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Developing Improved Civil Aircraft Arresting Systems (2009)

Chapter: Chapter 11 - Aggregate Foam Arrestor Concept

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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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Suggested Citation:"Chapter 11 - Aggregate Foam Arrestor Concept." National Academies of Sciences, Engineering, and Medicine. 2009. Developing Improved Civil Aircraft Arresting Systems. Washington, DC: The National Academies Press. doi: 10.17226/14340.
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110 11.1. Concept Description 11.1.1. System Overview An aggregate foam arrestor concept has been proposed. The arrestor would use rough-broken foam aggregate made from recycled glass (Figure 11-1). The foamed, or aerated, glass material has nominally 80% void space by volume. Its closed- cell microstructure makes it resistant to water absorption and degradation. The aggregate comes in a variety of gra- dations and is currently used in different civil engineering applications for light fill construction, insulation, and frost protection of road foundations and terraces (43). For the eval- uation effort, fragments were graded to fall between 0.4 and 2.4 in., with a loose fill density of 11.2 pcf (Figure 11-2). An arrestor using the aggregate foam would be constructed by creating a basin and filling it with the material (Figure 11-3). An engineered turf would serve as a top cover layer for the bed, which can serve several purposes: 1. Prevent material dispersion due to jet blast; 2. Mitigate material spraying during overrun by an aircraft tire, thus limiting engine ingestion hazard; 3. Regulate water drainage and potential ice crust formation in winter; and 4. Act as a structural component to prevent lightweight land vehicles from penetrating the arrestor bed. This simple fill-and-cover construction would likely pro- duce lower manufacturing and installation costs than the block construction methods used at present. This potential advan- tage is offset by the possibility that the material could settle over time and result in altered arresting performance; potential mechanisms of settling could include: 1. Consolidation, and 2. Mechanical/chemical degradation and shrinkage/break- ing of aggregate pieces. Because the aggregate foam material has a closed-cell micro- structure, it has an inherent resistance to moisture. However, the presence of water during freeze–thaw cycling could present a more aggressive erosion hazard for the material. Two poten- tial methods have been identified to handle precipitation (Fig- ure 11-4). 1. Drainage Approach. This approach would allow water drainage downward through the bed. The bed would be designed to prevent standing water within it using normal civil engineering design practices. 2. Waterproof Approach. In this approach, layers of geo- plastic and geo-textile materials would be employed at the top and bottom of the bed. The upper and lower plastic layers would be sealed at the edges to produce a watertight envelope around the fill material (a practice currently in use at some landfills). Precipitation in this design would run off to the perimeter of the arrestor bed. 11.1.2. Performance Considerations Mechanically, each piece of material is composed of crush- able glass foam, which has fairly conventional properties. However, when acting as a continuum, many irregular and loose-fitting pieces would exhibit flowing aggregate behav- ior. While the material would have compressive and shear strength properties, the loose aggregate would not exhibit tensile strength because nothing binds the loose pieces together. Overall, the aggregate foam concept falls somewhere in between a conventional foam block system and a conventional hard aggregate system. Falling between these two different material categories, the aggregate foam concept presented multiple issues for evaluation: • Overall dynamic response characteristics, • Required density/strength for effective arresting, C H A P T E R 1 1 Aggregate Foam Arrestor Concept

111 • Effect of aggregate size, • Rate dependence of the material, and • Durability to freeze–thaw exposure. 11.2. Testing and Modeling Approach 11.2.1. Overview The goal for the performance evaluation was to undertake testing that would allow calibration of high-accuracy com- puter models of the aggregate foam concept. The testing approach evaluated several characteristics of the material, which was sufficient for a concept-level evaluation. The sub- sequent modeling approach for the aggregate foam material was comprehensive in nature, and the outcome was a cali- brated numerical model for predicting arrestment loads. The testing and modeling approach for the aggregate foam concept is illustrated by Figure 11-5. Five major stages are illustrated by the larger process bubbles of the chart: 1. Arrestor Material Testing and Modeling. Laboratory testing generated test data, and computer models of the material were calibrated to match it. 2. Tire Modeling. Aircraft tire models for the three test air- craft were built and calibrated to match manufacturer performance specifications. 3. Aircraft Modeling. A generalized aircraft model was devel- oped to predict the suspension response of the plane and its deceleration during a ground roll. This model was then incorporated into an APC for determining stopping dis- tances and landing gear loads when an aircraft is driven through an arrestor bed. A library of aircraft definitions was created to represent the three test aircraft. 4. Metamodeling. The arrestor material and tire models were combined to produce an overrun model for determin- ing the loads exerted on the different aircraft tires by the arrestor bed. Large data sets were generated using simula- tion batches for each tire and arrestor combination. These data sets were then accessible by the arrestor prediction code (next step). 5. Performance Predictions. The preceding four develop- ment stages culminated in the final, bottom-most process Figure 11-1. Aggregate foam material: pile of aggregate (left) and close-up of microstructure (right) (43). Figure 11-2. Typical aggregate foam fragments. Figure 11-3. Aggregate foam arrestor concept. Aggregate Foam Bed Cover Layer of Engineered Turf Arrestor Basin

on the figure. The APC was used to predict arresting dis- tances, landing gear loads, and ideal arrestor bed designs for the different aircraft. The subsequent sections of this chapter will focus on areas (1), (4), and (5). Special attention will be given to the aggregate foam material testing that was conducted and the calibration of the computer models to match the tests. The development of the tire models (2) and aircraft model (3) will be reserved for Appendix F and Appendix G, respectively. 112 Figure 11-4. Aggregate bed methods for handling precipitation and drainage. Drainage Approach Waterproof Approach Precipitation Drains Through Bed Layers of Plastic Keep Aggregate Dry (5) Performance Predictions Predict Arresting Performance for Test Aircraft (APC/MATLAB) (2) Tire Modeling Manufacturer Tire Data Build Tire Model (LS-DYNA) Calibrate Tire Model to Match Data (LS-OPT) Final Tire Models (1) Arrestor Material Testing/Modeling Conduct Material Tests Build Models Replicating Tests (LS-DYNA) Calibrate Model to Match Test Data (LS-OPT) Final Material Model Test Data (3) Aircraft Modeling Develop Arrestor Prediction Code (APC) (MATLAB) Manufacturer Aircraft Data Develop Estimated Aircraft Parameters Aircraft Library (4) Metamodeling Build Combined Tire/Arrestor Models (LS-DYNA) Batch Simulations for Tire/Arrestor Combinations (LS-OPT) Metamodel Data Figure 11-5. Testing and modeling process for aggregate foam arrestor evaluation.

113 11.2.2. Special Considerations Overall, testing the aggregate foam proved challenging for three reasons: 1. The combined crushable and aggregate properties of the material did not readily fit into standardized test methods. 2. The large size of the foam pieces required large test fixtures to capture continuum-type properties of the loose fill; this precluded the use of some desirable tests. 3. The dual-mode material behavior did not allow compre- hensive definition by the material models of the available modeling software. 11.3. Testing Effort The testing effort for the aggregate foam material involved several mechanical and environmental tests. Table 11-1 depicts the overall test matrix for the aggregate foam material. All cylinder dimensions specify diameter followed by height. 11.3.1. Density and Dimension Measurements The original, as-received aggregate foam material had pieces graded to between 0.4 and 2.4 in., with a loose fill den- sity of 11.2 pcf. For a small sample of pieces measured, the average size was 1.9 in. As previously discussed (Section 11.2.2), the as-received aggregate size precluded some testing that might otherwise have been conducted. For other tests, the aggregate was bro- ken down into smaller pieces and sifted through a 1-in. grid, giving it a gradation of 0.4 to 1.0 inches. All tests, therefore, were conducted on either the original 2.4-in. material, or on the reduced 1.0-in. material. Table 11-1. Test matrix for aggregate foam material. Test Properties Characterized Detail Number of Tests Laboratory Tests Hydrostatic Triaxial Compression Test • Compressive strength (σu) • Shear strength (τu) • Effects of confining pressure on strength • Per ASTM D2850 • 6 x 10.5” cylinder • Reduced aggregate size • Pieces graded for 0.4” to 1.0” 0.24%/min compression rate • Maximum compression of 15 to 25% • Confining pressures of 5, 10, 20, 60, and 100 psi 5 psi 10 psi 20 psi 60 psi 100 psi Total: 1 2 2 2 2 9 • Version 1: Original Aggregate Size • Non-standard • Pieces graded for 0.4” to 2.4” • 12.375 x 9.5” confining cylinder • 3 in./min compression rate • Maximum compression of 75% 2 Confined Cylinder Compression Test • Compressive strength (σu) • Compressive stress– strain curve • Extrapolated: volumetric energy capacity • Version 2: Reduced Aggregate Size • Non-standard • Pieces graded for 0.4” to 1.0” • 12.375 x 7.5” confining cylinder • 3 in./min compression rate • Maximum compression of 45% Fresh Conditioned 1 2 • Version 1: Per ATSM C 666/C 666M- 03 (modified) • Trays of material, 10 gallons total • 50 freeze–thaw cycles • Material compression tested thereafter 2Environmental Chamber Tests • Durability to freeze– thaw cycles in fully immersed conditions • Version 2: Per AASHTO T 103 • Particles separated by gradation into different sieve sizes • Particle size distribution changes measured • 50 freeze–thaw cycles 2

