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Developing Improved Civil Aircraft Arresting Systems (2009)

Chapter: Chapter 9 - Glass Foam Arrestor Concept

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Suggested Citation:"Chapter 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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 9 - Glass 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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63 9.1. Concept Description 9.1.1. System Overview Glass foam is a crushable low-density material that has been proposed as an arrestor (Figure 9.1). The foam has a closed- cell microstructure that prevents water absorption and makes it an excellent thermal insulator. Glass foam has material properties that suggest excellent durability to the environment and good chemical resistance. Typically manufactured in solid blocks of various sizes, the foam can be cut into a variety of shapes for different applications. Because both materials are low-density crushable foams, the glass foam material exhibited properties similar to cellular cement. However, glass foam in general appeared to be less fragile, easier to handle, and potentially more durable than the cellular cement. Additionally, adhesives and moisture sealants are available for glass foams that permit joining and weatherproofing. Because the glass foam material is generally manufactured in blocks sized at approximately 24 × 18 × 6 inches, two vari- ants are possible (Figure 9-2): 1. Block Method. The block method would use 4-ft square blocks of the foam, analogous to the current EMAS con- struction approach. The blocks would be constructed by adhering multiple smaller blocks together, followed by the potential addition of top and/or bottom cap materials. The sides of the block could potentially be sprayed with a sealant to further weatherproof the blocks. These blocks would be transported to and installed at a runway in essen- tially the same manner as the current EMAS beds. 2. Monolithic Method. The monolithic method would be assembled on-site at the runway by stacking and gluing blocks into a single large structure. The final assembly would then be fitted with a continuous top cover layer composed of a roll/spray-on polymer. This layup would preclude the need for joint seams, sealants, and maintenance, which are required for the current EMAS design. Moisture sealing of the vertical sides of individual blocks would be unnecessary. Monolithic layups such as this have been used in building roof applications. 9.1.2. Performance Considerations Because glass foam material is a lightweight crushable foam bearing many similarities to cellular cement, the list of perfor- mance considerations at the outset was relatively short. Issues for evaluation included: • Compression performance of the material in terms of energy absorption and rebound characteristics, • Required density/strength for effective arresting, • Balance of compression and shear strengths of the material, • Rate dependence of the material, and • Durability to freeze–thaw exposure. 9.2. Testing and Modeling Approach The goal for the performance evaluation was to perform testing that would allow calibration of high-accuracy computer models of the glass foam concept. The testing and modeling approach for the glass foam material was comprehensive in nature, and the final outcome was a well-calibrated numerical model for predicting arrestment loads. The testing and modeling approach for the glass foam concept is illustrated by Figure 9-3. Five major stages are illustrated by the larger process bubbles of the chart: 1. Arrestor Material Testing and Modeling. Laboratory and pendulum 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. C H A P T E R 9 Glass Foam Arrestor Concept

64 Figure 9-1. Glass foam material: cylinder test specimen (left) and close-up of cellular microstructure (right). Glass Foam Arrestor Concept 1: Block Method Seams Between Blocks, Sealed with Tape/Caulk Each Major Block is a Glued Assemblage of Smaller Blocks Lid Glass Foam Arrestor Concept 2: Monolithic Method Monolithic Sealant Top Layer without Joints Small Blocks Glued Together in Brick Pattern Figure 9-2. Glass foam arrestor variants: block method (top) and monolithic method (bottom).

3. Aircraft Modeling. A generalized aircraft model was developed 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 distances 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 determining the loads exerted on the different aircraft tires by the arrestor bed. Large data sets were generated using simulation 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 development stages culminated in the final, bottom-most process on the figure. The APC was used to predict arresting distances, landing gear loads, and ideal arrestor bed designs for the different aircraft. 65 (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 9-3. Testing and modeling process for glass foam arrestor evaluation.

