) so that results early in the 

ballistic cycle (i.e., during ignition and flamespreading) could be compared to the null drag 

option. With significant interphase drag present, pressurization of the breech is much more rapid 

for the rear loading and the full chamber loading configurations (comparing figures 16a and 20a 

as well as figures 16c and 20c). As a result, pressure differentials are always positive as the 

propellant in the region behind the projectile base ignites promptly. For the shell loading 

configuration, a comparison of figures 16b and 20b shows that the results are quite similar while 

the addition of granular drag has the effect of smoothing the pressure curves—retaining the 

pressure waves while introducing some degree of damping to the ignition event. 

4. Conclusions and Future Efforts 

As first revealed in our earlier papers (1, 3, 7, and 8), the more rigorous studies of the current 

effort confirm that the telescoped ammunition concept provides a stressful configuration for 

analysis using even the most sophisticated of interior ballistic codes today. The displacement by 

a projectile afterbody of chamber volume ordinarily occupied by propellant, and equally 

significantly, a central ignition system, leads to configurational complexities likely to challenge 

the real-world charge designer as well as the theoretical interior ballistician. Moreover, the 

continued requirement for performance will likely require overall propellant loading densities 

necessitating the presence of propellant in the annular region external to the propellant afterbody. 

Ignition and combustion in this region complicates not only charge behavior, but also its 

interface with the adjacent projectile body. The current study has demonstrated the value of the 

hierarchy of interior ballistic codes in study of this problem: excellent agreement was seen in 

predicting maximum pressures using all three codes; good agreement was seen in the nature of 

pressure-wave simulations provided by XKTC and NGEN3; yet only the truly multidimensional 

representation provided by NGEN3 provided detailed insight into the controlling processes and 

interactions. Distribution of propellant in the chamber and charge permeability to gas flow 

during flamespreading and early pressurization, in addition to the projectile/chamber interface 

itself, are seen to be critical factors in achieving and acceptable interior ballistic environment. 
17

Figure 20. (a–b) Low pressure regions for simulation 

with granular drag; rear loading (left) and 

shell loading (right) (NGEN3); and (c) low 

pressure region for granular drag; full 

chamber loading (NGEN3).

(b)

(c) 

(a) 
18

Further value must now be gained from detailed evaluation of our predictive capability against 

carefully designed ballistic simulator experiments (13), allowing comparison of theoretical and 

experimental parameters not obtainable in the full gun environment. Our aim will then be to 

provide interior ballistic predictions of ever-increasing fidelity adequate to serve as the detailed 

loading profiles required to drive the transient projectile response problem (14). Clearly, what is 

needed is a routine and trustworthy approach for coupling an interior ballistic code such as 

NGEN3 to corresponding numerical codes that model the structural response of the projectile 

(e.g., DYNA3D and EPIC), so that the overall ammunition designer can arrive at a solution that 

guarantees the successful launch of a useful payload at a lethal velocity.
19

5. References 

1. Horst, A. W. Weapon System Constraints on Advances in Gun Propulsion, 5th Joint 

Classified Bombs/Warheads & Ballistics Symposium, Colorado Springs, CO, 1–2 June 2002. 

2. Gough, P. S. Interior Ballistics Modeling: Extensions to the XKTC Code and Analytical 

Studies of Pressure Gradient for Lumped Parameter Codes; ARL-CR-460; U.S. Army 

Research Laboratory: Aberdeen Proving Ground, MD, February 2001. 

3. Nusca, M. J.; Horst, A. W. Progress in Multidimensional, Two-Phase Simulations of 

Notional Telescoped-Ammunition Solid Propelling Charge. Proceedings of the 39th 

JANNAF Combustion Meeting, CPIA Publication JSC CD-25, December 2003. 

4. Gough, P. S. Modeling Arbitrarily Packaged Multi-Increment Solid Propellant Charges of 

Various Propellant Configurations. Proceedings of the 33rd JANNAF Combustion Meeting, 

CPIA Publication, 653 (1), 1996, pp 421–435. 

5. Gough, P. S. Extensions to the NGEN Code: Propellant Rheology and Container Properties. 

Proceedings of the 34th JANNAF Combustion Meeting, CPIA Publication, 662 (3), 1997, 

pp 265–281. 