11.3.2. Confined Cylinder Compression Tests Confined cylinder compression tests were performed on both the original and the reduced size aggregate foam. The tests were conducted by pressing a 12-in. diameter platen into a 12.375-in. diameter cylinder at a fixed rate of 3 in./min (Fig- ure 11-6). The 12-in. diameter was chosen as being greater than six times the characteristic dimension of the material, which as a rule of thumb helps to ensure continuum material behavior. The material was poured loosely into the cylinder without packing. Due to the irregular nature of the aggregate, initial depth measurements were approximate. The load data for replicate tests was remarkably consis- tent, despite the random nature of the selected aggregate pieces for each test. However, the gradation size of the aggre- gate appeared to have a substantial effect on the loading and the energy absorption. Figure 11-7 illustrates the load history for both gradations. The general load curve shape followed an exponential form; as a result, the energy absorption also followed an exponential trend (Figure 11-8). As shown, the reduced gradation material only obtained a compression of about 45%, whereas the orig- inal larger gradation reached 75% at the same load. This led 114 Figure 11-6. Confined cylinder test for aggregate foam showing the test fixture (left), pre-test specimen (top) and post-test specimen (bottom). Figure 11-7. Confined cylinder test average load histories for aggregate foam material. Compressive Strain (in./in.) St re ss (p si) 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0 20 40 60 80 100 120 140 160 180 Original Gradation - Loading Reduced Gradation - Loading

115 to a 39% decrease in energy absorption for the reduced grada- tion material. The gradation effect may be due to two geometric differ- ences between the specimens. First, in the larger gradation material, the particles may simply have more size variation. If the reduced gradation sizing was more uniform (smaller size range), then the packing ratio may have been higher, leading to less empty space between the particles. Second, some of the larger pieces had concave surface features that may have bolstered the void percentage of the original gra- dation material. Once broken into smaller 1-in. pieces, the particles may have become rounder, which would, again, reduce the empty space between the particles. In either case, or in a combined case, a reduction in the void space between pieces would lead to earlier compaction of the material and less energy absorption. It is clear from these tests that the aggregate foam gradation would need to be chosen carefully for an arrestor application. Further, the intended gradation would need to be protected during the installation process and over the life of the bed. Repetitive vehicle overruns, jet blast, or other physical loads that might induce compaction or fragmentation of the pieces would need to be avoided. If gradation shifts occurred over time, the bed properties could change and lead to unanticipated mechan- ical performance. Another important observation from the confined cylinder tests is that the material would, in effect, act as a depth-varying compressible material. Deeper tire penetrations would lead to an increase in vertical load, not simply because a larger surface area comes into contact with the material, but because the material continually hardens as the compression increases. This trend is dissimilar from a typical block of crushable foam, which reaches a fairly constant plateau strength for the majority of the compression range. The energy absorption for a crushable block increases linearly with compression; for the aggregate foam, the increase is exponential. This depth-varying property posed an interesting alternative for arrestor applications. 11.3.3. Hydrostatic Triaxial Tests The hydrostatic triaxial tests evaluated the aggregate foam performance at different confining pressures to determine if any strength increase or bulking deformation took place. Due to the aggregate mode of behavior, lateral bulging of the spec- imens was anticipated. The largest practical specimen size for testing was a 6-in. diameter cylinder. The reduced gradation material was devel- oped to satisfy the rule of thumb for six times the nominal particle dimension across this width. All hydrostatic triaxial specimens used the 1-in. graded material. The specimens were cylinders (6 × 10.5-in.) placed between platens using Sorbothane caps at the top and bottom (Fig- ure 11-9, left). The specimens were fitted with flexible mem- brane sleeves before immersion in a pressurized vessel of water. While at this hydrostatic pressure, the specimens were compressed axially until reaching 15 to 25% compression. As shown on the right in Figure 11-9, the specimens bulked lat- erally as anticipated. A wide range of hydrostatic pressures was explored in order to determine the overall effects on the material. The upper range was chosen as 100 psi because little was observed in the way of new trends by that pressure. Because the hydrostatic specimen data had more scatter than the confined cylinder data, a total of nine tests were conducted. Figure 11-10 compares the hydrostatic triaxial data for the three main material candidates: glass foam block, aggregate glass foam, and hard engineered aggregate. The data points in the figure represent the top center for the Mohr’s circles at the different confining pressure conditions. Of the three materi- als, the only material exhibiting a change in behavior is the aggregate foam, which has a distinct transition occurring at compressive stress of 20 psi. This 20-psi stress value would occur for a test specimen at a confinement pressure of about 15 psi. Prior to the transition point, the trend-line fit to the aggre- gate foam has a steep slope and a very low initial value (near zero); these characteristics are analogous to the hard aggregate material. As the compressive stress drops to zero, so does the shear strength of the material. This is consistent with intuitive observations since lower confinement pressures allow the material pieces to roll and flow past one another more readily. Compressive Strain (in./in.) St re ss (p si) En er gy A bs or pt io n (p si ) 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0 0 20 6 40 12 60 18 80 24 100 30 120 36 140 42 160 48 180 54 Original Gradation - Loading Original Gradation - Energy Absorption Figure 11-8. Confined cylinder test stress and energy absorption for original gradation aggregate foam material.

After the transition point, the trend-line fit to the aggre- gate foam has a shallow slope and a much higher initial value (∼6 psi); these characteristics are analogous to the foam block material, albeit with lower values. The pressure dependence of the material decreases substantially, producing solid-like behavior, where the individual pieces are compressed together too much to permit flow. Because the transition point occurred at a fairly low confine- ment pressure (∼15 psi), an important simplification became possible. By observing the confined cylinder load data from Figure 11-8, only about 3 to 5% of the total material energy absorption would occur by this point. Figure 11-11 illustrates the relative energy capacity of the material, divided by behav- ioral regime. The clear dominance of the solid regime from an 116 Figure 11-9. Hydrostatic triaxial compression test of aggregate foam pre-test (left) and post-test (right), 100 psi confining pressure. Figure 11-10. Comparison of hydrostatic triaxial data for three primary material candidates. Sorbothane Cap Membrane Around Specimen Bulking Post-Test 0 5 10 15 20 25 30 0 10 20 30 40 50 60 70 80 Pe ak S he ar S tre ss , q (p si) Compressive Stress, p (psi) Glass Foam Block Aggregate Foam - Late Stage Hard Aggregate Trend - Glass Foam Block Trend - Aggregate Foam - Early Stage Trend - Aggregate Foam - Late Stage Trend - Hard Aggregate