The subsequent sections of this chapter will focus on areas (1), (4), and (5). Specific attention will be given to the glass 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. 9.3. Testing Effort The testing effort for the glass foam material involved an extensive battery of mechanical and environmental tests. Table 9-1 depicts the overall test matrix for the glass foam material. All cylinder dimensions specify diameter followed by height. Conditioned specimens were environmental test specimens that were compression-tested following freeze–thaw exposure. These cylinders were only 2.5 in. tall to maintain the aspect ratio of prior “short” cylinder specimens. The specimen diameter was constrained by the environmental test apparatus. 9.3.1. Density and Dimension Measurements Several densities of the glass foam material were examined. The higher density samples had higher compressive strengths. Preliminary screening indicated that the lowest density spec- imens, at nominally 6 pcf, were most promising for this appli- cation. The remainder of the testing discussion is confined to this density of the material. Specimens were obtained in cylindrical shapes with 3.65- and 5.625-in. diameters and larger 24 × 18 × 5-in. blocks. 9.3.2. Platen Compression Tests 9.3.2.1. Low Rate Platen compression tests were performed for the initial mat- erial density screening, enabling a rapid down-selection to the 6 pcf density as the most likely candidate. The platen compres- sion tests further permitted the evaluation of the energy absorp- tion capacity of the material and the effects of loading rate. 66 Test Properties Characterized Detail Number of Tests Laboratory Tests Hydrostatic Triaxial Compression Test • Compressive strength at failure (σu) • Shear strength (τu) • Per ASTM D2850 • 3.65 x 8” cylinder • 0.0592 in./min compression rate • Maximum compression of 5 to 10% • Confining pressures of 0, 5, 10, and 20 psi 0 psi 5 psi 10 psi 20 psi Total: 1 2 2 2 7 • Version 1: low speed, tall • Non-standard • 3.65 x 8” cylinder • 3 in./min compression rate • Maximum compression of 85% • Unconfined flat disk specimen compression Fresh Capped: Uncapped: Conditioned Uncapped: 2 3 2 • Version 2: low speed, short • Non-standard • 5.625 x 4” cylinder • 3 in./min compression rate • Maximum compression of 85% Capped: Uncapped: 2 3 Parallel Platen Compression Test • Compressive strength at failure (σu) • Compressive stress– strain curve • Extrapolated: volumetric energy capacity • Determined for different strain rates • Version 2: high speed, short • Non-standard • 5.625 x 4” cylinder • 60 in./sec compression rate • Maximum compression of 70% Capped: Uncapped: 2 3 Punch Compression Test • Combined compression and shear strain behavior • Somewhat similar to EMAS material testing method • Non-standard • 5” thick material block • 1.5” punch diameter • 3 in./min compression rate • Maximum compression of 70% 5 Environmental Chamber Tests • Durability to freeze– thaw cycles • Effectiveness of sealant treatments, where applicable • Per ATSM C 666/C 666M-03 (modified) • 3.65 x 2.5” cylinder • 78 freeze-thaw cycles 2 Pendulum Test Pendulum One- Wheel Bogey Test • Dynamic response for the material to a moderate-speed tire overrun characterized • Non-standard • 16.9 mph collision • 24 x 18 x 5” blocks, glued into stacks 2 Table 9-1. Test matrix for glass foam material.