6. Nusca, M. J.; Gough, P. S. Numerical Model of Multiphase Flows Applied to Solid 

Propellant Combustion in Gun Systems, AIAA Paper No. 98-3695, July 1998. 

7. Newill, J. F.; Nusca, M. J.; Horst, A. W. Advances in Coupled Projectile-Dynamics/InteriorBallistics Simulations: Coupling the DYNA3D Code and the ARL-NGEN3 Code. 

Proceedings of the 21st International Symposium on Ballistics, NDIA, Adelaide, South 

Australia, April 2004. 

8. Nusca, M. J.; Horst, A. W.; Newill, J. F. Multidimensional, Two-Phase Simulations of 

Notional Telescoped-Ammunition Propelling Charge. Proceedings of the 52nd JANNAF 

Propulsion Meeting, CPIA JPM CD-04, May 2004 (see also, U.S. Army Research 

Laboratory Technical Report ARL-TR-3306, September 2004). 

9. Whirley, R. G.; Engelmann, B. E. DYNA3D – A Nonlinear, Explicit, Three-Dimensional 

Finite Element Code for Solid and Structure Mechanics; UCRL-MA-107254 (rev 1); 

Lawrence Livermore National Laboratory: Oak Ridge, TN, November 1993. 

10. Minor, T. C.; Horst, A. W. Theoretical and Experimental Investigation of Flamespreading 

Processes in Combustible-Cased Stick Propelling Charges; BRL-TR-2710; U.S. Army 

Ballistic Research Laboratory: Aberdeen Proving Ground, MD, February 1986.
20

11. Nusca, M. J. Numerical Modeling of the Modular Artillery Charge System Using the NGEN 

Multiphase CFD Code – Effects of case Combustion. Proceedings of the 37th JANNAF 

Combustion Subcommittee Meeting, CPIA Publication, 701 (2), November 2000, pp 87–101. 

12. Chang, L.-M.; Howard, S. L.; Hui, P. Y. Experimental Evaluation of Laser-Ignited MACS 

Increments with a Modified Centercore Ignition System. Proceedings of the 37th JANNAF 

Combustion Subcommittee Meeting, CPIA Publication, 701 (2), November 2000, 

pp 102–112. 

13. Williams, A. et al. Full-scale Simulator Studies of a Notional Telescoped-Ammunition 

Charge. Proceedings of the 40th JANNAF Combustion Meeting, June 2005, in press. 

14. Ray, S. E.; Nusca, M. J.; Newill, J. F. A Study of Projectile Response to Ballistics 

Environment. Proceedings of the 40th JANNAF Combustion Meeting, June 2005, in press. 

15. Anderson, R.; Fickie, K. IBHGV2 – A User’s Guide; BRL-TR-2829; U.S. Army Ballistic 

Research Laboratory: Aberdeen Proving Ground, MD, July 1987. 

16. Boris, J. P. et al. LCPFCT – A Flux-Corrected Transport Algorithm for Solving Generalized 

Continuity Equations. NRL-MR/6410-93-7192, April 1993. 

17. Nusca, M. J. Application of the NGEN Code to Solid Propellant Charges of Various 

Propellant Configurations. Proceedings of the 33rd JANNAF Combustion Meeting, CPIA 

Publication, 653, November 1996, pp 421–435. 

18. Nusca, M. J. Application of the NGEN Interior Ballistics Code to High Loading Density and 

3-D Propellant Charges. Proceedings of the 34th JANNAF Combustion Meeting, CPIA 

Publication, 662, October 1997, pp 265–281. 

19. Conroy, P. J.; Nusca, M. J. Interior Ballistic Modeling of High-Loading Density Charges in 

an ETC Gun Using the NGEN Multiphase CFD Code. Proceedings of the 38th JANNAF 

Combustion Meeting, CPIA Publication, 712 (2), April 2002, pp 449–462. 

20. Nusca, M. J. Numerical Simulation of Plasma Ignition for High-Loading Density Charges in 

an ETC Sun Using the NGEN Code. Proceedings of the 39th JANNAF Combustion 

Meeting, CPIA Publication JSC CD-25, December 2003.
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