117 energy absorption standpoint eventually led to a modeling decision to use a solid crushable foam model to simulate the material. 11.3.4. Pendulum Tests The pendulum tests that were conducted for the other can- didate arrestor concepts were omitted for the aggregate foam material. It was determined that the ungraded material would be required for such tests. Having a nominal 2-in. particle size, this presented substantial scaling issues for the pendu- lum’s 1⁄3-scale tire form. The wheel form was only 5 in. in width, leaving it at just over two particles wide. Further, the effective penetration depth of the wheel would likely have been about 7.5 in., or half the diameter. This would likely have created a spraying of the upper surface particles, but minimum compression of the material. As such, minimal energy dissipation would have occurred. Finally, the turf cover layer of the concept could not be scaled in thickness, mass, or strength, which would have made it disproportionate to the small wheel. Making a larger wheel and strut assembly for the pendulum was considered, but ultimately not feasible. Larger wheels would produce more significant loads on the pendulum mass; the wheel size for the pendulum was chosen as the upper feasible limit without causing bouncing or path deviation of the mass. 11.3.5. Environmental Tests Two sets of environmental tests were conducted to deter- mine the necessity for weatherproofing the aggregate foam material. The first test involved subjecting 10 gal of 1-in. graded material to freeze–thaw testing in various sized sieves, per AASHTO T103. The particles were fully immersed in water and subjected to 75 freeze–thaw cycles. The relative degrada- tion of the different sized pieces was recorded. It was con- cluded that the particles generally eroded in size during the course of the testing. However, this test did not ultimately provide a great deal of insight regarding the performance impact of the degradation. The second test also subjected 10 gal of 1-in. graded material to freeze–thaw testing. The test methodology was non- standard, but similar in nature to ATSM C 666/C 666M-03. The aggregate foam was placed in two open plastic tubs and fully immersed in water (Figure 11-12). The specimens were then subjected to 50 freeze–thaw cycles. After the test, the Solid Regime 95% Aggregate Regime 5% Figure 11-11. Approximate compaction- based energy dissipation in aggregate foam, divided by behavioral regime. Figure 11-12. Aggregate foam environmental test specimens before (left) and during (right) test.

specimens were oven dried to eliminate the absorbed water mass (Figure 11-13). Following the environmental tests, the specimens were sub- jected to a confined cylinder compression test (Figure 11-14) to determine the performance degradation. When compared with the fresh material, the samples exhibited a 47% decrease in energy absorption capacity. Mechanically, the closed-cell microstructure of the foam limits water absorption such that water penetrates only the outer-most open pores of the foam. Upon freezing, the expand- ing water cracks the cells, permitting progressively deeper pen- etration into the specimen as the cyclical testing proceeds. The degradation observed is, therefore, not surprising. These environmental tests represent the most severe of circumstances, where the specimens were fully immersed in water, without normal countermeasures of drainage or a pro- tective plastic envelope. Information provided by the manu- facturer indicates that cyclical temperature and humidity alone do not degrade the material over time. The presence of water immersion, therefore, is the cause for the observed degradation. Overall, these tests indicate that the aggregate foam mat- erial should be protected from immersion conditions caused by standing water. Additional testing could be conducted to characterize durability in non-immersion scenarios where a drainage bed approach is used. For the waterproof bed approach using a fully sealed plastic envelope, degradation over time is not anticipated. 11.4. Modeling Effort The modeling effort involved several stages, as shown previ- ously in the flowchart of Figure 11-5. A high-fidelity model for the aggregate foam material was calibrated to match the test data (Figure 11-5, block 1). Using this material model, an arrestor bed model was constructed and coupled with tire models for the different aircraft (Figure 11-5, block 4). Finally, large batches of simulations were conducted using these paired models, which generated volumes of data for use by the APC (also block 4). This section will discuss the arrestor model development and batch simulation process. Performance predictions for the aggregate foam arrestor concept are reserved for the fol- lowing section (Section 11.5). 11.4.1. Selection of Modeling Approach Of the three candidate systems evaluated, the aggregate foam concept was the most difficult to represent with a robust high-fidelity computer model. Because it had both loose aggregate and crushable foam properties, it did not readily fit into existing material models. Two modeling code choices were available: 1. LS-DYNA is a general purpose FEM code that supports multiple numerical methods. LS-DYNA could readily sup- 118 Figure 11-13. Aggregate foam material environmental test specimens pre-test (left) and post-test (right). Figure 11-14. Confined cylinder testing of aggregate foam environmental test specimen (post-test).

port crushable material models. However, loose aggregates could only be represented by defining many separate pieces of material and allowing them to interact on the basis of defined contacts between the pieces. This approach repre- sented a computationally expensive path that would have required infeasible simulation times. 2. EDEM is a DEM code that supports the modeling of hard aggregates. EDEM could readily model massive beds of aggregate pieces while maintaining efficient simulation times. However, EDEM did not support crushable material modeling of any kind. In light of these considerations, some simplifying assump- tions were made, and LS-DYNA was selected to model the material using a crushable foam model. The model represented the aggregate foam as a continuum of material (Lagrangian), rather than representing individual pieces of crushable foam. Because the material definition did not discretely represent the separate pieces of the material, the aggregate mode of behavior was not captured by the model. This was deemed an acceptable loss in fidelity based on several assumptions: 1. Dominant Mode of Energy Absorption. As the hydrostatic compression tests showed, the majority of the energy absorption for the material takes place in the solid regime of behavior, rather than the aggregate regime (Section 11.3.3). About 95% of the energy absorption capacity of the material exists when the material behaves in a solid-like fashion. 2. Cover Layer Effect. In arrestor applications, the aggregate- type behavior would appear in the uppermost portion of the arrestor bed, where the confinement pressures are low and the pieces are allowed to flow past and roll over one another. This upper portion of the aggregate would likely spray forward and away from the tire, as with solid aggre- gate arrestor beds. Material projected away from the tire would not participate significantly in the arresting process, decreasing the effective thickness of the arrestor bed. However, the design concept includes a cover layer that would attenuate or prevent this spray from occurring. The aggregate at the top would presumably be held in place until overrun by the tire and preserve the effective bed thickness. 3. Material Density. Unlike a hard aggregate arrestor bed, the foam aggregate has a low density (11 pcf). This leads to minimal mass-based momentum transfer effects, which can occur when an aggregate is projected, or sprayed, at high speed. Because the material is light, the momentum effects would be much less significant than the energy absorbed by material compaction. 4. SPH Formulation. Within LS-DYNA, the material was rep- resented using an SPH formulation. While this formula- tion does not represent aggregate particles, it does represent solids in a way that allows large dislocations, permitting the material to move over and past itself. The SPH methodology enhanced the continuum representation of the material. Additionally, because the material acted as a continuum, this was assumed to at least partially represent the turf layer’s mild confinement effects on the bed. 11.4.2. Smoothed Particle Hydrodynamics Formulation The aggregate foam arrestor models were developed in LS- DYNA, a general-purpose finite element modeling code. Within LS-DYNA, a number of formulations exist for repre- senting solids and fluids. Due to the high compressibility of the aggregate foam material and its loose fill properties, an SPH mesh-free formulation was employed. SPH offered the ability to represent high-dislocation solids with accuracy while maintaining time-efficient simulations. Because SPH uses particles instead of the more typical finite elements, the illustrations in this section depict the material as a collection of small spheres. Although it would be desirable given the nature of the aggregate foam material, these particles do not represent disjointed pieces of aggregate. Rather, they are mathematically interconnected to represent a continuous solid material (Lagrangian formulation). 11.4.3. Calibration to Physical Tests 11.4.3.1. Constitutive Model LS-DYNA currently offers about 200 constitutive models; of these, about 18 are applicable to various foams. Based on prior experience and on a review of the LS-DYNA keyword manual, several candidates were singled out for evaluation. After some experimentation, *MAT_CRUSHABLE_FOAM was selected as the best overall choice. This material model has parameters as given by Table 11-2. The calibration process required defin- ing these material parameters such that the model performance matched that of the physical material tests. 119 Parameter Symbol Description MID Material ID number RO ρ Density E E Young’s modulus PR ν Poisson’s ratio LCID Load curve ID for nominal stress versus strain TSC Tensile stress cutoff DAMP Rate sensitivity via damping coefficient Table 11-2. Parameters for *MAT_063 or *MAT_CRUSHABLE_FOAM.