Low-rate platen tests were conducted by placing a cylindrical specimen between two metal platens. The upper platen was then displaced downward at a fixed rate of 3 in. per minute until the specimen was fully compressed. The platen tests were conducted with and without Sor- bothane capping material on the top and bottom of the spec- imens to determine whether capping was beneficial or not. For the uncapped specimens, the material crushed in a local- ized failure zone immediately adjacent to the platen. These specimens maintained their shape throughout the duration of the loading up to the final 85% compression state, which aided in capturing the full energy absorption of the material (Figure 9-4). When caps were used, the specimens tended to split along vertical planes, causing a shape change that often led to early terminations of the compression test (Figure 9-5). The un- capped approach was eventually favored due to its superior test repeatability and its ability to capture the full energy capacity of the material. Tall and short cylinder specimens were tested for compar- ison with the hydrostatic triaxial tests, which used tall cylinders (3.65 × 8 in.), and the high-rate platen tests, which used short cylinders (5.625 × 4 in.). The glass foam material exhibited a characteristic crush- able foam load history that rose to a plateau value, where it remained until the material approached full compression (Figure 9-6). Near full compression, the material hardened and the loading increased. The stress–strain curve was integrated to produce the energy absorption curve shown in the figure. No measurable rebound occurred after compression. 9.3.2.2. High Rate High-rate compression tests (60 in.-s) were conducted on short cylinders (5.625 × 4 in.) to evaluate the effects of rapid compression on the material. For a B747 main-gear tire travel- ling at 70 knots, the average vertical compression rate would be 40 in./(in.-s). The test produced an average strain rate of 150 in./ (in.-s), which exceeds the overrun case by a substantial margin. When comparing the high- and low-rate loading curves, only a minimal difference is observed, which could be due only to inertial effects of the material. Overall, the high-rate platen 67 Figure 9-4. Low-rate uncapped platen test of glass foam pre-test (left) and post-test (right). Figure 9-5. Mid-test splitting of low-rate capped platen test cylinder of glass foam.

tests indicate that there is little to no practical rate effect in the material for the loading regimes of interest (Figure 9.7). 9.3.2.3. Hydrostatic Triaxial Tests The hydrostatic triaxial tests evaluated the glass foam per- formance at different confining pressures to determine if any strength increase or bulking deformation took place. Neither was anticipated because the glass foam material behaved as a one-dimensional foam with little to no observed lateral bulking (low Poisson ratio). The specimens were tall cylinders (3.65 × 8 in.) placed between platens using Sorbothane caps at the top and bottom (Figure 9-8, left). The specimens were fitted with flexible membrane sleeves before immersion in a pressurized vessel of water. While at this hydrostatic pressure, the specimens were compressed axially until failure. As observed in the platen tests, the use of capping induced vertical failure plane formation in the specimen (Figure 9-8, right), which typically occurred at a compression displacement of about 5%. This failure load did show a mild dependence on 68 Strain [in./in.] St re ss [p si] En er gy A bs or pt io n [p si/ in2 ] 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 0 10 6 0 20 12 30 18 40 24 50 30 60 36 70 42 80 48 Stress, Average of Test Data Energy Absorbed Per Unit Volume Figure 9-6. Average load history for glass foam material from low-rate platen tests, tall specimen. Strain [in./in.] St re ss [p si] En er gy A bs or pt io n [p si /in 3 ] 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0 10 8 0 20 16 30 24 40 32 50 40 60 48 70 56 Stress, Average of Test Data Energy Absorbed Per Unit Volume Figure 9-7. Average load history for glass foam material from high-rate platen tests, short specimen. Sorbothane Cap Figure 9-8. Hydrostatic triaxial compression test of glass foam pre-test (left) and post-test (right).

confinement pressure, though no bulking behaviors were observed. 9.3.3. Punch Tests Punch tests were performed by pressing a 1.5-in. diameter, smooth-sided punch into a 5-in. thick block of the material (Figure 9-9). The glass foam material compressed cleanly, leaving a smooth-sided hole in the block. These tests were similar to the characterization tests performed by the current EMAS manufacturer (34). However, the tests differed with regard to the proportional aspect ratio of the punch, the punch shape, the specimen size, and the loading rate. The resulting load was due to a combination of the com- pression strength of the material (area under the punch) and the shear strength of the material (circumferential edge of the punch). Figure 9-10 illustrates the averaged load history for the punch tests. As shown, the effective “pressure” on the 69 Figure 9-9. Punch test of glass foam test specimen pre-test (left) and post-test (right). Strain (in./in.) Pr es su re (p si) En er gy A bs or pt io n (p si/ in3 ) 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 0 0 20 10 40 20 60 30 80 40 100 50 120 60 140 70 160 80 180 90 200 100 Compression Force Average Energy Absorption Figure 9-10. Average load history for glass foam material from punch test with stress normalized for cross-sectional area of 1.5-in. diameter punch.

punch was nominally 140 psi, which compares at 2.5 times the nominal 55 psi strength of the material exhibited in the platen tests (Figure 9-6). This increase in apparent strength is due to confinement and shear contributions. In arrestor applications, the shear strength of the material would be rele- vant along the vertical walls of the tire rut. 9.3.4. Pendulum Tests 9.3.4.1. Pendulum Apparatus In addition to the small-scale laboratory tests, a larger-scale one-wheel bogy test was conducted using a large pendulum test apparatus. The pendulum test apparatus featured a heavy 4,400-lb mass that hung from an overhead support frame, giving it a swing arc of 24.5 ft. The mass was hoisted to the desired height and then released; the speed of the mass was controlled by the release height. The pendulum mass was fitted with a strut and wheel assembly, which Figure 9-11 illustrates in an exploded view. The strut was instrumented with three load cells that measured loads at the connections. These three connection loads were resolved into orthogonal vertical and drag loads on the strut and wheel assembly. To reduce the number of variables in the design, a rigid aluminum wheel form was used rather than a pneumatic tire. The diameter and width were 14.8 and 5.5 in., respectively. These proportions were based on a nominal one-third scale B737-800 main-gear tire, which has a diameter and width of 44.5 and 16.5 in., respectively. 9.3.4.2. Glass foam retention box. Below the pendulum assembly, a box was constructed to retain larger blocks of the glass foam, which each measured 24 × 18 × 5 in. The blocks were arranged in one row, 6 blocks long and 2 blocks in depth, for a total of 12 blocks. The over- all dimensions were 9 ft in length, 2 ft in width, and 10 in. in height. The upper and lower blocks were glued in 6 pairs, which were placed in the retention box and held in place by an upper cap rail around the perimeter. The block pairs were not adhered to one another. Figure 9-12 shows the overall pendulum apparatus with the glass foam blocks beneath it. 9.3.4.3. Tests Executed For the tests, the pendulum was set to swing such that the wheel penetrated to a depth of nominally one-third diameter, or 5.0 in. One swing height was used to produce an average overrun speed of 16.9 mph. Due to the single-use nature of the material and the limited supply of evaluation samples, two tests were conducted for these conditions. Both tests gave consistent results. Figure 9-13 shows the rut created through the material by the strut and wheel of the pendulum; the wheel cut a clean path through the material, leaving the RG-1 material adjacent to the rut essentially undamaged. This behavior was consistent with expectations based on the small-scale testing. 9.3.4.4. Landing Gear Loads The loading history for the pendulum strut showed oscil- lations in both the drag and vertical loads (Figure 9-14). It is believed that these pulses were caused by the seams between the block pairs in the bed. As the wheel approached the end of a block, the load decreased; when the wheel began to overrun the next block, the load increased. Since the blocks were not glued at the seams, these joints presented a discontinuity in the material. The pulsing effect was further amplified by the 70 Figure 9-11. Pendulum test device with one-wheel bogy.

71 Pendulum Mass Strut and Wheel Assembly Retention Box with Glass Foam Blocks Cap Rail Figure 9-12. Overview of pendulum test setup for glass foam. Figure 9-13. Post-test results from pendulum test for glass foam.