The constitutive model assumes a normal homogeneous crushable foam material. In a typical foam, the compressibil- ity derives from the presence of voids within the cells of the foam microstructure. However, the aggregate foam includes not only these microstructure void spaces (in each aggregate piece) but also larger voids between the aggregate pieces. The approach taken for applying this material model was to match the overall net behavior of the material. The load curve definition (LCID) was therefore defined per the con- fined cylinder test data, with an exponential shape. In testing, this shape was produced by the sum of the microstructure compression and inter-particle void space compaction. Using this material model, the same summed behavior was assumed. 11.4.3.2. Material Calibration Model A confined cylinder calibration model (Figure 11-15) was developed to determine the best-fit properties for the consti- tutive model. The material parameters of the model were optimized using LS-OPT, an optimization software package developed by LSTC, the makers of LS-DYNA. LS-OPT ran the simulations in batches iteratively. After each iteration, it nar- rowed the region of interest, effectively zooming in closer to the predicted optimum calibration point. After 5 iterations of 18 simulations each, the design was optimized for a best-fit set of material parameters. Table 11-3 gives a summary of the calibration process, including the final accuracy of the calibrated model. Attempts were made to calibrate the model to also match the hydrostatic triaxial specimens. These specimens were somewhat difficult to model due to a nuance of the SPH formulation, which proved challenging to fit with a hydrostatic membrane load. Application of high confinement pressures led to instabil- ities due to the low initial slope of the material compression curve. Since the aggregate foam material exhibited minimal pressure dependence for the higher confinement pressures (Fig- ure 11-10), this was deemed a low-priority calibration point. 11.4.4. Tire and Arrestor Simulations Using the calibrated aggregate foam material model previ- ously described (Section 11.4.3), a large-scale arrestor model was created in LS-DYNA to simulate overruns by aircraft tires. Figure 11-16 illustrates the model with a 36-in. depth and a B737-800 main-gear tire (Goodyear H44.5x16.5) at 50% penetration depth. No turf cover layer for the bed was included in the model. Several possible turf layer designs were feasible, each with different thicknesses and material properties. It was assumed that, while a turf layer would confine the top layer of aggre- gate foam to prevent spraying, it would optimally not affect the mechanical response substantially. Because the spraying behavior was inherently mitigated by the continuum repre- sentation of the material, a discrete top layer was not neces- sary to determine the characteristic arrestor response. As Figure 11-17 shows, the compressed material area extended all the way to the bottom of the arrestor bed, which was in con- 120 Figure 11-15. confined cylinder model for aggregate foam material calibration. Test to Replicate Error of Calibration Confined Cylinder Compression Test Description • 12.375 x 9.5-in cylinder • Match stress–strain load curve with RMSE to 75% compression <2.0% Table 11-3. Specifications for aggregate foam material calibration.

trast to the localized compression region for the solid foam block candidate (Section 9.4.3). Waviness in the rut sidewalls and a longer compression region in front of the tire were also manifest. Figure 11-17 shows that under a static downward load, the material exhibited a pyramid-shaped compression zone. These behaviors are consistent with the anticipated mechanical response of the aggregate foam material. 11.4.4.1. Arrestor Bed Models The arrestor bed models were constructed using half- symmetry to reduce computation time. They varied in size depending on the aircraft tire being used. The bed length was determined by the distance required for the tire to make a certain number of rotations, such that the loading settled to a steady-state condition. The bed width was determined by the tire width, such that artificial boundary effects were min- imal and the response approximated that of a wide bed of the material. The smallest bed, used for the 18-in. nose tire of the CRJ-200, was 120 in. long and 9 in. wide. The largest bed, used for the 49-in. main tire of the B747-400, was 300 in. long and 36 in. wide. All beds were constructed with a 36-in. depth. However, the effective depth of the bed was adjusted by use of a mov- able rigid plane (Figure 11-18). Only the upper part of the material, above the rigid plane, was involved in the overrun compression. This approach enabled various depths to be rapidly configured within a single arrestor bed model. 121 Arrestor and Tire Model Uses Half- Symmetry Small Pieces of Aggregate Foam Fragment Off of Bed Material Compressed by Tire Tire Penetrates Vertically to a Prescribed Depth Deformable Finite Element Aircraft Tire Model SPH Arrestor Bed Figure 11-16. Model of combined tire and aggregate foam arrestor system. Pyramid-Shaped Compression Zone Figure 11-17. Compression zone under tire without forward motion.

SPH particle sizes were chosen based on the tire size. An error estimation process was undertaken to determine the required particle size to maintain an acceptably low discretization error. For the larger tires, a 2-in. particle size was found to have less than a 4% error for the predicted drag and vertical loads. For smaller tires, the particle size was reduced to 1 in. to maintain similar size proportionality. Based on the particle and bed size variations, a typical bed model had nominally 45K particles. 11.4.4.2. Tire Models The tire models were fully deformable FEM, as discussed in Appendix F. Table 11-4 summarizes the tire models devel- oped. Each tire model was calibrated to match the actual tire’s load-deflection performance up to 80% of the maximum bottoming load. This 80% load became a limit criterion dur- ing the batch simulations. The deformable nature of the tires produced an accurate representation of the interface between the tire and the arrestor material. As the load on the tire increased due to deeper bed penetration, the contact area became flatter with an increased surface area. This shape change created a corresponding increase in the load on the tire. 11.4.4.3. Sequencing of Simulations Because the tires in the LS-DYNA simulations were deform- able and were allowed to spin freely, a sequencing method was required to create stable, fast-running simulations. These two factors were additional complications that were not present in the EDEM aggregate simulations. However, the inclusion of these factors led to higher-fidelity results. From a mechanical standpoint, as the axle of the wheel penetrated the bed vertically, both the tire and the arrestor material underwent compression. The interplay of the tire and arrestor compression created oscillations in the mea- sured loads. This oscillating behavior was further amplified by the free-spinning nature of the tire. Eventually the tire would settle to a constant rate of rotation, which proved to be a function of the forward speed, depth, and interface friction. It was found that a minimum forward travel dis- tance was required for the loads and rotation to reach steady-state conditions before load measurements could be made. Sequencing options included several factors: • Prescribed vertical penetrations versus prescribed vertical loads; • Applying the vertical penetration/load before or after beginning the forward motion; and • Applying the forward motion before or after making con- tact with the bed Depending on the sequencing method used to accelerate the tire and set the penetration depth, the initial oscillations could be more or less severe. This in turn could require longer or shorter simulation times, and longer or shorter arrestor beds. Because the arrestor–tire models were to be run repeat- edly in large batches, it was important to develop a sequenc- ing methodology that would produce efficient simulation run times. Multiple methods were attempted through experimentation before settling on the approach illustrated by Figure 11-19. The tire was first pressed downward into the material to a pre- scribed depth. Then the tire was accelerated to the desired for- ward speed and spun-up to an initial rotation speed (typically about one-third of the ideal rotation rate expected on hard pavement). The prescribed spin rate was then released, allow- 122 36-inch Deep Bed Rigid Plane at 18- inch Depth Only Upper Part of Material Involved in Compression Figure 11-18. Adjustable height of aggregate foam arrestor bed. Aircraft Landing Gear Tire Designation Main Gear H29x9.0-15 CRJ-200 Nose Gear R18x4.4 Main Gear H44.5x16.5-21 B737-800 Nose Gear H27x7.7-15 Main Gear H49x19-22 B747-400 Nose Gear H49x19-22 Table 11-4. FEM tire library for aggregate foam arrestor models.

ing the tire to settle to a natural rotation rate, while the forward motion continued at a constant speed. After the oscillations settled out of the system, the steady-state vertical and drag loads were measured. The final result was an accurate prediction of the loads on the tire under free-spinning un-braked conditions. 11.4.5. Batch Simulations Using the arrestor bed model, large batches of simulations were conducted to generate substantial bodies of data for a wide range of overrun conditions. This data was then assem- bled into “metamodels” for uploading and use by the APC. 11.4.5.1. Methodology Batch simulations were conducted for each tire with three open variables: • Speed, from 10 to 70 knots (Speeds below 10 knots were impractical due to the long simulation times required for a tire to travel the required minimum distance. Loading at speeds below 10 knots was based on the extrapolated meta- model data fit.); • Bed depth, in incremental depths from 3 to 36 in. (Fig- ure 11-20); and • Penetration into the bed, from 10% to 100% of maximum penetration depth. 123 Tire presses down into arrestor Tire accelerates forward and is given an initial spin rate Tire continues forward and is allowed to freely spin Steady-state vertical and drag loads are measured Tire begins above arrestor Figure 11-19. Sequencing method for aggregate foam arrestor model. Bed Depth Bottoming Depth Rut DepthPenetration Depth Tire Deflection Figure 11-20. Depth definitions for aggregate foam bed models.