rigid wheel used in the strut assembly, which prevented natural smoothing that might result with a pneumatic tire. In arrestor applications, the severity of the loading pulses depends on several variables: • Relative compressive strength of the material, • Flexibility of the pneumatic aircraft tire, • Relative penetration depth of the tire, and • Glued or non-glued approach to joints in the material. In applications involving a flexible tire, or where the blocks are glued at all joints, the loading pulses are expected to smooth substantially. However, the appearance of these pulses gener- ated questions regarding the feasibility of using separate blocks of the material in an analogous manner to the current approach for EMAS. Whether or not these pulses would occur during overruns into the existing EMAS beds is unclear. The nature of the pulses suggests that an arrestor design using a foam block material should explicitly include this seam effect; it is not sufficient to make a general assumption that separate blocks essentially give the same loading as a continuous bed of the material. 9.3.5. Environmental Tests A basic set of environmental tests was conducted to determine the necessity for weatherproofing the glass foam material. Two 3.65 × 2.5-in. cylinders were subjected to fully immersed freeze–thaw testing, per ATSM C 666/C 666M-03. The specimens were subjected to 78 freeze–thaw cycles, during which they absorbed water and partially eroded (Figure 9-15). (Standardized testing generally uses 300 freeze–thaw cycles. Due to the test duration required, this was abbreviated to a planned 50 cycles for this effort. An unplanned cycling overrun at the test facility resulted in a total of 78 cycles.) Following the environmental tests, the specimens were sub- jected to a low-speed platen compression test (Figure 9-16) to determine the performance degradation. When compared with the fresh material, the samples exhibited a 60% decrease in energy absorption capacity and compression strength. Mechanically, the closed cell foam limits water absorption, such that water penetrates only the outer-most open pores of the foam. Upon freezing, the expanding water cracks the cells, permitting progressively deeper penetration 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 are fully immersed in water, without normal countermeasures of drainage, protective packaging, or sealants. Information provided by the manu- facturer indicates that cyclical temperature and humidity do not degrade the material over time, and a number of sealants are available to prevent water absorption, if required. Overall, these tests indicate that the glass foam material should be protected from immersion conditions caused by standing water, as is done for the current cellular cement material. Additional testing could be conducted to characterize durability in non-immersion scenarios, or in immersion con- ditions where a sealant has been applied to the material. 9.4. Modeling Effort The modeling effort involved several stages, as shown pre- viously in the flowchart of Figure 9-3. A high-fidelity model for the glass foam material was calibrated to match the test data (Figure 9-3, block 1). Using this material model, an arrestor 72 Time (s) Fo rc e (lb f) 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0 1000 2000 3000 4000 5000 6000 7000 8000 Drag Force Vertical Force Figure 9-14. Loading history for glass foam pendulum test. Figure 9-15. Glass foam specimen after environmental freeze–thaw testing.

bed model was constructed and coupled with tire models for the different aircraft (Figure 9-3, 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 glass foam arrestor concept are reserved for the following section (9.5). 9.4.1. Smoothed Particle Hydrodynamics (SPH) Formulation The glass foam arrestor models were developed in LS-DYNA, a general-purpose finite element modeling code. Within LS-DYNA, a number of formulations exist for representing solids and fluids. Due to the high compressibility of the glass foam material, 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. These particles are not dis- jointed pieces of aggregate, but are instead mathematically inter- connected to represent a continuous solid material (Lagrangian formulation). 9.4.2. Calibration to Physical Tests 9.4.2.1. 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 9-2. The calibration process required defining these material parameters such that the model performance matched that of the physical material tests. 9.4.2.2. Multi-Tester Model A multi-tester model was constructed to simultaneously replicate the four laboratory material tests (Figure 9-17). The material parameters of the multi-tester model were optimized using LS-OPT, an optimization software package. LS-OPT ran the simulations in batches iteratively; after each iteration, it narrowed the region of interest, effectively zooming in closer to the predicted optimum calibration point. After 8 iterations of 12 simulations each, the design was optimized for a best-fit set of material parameters. Table 9-3 gives a summary of the calibration process, includ- ing the final accuracy of the calibrated model. Test metrics marked with an asterisk were optimization criteria, which LS-OPT attempted to minimize. The remaining metrics were measured, but did not act as optimization criteria. Of the various metrics given, the energy absorption values were the most critical to match accurately. As shown, all energy 73 Figure 9-16. Platen testing of glass foam environmental test specimen. 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 9-2. Parameters for *MAT_063 or *MAT_CRUSHABLE_FOAM.