Due to the two sources of compression (arrestor material and tire), the definition for penetration depth was more com- plex than for the rigid tire approach used for the aggregate arrestor models. Two conditions defined the maximum pen- etration depth: 1. The maximum penetration depth was considered to be 85% of the bed depth (fully compressed material) plus the deflection of the tire at 80% of the bottoming load. Beyond this degree of penetration, the tire models were no longer accurate. 2. For small tires in deep beds, the maximum penetration was further limited to be no greater than the tire diameter. At depths beyond this, the simulations often did not settle to steady-state conditions. The large batch simulations were conducted using LS- OPT. Based on the initial model files, LS-OPT generated permutations with various speeds, bed depths, and penetra- tion levels. It sequentially executed the simulations and extracted the load data from them. Generally, the batches were conducted in multiple iterations of 10 simulations each. Additional iterations were added to improve accuracy as needed. 11.4.5.2. Summary Tables of Metamodels The output from the batch simulations was extracted and assembled automatically by LS-OPT, where metamodels were constructed for the drag and vertical load forces. Meta- modeling is analogous to fitting a curve through experi- mental data, except it is applied to multi-dimensional data sets. These data sets were four-dimensional, including speed, depth, penetration, and load (either vertical or drag). The metamodels were RBF networks, which can effectively capture non-linear behaviors including multiple concavity changes across the data set. Table 11-5 summarizes the fit quality for the metamodels. The RMS error showed the highest variation among the three candidates simulated, which was typically below 7%, but had three instances of higher values for the 29-in. and 18-in. tires. The R-squared value was typically above 0.97, but again dropped below 0.95 for the 18-in. tire. Typically, these prob- lems would be resolved by adding additional simulation points. In this case, however, the scatter in the data persisted even when more simulations were conducted. A closer examination of the batch simulation output reveals that the cases most responsible for the scatter involve deeper arrestor beds. Because the material essentially has a depth- varying quality to it, with an exponentially shaped load curve, it tended to exhibit slip–stick behavior under the tire. The material would compact into a clump until enough force was built up for a section to break free, forming a fissure. Two potential solutions to this data scatter would have involved (1) longer rolling distances in the simulations with corre- spondingly increased run times, and (2) revisions to the con- stitutive model for the aggregate foam material. Project time and budget constraints prohibited further pursuit of improve- ments, so the existing metamodels were used. As the table shows, the points used in the metamodels were often less than the total number of simulations conducted. This discrepancy was caused by simulations that failed prior to termination due to tire overloading, or simulations that had not adequately settled to steady-state conditions for accurate load measurement. The smaller tires experienced a greater percentage of omitted runs than the larger tires due to the relative loading severity and greater proportional pene- tration depths. 124 Tire SimulationsConducted Points Used Response RMS Error R 2 Drag 7.10% 0.995 H49 50 50 Vertical 5.96% 0.994 Drag 3.57% 0.999 H44 60 58 Vertical 3.05% 0.998 Drag 16.70% 0.971 H29 60 47 Vertical 7.70% 0.986 Drag 3.39% 0.998 R27 40 40 Vertical 3.78% 0.997 Drag 23.80% 0.948 H18 80 73 Vertical 21.40% 0.943 Table 11-5. Metamodel accuracy summary for aggregate foam arrestor bed.

11.4.5.3. Parameter Sensitivities Using the metamodels, it is possible to review how sensi- tive the landing gear loads are to the different variables of speed, bed depth, and penetration percentage. Figure 11-21 shows a surface plot for the 44.5-in. main-gear tire of the B737 in an 18-in. deep aggregate foam arrestor bed. Stronger drag loads (shown as the lower, more negative val- ues) occur when the penetration ratio increases, up to the max- imum of 1.0 (or 100%). The loading is, therefore, strongly dependent on the depth of penetration into the bed, as would be expected. However, unlike with crushable block foam, the increase is non-linear, as shown by the downward concavity of the surface. By contrast, the variation with speed shows very little change between 10 and 70 knots. The loading is fairly insensitive to speed, reflecting the low rate-sensitivity that would be expected from crushable glass foam. Practically speaking, this means that the aggregate foam system will exert nearly the same decel- eration load on an aircraft travelling at high or low speed. This behavior is desirable for an arrestor and is consistent with the general behavior of the current cellular cement material. While overall rate insensitivity is expected for a glass foam material, the aggregate foam glass presents an additional layer of complexity. The inter-particle void compaction may not be a rate-independent process, and that could affect the over- all rate sensitivity of the medium. This trait requires confir- mation through larger-scale testing. 11.4.5.4. Data Transformation The final metamodel data for each tire was converted for use by the APC. LS-OPT was used to extract nominally 9,000 data points from each metamodel and export it into tabular form. A MATLAB conversion program was written to map this data into multi-dimensional matrix form that could be quickly accessed by the APC. 11.5. Arrestor Performance Predictions 11.5.1. Scope of Simulations Using the APC, a separate optimal arrestor was designed for each of the three trial aircraft: CRJ-200, B737-800, and B747-400. Subsequently, an optimal mixed-fleet arrestor was designed as a compromise best-fit for all three aircraft. All arrestment predictions assumed the following: • 50-ft setback distance; • 50-ft gradual decline to the maximum bed depth; 125 D ra g Fo rc e (lb f) Speed (knots) Penetration Ratio Figure 11-21. Metamodel drag load surface plot for 44.5-in. tire in an 18-in. deep aggregate foam bed.

• 70-knot starting speed for the aircraft; • No reverse thrust; • Braking factor of 0.25 before and within the bed; and • Arrestor bed loads based on interaction with tires, neglect- ing strut and axle components. Arrestor beds were designed for two different nose-gear loading criteria: 1. Limit Load Criterion, where the drag load applied to the nose strut cannot exceed the limit load for the nose gear (FAR Part 25.509); 2. Ultimate Load Criterion, where the drag load applied to the nose strut cannot exceed the ultimate load for the nose gear; Since the ultimate loading criterion permits higher loads on the strut, deeper beds and shorter stopping distances resulted from those cases. It was determined through experimentation that the aggre- gate foam arrestor design functioned equally well for a par- tially or fully recessed bed. In practice, a partially recessed bed would require a shallow basin, and a raised, flat mound of arrestor material covered with turf. The simulations conducted in this section, however, assume a simpler, fully recessed bed, where the top of the arrestor is level with the runway (Fig- ure 11-3). Two design variables were considered for the arrestor: the bed thickness and the material strength. The material strength was adjusted by applying a scale factor to the metamodel load- ing data in the APC during a simulation. It was considered an open variable because the aggregate foam can be manufactured at a variety of strength levels (varying density). 11.5.2. Performance for Test Aircraft Table 11-6 lists best-case arrestor designs for each aircraft taken individually. Each arrestor bed listed uses a different material strength and depth optimized for the design aircraft. Compared with the similar EMAS design cases on the right, the distances are all somewhat longer. Material strength in the table is given as a percentage of the tested material (original gradation) since there is not a simple plateau strength as with normal crushable foams. The scaling assumes that a 75% maximum compression is retained and that the gradation remains the same. However, the composi- tion of the aggregate foam material itself is altered in density to effect a stronger or weaker foam aggregate. Generally, a range of acceptable strength and depth combinations was available, so a different material strength could be chosen than the one given. Table 11-7 shows the compromise design case with the best arrestor design for all three aircraft. With the material strength and depth as specified, the B747 would require 592 feet to decelerate from 70 knots. Per typical practice for EMAS design, the bed length may be specified such that all aircraft satisfy the minimum 40-knot exit speed requirement. The bed designs in the table are fixed with a 400-ft length for comparison with the other alternatives. At this length, the B747-400 would have a maximum exit speed of approximately 56 knots, which satis- fies the requirements of AC 150-5220-22a. Interestingly, this multi-aircraft design is the best one-size- fits-all arrestor bed from among the three candidate systems. Some of the other candidate design cases provided better single-plane performance, but they lagged comparatively 126 Nose-Gear Limit Load Criterion Nose-Gear Ultimate Load Criterion Current EMAS, Optimal Designs Aircraft Strength (%) Depth (in.) Bed Length (ft) Strength (%) Depth (in.) Bed Length (ft) Depth (in.) Bed Length (ft) CRJ-200 68% 19.0 326 59% 21.1 309 22 258 B737-800 103% 29.3 356 72% 36.0 315 22 287 B747-400 61% 36.0 580 61% 36.0 580 28 495 Table 11-6. Individual aircraft 70-knot arrestor beds for aggregate foam arrestor system. Nose-Gear Ultimate Load Criterion Bed Dimensions Bed Depth: 36.0 in. Bed Length: 400 ft Material Strength Scaling: 86% Aircraft Exit Speed (knot) Stopping Distance (ft) CRJ-200 70+ 367 B737-800 70+ 327 B747-400 56 400 Table 11-7. Fleet design arrestor bed for aggregate foam arrestor system.

with respect to the multi-plane case. The reason for the supe- rior performance appears to be the depth-varying nature of the material. Deeper beds were feasible for the aggregate foam, which performed better in arresting the B747, with its large-diameter tires. Yet, because of the depth-varying prop- erties, the smaller aircraft each found their own equilibrium rut depths without being overloaded. The equilibrium depth effect is further discussed in Section 11.5.3.2. 11.5.3. General Observations 11.5.3.1. Deceleration and Nose-Gear Loading The overall deceleration and loading trends on the three air- craft showed several common characteristics. Figure 11-22 and Figure 11-23 illustrate sample deceleration and nose-gear load plots generated by the APC for the B737 aircraft. Figure 11-22 represents the arrestor bed using the limit load design criterion, while Figure 11-23 is based on the ultimate load criterion. The overall bed lengths and depths reflect the values of Table 11-6. The upper plot in each figure shows the aircraft speed decrease relative to the nose wheel location in the arrestor bed. The lower surface line of the arrestor is depressed by the bed thickness, as discussed in Section 11.5.1. The middle plot in each figure shows the deceleration of the aircraft in g’s. The deceleration first increases when the nose wheel penetrates the bed at zero feet, while the main-gear tires are still on pavement. The deceleration then increases strongly when the main-gear tires enter the arrestor, at a nose-wheel location of 50 to 100 ft. After the initial transition, the deceler- ation oscillates in both cases, at about 0.65 g for the limit design and about 0.8 g for the ultimate design. The oscillations are substantially more severe in the ultimate design case due to the combined effects of the deeper bed and the depth-varying material properties. Among the three arrestor candidates, this oscillatory tendency is unique to the foam aggregate concept. The reason for the oscillations becomes clearer when analyzing the wheel rut characteristics in the next section. The lower plot in each figure shows the nose-wheel drag loading, which proved to be the limiting load for the arrestor bed design. The loading is highest between 100 and 120 ft, which is after the main-gear tires have entered the bed. When the main-gear tires enter, the deceleration increases sub- stantially, and this causes the aircraft to pitch forward and presses the nose wheel deeper into the material. The deeper penetrations lead to higher drag loads on the nose wheel. 11.5.3.2. Rut Depth Effects Figure 11-24 gives the corresponding main and nose-wheel rut depths for the limit case of Figure 11-22. As shown by the lower dashed line in each figure, the rut is only about 10 to 12 in. deep in the material. For a normal crushable block foam system, the rut would tend to cut near the bottom of the bed at the bottoming compression depth of the material. In this case, however, the depth-varying properties of the foam aggregate (exponential curve of Figure 11-7) allow the tires to find their own natural equilibrium depth in the material, which is at a much shallower penetration. The tires are, in a sense, floating atop a layer of partially compacted material. Because the material has not been fully compressed, oscilla- tions can result as the tire bounces above and below the equi- librium rut depth. Because of the rut depth effect, large oscillations were observed in a number of design cases attempted. New design practices were required to stabilize the ruts. In general, two main principles emerged: 1. Oscillations are more likely for beds that are deep relative to the tire diameter, and 2. Oscillations are more likely where the decline distance is not substantially longer than the wheel base of the aircraft (discussed further in the next section). The B747 was not observed to form any lasting oscillations in any design case. Any bouncing quickly settled to a steady- state condition. It was determined that this was because all tires on the B747 are 49-in. diameter, which is substantially larger than the maximum 36-in. bed depth. 11.5.3.3. Decline Distance Effects The decline distance is the length of the depth-tapering region of the arrestor bed. For the two other candidate sys- tems, a 50-ft decline distance was sufficient to produce stable arrestors. However, it was found that a 100-ft decline was considerably more stable for the aggregate foam concept due to its relative size with respect to the aircraft wheelbase. A shorter 50-ft decline produced sometimes drastic oscil- lations, as shown in Figure 11-25 and Figure 11-26. This design case for the B737 is identical to that of Figure 11-22 except for the decline distance, which was 100 ft. The cause of the oscillations appears to be the initial pitch rate caused by the sudden drop of the nose wheel, which set up a “por- poising” behavior in the aircraft. Since the rut bottom was floating, rather than bottomed, the motion damped out slowly. 11.5.4. Braking Effects In the APC simulations, braking loads for the main-gear tires while in the arrestor bed were added to the drag loads of the tire–arrestor metamodel. Thus, the net drag load on the main-gear tires was due to the braking plus the arrestor 127

128 -40 -20 0 20 40 60 80 Sp ee d [kn ot ] a nd D ep th [in .] Aircraft Velocity and Bed Profile Aircraft Speed Upper Surface of Arrestor Lower Surface of Arrestor -50 0 50 100 150 200 300250 350 400 0.2 0.3 0.4 0.5 0.6 0.7 0.9 0.8 Nose-Wheel Location [ft] -50 0 50 100 150 200 300250 350 400 Nose-Wheel Location [ft] -50 0 50 100 150 200 300250 350 400 Location [ft] D ec el er at io n [g ] Aircraft Deceleration Aircraft Deceleration 0 5 10 15 20 25 30 35 40 Fo rc e (ki p) Landing Gear Forces - NOSE STRUT Nose Gear Drag Nose Gear Limit Load Nose Gear Ultimate Load Figure 11-22. Limit criterion aggregate foam arrestor design plots for B737-800 showing speed (top), deceleration (middle) and nose-gear drag load (bottom).

129 -40 -20 0 20 40 60 80 Sp ee d [kn ot ] a nd D ep th [in .] Aircraft Velocity and Bed Profile Aircraft Speed Upper Surface of Arrestor Lower Surface of Arrestor -50 0 50 100 150 200 300250 350 0.2 0.3 0.4 0.5 0.6 0.7 1 0.9 0.8 Nose-Wheel Location [ft] -50 0 50 100 150 200 300250 350 Nose-Wheel Location [ft] -50 0 50 100 150 200 300250 350 Location [ft] D ec el er at io n [g ] Aircraft Deceleration Aircraft Deceleration 0 5 10 15 20 25 30 35 40 Fo rc e (ki p) Landing Gear Forces - NOSE STRUT Nose Gear Drag Nose Gear Limit Load Nose Gear Ultimate Load Figure 11-23. Ultimate criterion aggregate foam arrestor design plots for B737-800 showing speed (top), deceleration (middle) and nose-gear drag load (bottom).

resistance. As an approximate solution, this approach worked well and was appropriate. However, a small side study conducted during the metamod- eling process indicated that braking applied to the main-gear tires could cause the penetration depth to increase. Since the arrestor metamodel loads assumed a non-braked free-spinning wheel, this depth change was not captured in the predictions. For a bottomed tire, little depth change is feasible and this effect would likely not occur. For a non-bottomed tire, how- ever, the tendency to penetrate deeper would lead to higher arrestor drag loads on the tire. Since this effect would only apply to the braked wheels of the main gear, it would benefit the deceleration process while not affecting the nose gear. Stopping distances could be somewhat reduced by this effect, though it is unclear by what amount or in what limited sub- set of arrestor design cases. Additionally, the sudden application of, or increase in, air- craft braking could induce a porpoising behavior similar to the short decline distance effects of Section 11.5.3.3. This brake- induced oscillatory behavior could be studied using the APC, but such an investigation fell outside the scope of the current effort. In actual practice, it seems doubtful that pilots in over- run situations would release and re-apply the brakes of the air- craft. It may prove most useful to characterize the torsional forces associated with a fully applied brake and assume those loads throughout the simulations. 11.5.5. Short Landings Short landings involving an aircraft touch down inside the arrestor bed were not simulated. However, the potential for short landings presents two possible issues. 130 -150 -100 -50 0 50 100 150 200 250 300 350 Location [ft] Location [ft] D is pl ac em en t [ in. ] Vertical Displacements - MAIN STRUT Axle Displacement Rut Depth Lower Bound of Bed Upper Bound of Bed -50 0 50 100 150 200 250 300 350 400 -30 -20 -10 0 10 20 D is pl ac em en t [ in. ] -30 -20 -10 0 10 20 Vertical Displacements - NOSE STRUT Axle Displacement Rut Depth Lower Bound of Bed Upper Bound of Bed Figure 11-24. Limit criterion aggregate foam arrestor design plots for B737-800 showing axle and rut depth for the main strut (top) and nose strut (bottom).

131 Aircraft Speed Upper Surface of Arrestor Lower Surface of Arrestor Aircraft Deceleration Nose Gear Drag Nose Gear Limit Load Nose Gear Ultimate Load -40 -20 0 20 40 60 80 Sp ee d [kn ot ] a nd D ep th [in .] Aircraft Velocity and Bed Profile -50 0 50 100 150 200 300250 350 1.2 0.2 0.4 0.6 0.8 1 Nose-Wheel Location [ft] -50 0 50 100 150 200 300250 350 Nose-Wheel Location [ft] -50 0 50 100 150 200 300250 350 Location [ft] D ec el er at io n [g ] Aircraft Deceleration 0 5 10 15 20 25 30 35 40 Fo rc e (ki p) Landing Gear Forces - NOSE STRUT Figure 11-25. Limit criterion aggregate foam arrestor—50-ft ramp—design plots for B737-800 showing speed (top), deceleration (middle) and nose-gear drag load (bottom).

132 0 20 40 60 80 100 Fo rc e (ki p) Landing Gear Forces - MAIN STRUT Main Gear Drag Main Gear Limit Load Main Gear Ultimate Load -150 -100 -50 0 50 100 150 200 250 300 Location [ft] -150 -100 -50 0 50 100 150 200 250 300 Location [ft] Vertical Displacements - MAIN STRUT Axle Displacement Rut Depth Lower Bound of Bed Upper Bound of Bed D is pl ac em en t [ in. ] -30 -20 -10 0 10 20 Figure 11-26. Limit criterion aggregate foam arrestor—50-ft ramp— design plots for B737-800 showing main-gear strut loading (top) and main strut axle and rut depth (bottom). First, the aggregate foam arrestor exhibits a depth-varying strength property that is advantageous for arresting varied aircraft fleets. However, during a landing in an arrestor, the initial downward loads would be high, and this could cause a deep penetration and a corresponding drag-force overload to the landing gear. Second, the basin geometry of the arrestor concept would force the aircraft to roll up the decline slope in the reverse direction from normal, acting as a ramp that would cause a strong load to the landing gear. This issue could be eliminated by only partially recessing the bed, as in the ideal EMAS design cases. Mechanical performance was found to be nearly equivalent with either version. However, the turf covering of the arrestor would be above grade in such cases, presenting some construction and aesthetic complications. 11.6. Estimated System Cost and Upkeep 11.6.1. Installation Process The loose aggregate solution offers the advantage of construct-in-place simplicity, which could produce instal- lation cost savings over a traditional EMAS. It reduces site preparation and eliminates block manufacturing, placement, and joint sealing. However, turf preparation and placement would be additional tasks not required for the current EMAS. The foam aggregate concept would require excavation of an arrestor bed basin with a depth nominally equivalent to that of an EMAS bed. This basin may or may not require paving before being filled with foam aggregate. However, the below-grade nature of the basin would require drainage from

the bed to be included in the design using standard roadway engineering practices. The basin would be filled with foam aggregate using earth- moving equipment. Due to the crushability of the material and the importance of maintaining the original gradation and den- sity (Section 11.3.2), the foam aggregate would require care during installation. Bulldozers, or an equivalent, would reverse fill the bed starting from the distal end and working backwards, such that they did not overrun previously deposited material. Only vehicles with low ground pressure would be permissible on top of the bed material, such that the material was not com- pacted beyond design specifications. If a roller was to be used for creating a level surface on top of the bed before applying the turf cover layer, a prescribed degree of compaction would need to be pre-defined (Figure 11-27). However, grading the bed could require a fairly manual process to obtain the desired results. Depending on the design of the reinforced turf layer, it could be grown in place, or it could be grown in advance, then cut into segments and placed atop the bed. The latter alterna- tive would require the use of a front-end loader equipped to provide low ground pressure. Once completed, the bed would permit normal traffic from grounds keeping vehicles and pedestrians. Vehicles with high tire pressures, such as emergency vehicles and aircraft, would be restricted from driving on top of the bed. 11.6.2. Cost to Establish System A preliminary estimate was made for the cost to establish an aggregate foam arrestor system. It must be noted that the cost estimate from this section is only a basic approximation for the purposes of comparing the different arrestor alterna- tives. The cost estimate is based on a mixture of information from the manufacturer, the airport survey, and FAA Order 5200.9. To develop a more accurate estimate of the costs to install such a system, it is recommended that a detailed cost quote be sought from a firm qualified to undertake an instal- lation effort. Where possible, the methodologies used were consistent with the prior survey information collected regard- ing the existing EMAS (Section 3.5). The costs may be broken into two major categories: site preparation and installation. The site preparation costs were estimated for two cases. The aggregate foam arrestor would use a basin for the arresting materials rather than a flat runway-type surface as is used for the current EMAS design. The bottom of the basin could either be paved or earthen. Drainage, excavation, and leveling would be required for either option. Assuming that a full paved surface is not pro- vided under the bed, the cost for site preparation was assumed to be reduced by half; this value was used for the lower-bound cost estimate for the system. If a full paved sur- face is provided, then the preparatory costs were assumed to be the same as for the current EMAS; this provided the upper- bound cost estimate. The installation cost estimate was separated into specific materials and general installation labor needs. Because these costs were specific to the aggregate foam arrestor concept, they do not have a direct connection to any prior EMAS data. Discussions with the manufacturer produced cost estimates for the aggregate foam, reinforced turf cover layer, and geo-textile/ geo-plastic layers. Where applicable, materials included freight costs for trans-Atlantic shipping. The labor costs were based on estimates from the manufacturer established from similar installation efforts. Finally, the site preparation and estimated EMAS costs were computed in two ways: (1) assuming average survey costs from this research, and (2) assuming FAA Order 5200.9 costs. The final cost estimates for both options are given in Table 11-8 and Table 11-9, respectively. Using the survey cost assumptions of Table 11-8, a 300-ft arrestor bed would cost between 47% and 60% less than the current EMAS. If the Order 5200.9 costs are assumed, the cost advantage drops to between 37% and 44% (Table 11-9). 133 Figure 11-27. Roller compaction of aggregate foam material during typical installation (43). Aggregate Foam System Cost Category Lower Bound Upper Bound Current EMAS Site Preparation $ 1.08 $ 2.17 $ 2.17 Installation $ 2.16 $ 2.16 $ 6.03 Cost to Establish $ 3.24 $ 4.32 $ 8.19 Percent of EMAS 40% 53% Table 11-8. Estimated costs to establish aggregate foam arrestor, 150 x 300 ft, assuming survey average costs for current EMAS, units of millions USD.

In addition to the tables in this section, longer-term life- cycle issues could also be considered. FAA Order 5200.9 includes a standard 10-year replacement interval for an EMAS, which translates into present-value life-cycle costs. Such a replacement could arguably be unnecessary for this arrestor concept (Section 11.6.4). Eliminating the assumed 10-year replacement could effectively trim about $2.6M of present-value life-cycle costs (based on the EMAS replace- ment cost estimates of the survey). 11.6.3. Maintenance Maintenance for the aggregate foam concept would be rel- atively simple, and should be limited to standard grounds- keeping measures for the protective turf layer. Drainage of the area to prevent standing water is required, and periodic inspections would be advisable to ensure that no issues arise due to seasonal weather changes. Due to the lack of joints and blocks, many protective measures used in current EMAS con- struction would not be necessary. The aggregate foam material is expected to have a superior durability when compared with cellular cement. Aggregate foam does not exhibit the tendency to crumble during han- dling as occurs with cellular cement, nor does it exhibit overt sensitivity to moisture due to its closed cell microstructure. Many industrial applications of glass foam materials indicate long service life is possible with little degradation, as long as mild protective measures are taken. However, the aggregate foam carries some maintenance issues not present for solid foam block material. The loose aggregate could settle over time, or the aggregate pieces could eventually break down into smaller pieces, changing the gradation of the material. These are both issues to address through inspections. Plate-shaped dimensional markers could be inserted below the turf layer and checked periodi- cally using surveying equipment for changes in height, and aggregate samples could be tested periodically (discussed fur- ther in the next section). 11.6.4. Replacement and Overhaul FAA Order 5200.9 plans a full replacement for an EMAS after 10 years. Following this schedule for replacement may be required for the drainage version of the concept, but this would probably not be required for the sealed watertight bed. Given the minimal cost increase, adding the watertight membrane around the foam aggregate would likely afford considerable long-term savings. Eliminating the 10-year replacement could effectively trim about $3.5M of present- value life-cycle costs, based on the replacement cost esti- mates of the survey. Rather than a scheduled replacement, periodic bed material sampling could be undertaken, perhaps in 3- to 5-year inter- vals. Small areas of the cover layer could be removed, and a sample of the aggregate foam could be extracted and tested. The removal area would then require refilling and repair of the top layer. These periodic small-scale tests would provide assurance that the bed has not degraded and still meets per- formance specifications. 11.6.5. Repair After an overrun event, the rut areas would require repair. The foam aggregate remaining in the ruts would require removal and replacement. The damaged cover layer would also require removal from the rut areas and subsequent replace- ment. Geo-textile and geo-plastic layers would also require repair and fusing with the existing material to ensure contin- ued waterproofing, as applicable. 11.7. Transition to a Fielded System In order to transition the aggregate foam concept to a fielded system, the following additional development steps may be advisable. 11.7.1. Material Density and Compaction Calibration As shown in the example arrestor beds of Table 11-6 and Table 11-7, the compressive strengths for optimal designs ranged from 59% to 103% of the tested material. To estimate the needed strength in the arrestment predictions, the meta- model data was simply scaled, which was an appropriate simplification. However, in migrating to a fielded system, several standard densities and gradations would probably be selected to pro- vide flexibility during the design of actual arrestors. For each density and gradation, some testing would be required to generate calibration data. 134 Table 11-9. Estimated costs to establish aggregate foam arrestor, 150 x 300 ft, assuming Order 5200.9 costs for current EMAS, units of millions USD. Aggregate Foam System Cost Category Lower Bound Upper Bound Current EMAS Site Preparation $ 0.34 $ 0.68 $ 0.68 Installation $ 2.16 $ 2.16 $ 3.83 Cost to Establish $ 2.50 $ 2.84 $ 4.50 Percent of EMAS 56% 63%

The installation process will likely affect the performance properties of the material because compaction by construc- tion equipment and foot traffic will take place. Since this process is inevitable, it would be advisable to create several larger-scale testbeds using the anticipated installation process and equipment. Tests could then be conducted on these small beds using larger test apparatuses than those of the current research. From the results, a revised material model could be calibrated, and the arresting performance reassessed. In this manner, a calibrated prediction capability would exist for the as-installed material. Thereafter, installation of actual arrestor beds could proceed per the developed methodology with high confidence of the final system performance. Such tests could be conducted in conjunction with the cover layer design (next section). 11.7.2. Cover Layer Design The cover layer options that have been discussed include reinforced turf with or without additional geo-textile/ geo-plastic layers. Additional testing and modeling would be required for whichever method is selected since the membrane behavior of a cover layer will affect the dynamic mechanical performance of the arrestor bed. The cover layer performance should be further characterized under frozen conditions since it will be exposed to such condi- tions even where the foam aggregate layer of the system is waterproofed. 11.7.3. Braking Dynamics The research performed has identified that braking in the aggregate foam material may require further study. Of the three major concepts evaluated, braking dynamics appear likely to affect the aggregate foam concept the most. Because the depth-varying properties of the material have demon- strated tire “flotation” at mid-depth bed penetrations, there is room for potential depth shifts of the tires due to braking. If they occurred, such shifts could lead to excessive landing gear loads. The current modeling method of the APC makes simplify- ing assumptions regarding this phenomenon that are suffi- cient for a concept-level evaluation. However, to ensure accuracy of design predictions, some additional tests would be beneficial. One method would involve using a one-wheel bogy apparatus fitted with brakes and a load measurement system that is towed through the material. 11.7.4. Full-Scale Testing A full-scale aircraft overrun test of a foam aggregate arrestor bed is advisable because this concept represents a substantial departure from the current EMAS design in terms of mechanical loading and the materials used. 11.8. Summary The aggregate foam arrestor concept was found to have a mechanical response similar to a crushable foam material, except with depth-varying properties. It would absorb energy from the aircraft primarily though material compaction rather than through displacement. An aggregate foam arresting bed would be constructed using a shallow basin of the material topped with a reinforced turf cover layer. The material is closed-cell glass foam that provides inher- ent moisture and chemical resistance, and improved handling durability as compared with cellular cement. Manufacturer information indicates that long service life is possible, poten- tially eliminating the standard 10-year replacement assumed in FAA Order 5200.9. Water immersion must still be avoided, so the preferred design approach would use a sealed plastic geo-membrane envelope coupled with standard drainage provisions. Installation of the system would likely be simpler and less expensive than the current EMAS since placement of blocks and sealing joints are both unnecessary. Heavy equipment would place the material in the bed basin and top it with reinforced turf. Geo-membrane and geo-textile layers, as applicable, would be placed and joined manually. The arrestor basin could be constructed with or without paving, which could provide for preparatory cost reduction. In order to pre- serve material gradation and prevent over-compaction dur- ing installation, an appropriate installation process would be required. The cost to establish such a system would be nominally 40% to 53% of the survey cost of the existing EMAS, which provides the most substantial estimated cost reduction of the candidates evaluated; much of the cost reduction is due to the markedly less expensive aggregate foam material. Life-cycle costs could be further reduced due to longer bed life. Main- tenance needs appear to be simplified, requiring standard grounds-keeping measures, but no block or joint repairs. The APC predictions for the aggregate foam arrestor show fairly constant deceleration throughout the arrestment with little speed dependence, which are desirable characteristics. The depth-varying material characteristic produced a unique “floating” rut depth that was not observed for the other mat- erials evaluated. This characteristic allowed each tire to settle to its own natural depth in the material, creating more even load distribution among tires of different sizes. Bed lengths on a per-plane basis were nominally 15% longer than for the current EMAS. However, the multi-aircraft design case demonstrated the best one-size-fits-all perfor- mance from among the three candidate systems. A 400-ft 135

arrestor bed demonstrated 70+ knot exit speeds for the B737- 400 and CRJ-200 and a 56-knot exit speed for the B747. The reason for the superior performance appears to be the depth- varying nature of the material. The floating rut characteristic also gave rise to oscillating tendencies in which the plane could exhibit porpoising behav- ior. This was mitigated through appropriate bed geometry design. However, additional investigation would be required to establish the effects during short landings and in cases where the pilot applies intermittent braking. Transition to a fielded system would require finalizing a composite turf cover-layer design and calibrating a predic- tive model to match the response. Characterization would be advisable for the soil layer under various freezing condi- tions to assess the impact on arresting dynamics. Additionally, investigation should be made regarding the basin geometry to determine whether above- or below-grade construction is preferable. Because the material performance can be affected by size gradation and compaction, development of a suit- able installation process would be required. This installa- tion process would be coupled with a predictive model matching the as-installed material characteristics. Finally, full-scale testing is advisable for evaluation of the complete system. 136

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TRB’s Airport Cooperative Research Program (ACRP) Report 29: Developing Improved Civil Aircraft Arresting Systems explores alternative materials that could be used for an engineered material arresting system (EMAS), as well as potential active arrestor designs for civil aircraft applications. The report examines cellular glass foam, aggregate foam, engineered aggregate, and a main-gear engagement active arrestor system.

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