74 Shell membrane with 10 psi confinement Punch Specimen Hydrostatic Triaxial Specimen High Rate Platen Specimen Low Rate Platen Specimen (1/4 Symmetry) Specimens After Compression Figure 9-17. Multi-tester model for glass foam material calibration. Table 9-3. Multi-tester specifications for glass foam material calibration. Test to Replicate Description Accuracy of Calibration • 3.65 x 8-in. cylinder • Match stress–strain load curve with root-mean-squared error (RMSE) to 85% compression* 7.0% Low-Speed Platen Test • Match energy absorption at 85% compression* 5.0% • 5.625 x 4-in. cylinder • Match stress–strain load curve with RMSE to 50% compression* 17.7% High-Speed Platen Test • Match energy absorption at 50% compression* 2.9% • 3.65 x 8-in. cylinder • 10 psi hydrostatic pressure Hydrostatic Triaxial Test • Match stress at 5% compression 19% • Large block • Match stress–strain load curve with RMSE to 70% compression 8.4% Punch Test • Match energy absorption at 70% compression 1.2%

metrics were within 5% of the actual test values. The matching of the stress–strain load curves proved less consistent, which was expected since those curves tended to vary among the test specimens as well. The hydrostatic triaxial specimens proved somewhat dif- ficult to calibrate due to a nuance of the SPH formulation, which proved challenging to fit with a hydrostatic membrane load. Since the glass foam material exhibited little pressure dependence in testing, this was deemed a low-priority calibra- tion point. 9.4.2.3. Pendulum Model Using LS-DYNA, a model was constructed to replicate the pendulum tests (Figure 9-18). Because the material had been well-calibrated via the multi-tester, the pendulum model was used for validation of the material model, rather than for calibration. The pendulum strut was omitted because the penetrate depth of the actual test only allowed the wheel to contact the glass foam material. The foam bed was constructed with half-symmetry, such that it measured 9 ft in length, 1 ft in width, and 10 in. in depth. The bottom and outer sides of the bed were constrained to simulate the presence of the confining box. The wheel followed an arced path approximating that of the actual strut, with the same 1⁄3-diameter penetration depth into the arrestor bed. The wheel began with no rotation and was allowed to freely spin upon contact with the arrestor bed, just as in the actual test (Figure 9-19). The SPH particles were sized to 0.75 in., which was based on particle size scaling relationships developed in other simulation sets to maintain reasonable accuracy. This provided four par- ticles across the half-width of the wheel, which was the smallest characteristic length of the interface. A separate particle size convergence study was not undertaken for this model. The resulting loads matched the test data well, though no pulses were observed due to the lack of material seams in the model (Figure 9-20). The average drag and vertical loads were within 6% and 1% of the average test results, respectively. Since the test data showed a 6 to 10% variation for these two metrics, the model is within the experimental data scatter. Overall, the pendulum model validated that the glass foam material was well calibrated. 9.4.3. Tire and Arrestor Simulations Using the calibrated glass foam material model previously described (Section 9.4.2), a large-scale arrestor model was created in LS-DYNA to simulate overruns by aircraft tires. No 75 Figure 9-18. Overview of glass foam pendulum model. Figure 9-19. Action sequence from glass foam pendulum model. Time (s) Fo rc e (lb f) 0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0 2000 4000 6000 8000 Test Drag Force Test Vertical Force Model Drag Force Model Vertical Force Figure 9-20. Comparison of test and model load histories for glass foam pendulum test.

protective cover layer for the bed was included in the model; it was assumed that a well-designed cover layer (likely made of thin plastic or a spray-on polymer) should have minimal impact on the mechanical response. The model further assumed a continuous material, and material seams were not included. Figure 9-21 illustrates the model with a 36-in. depth and a B737-800 main-gear tire (Goodyear H44.5 × 16.5) at 50% penetration depth. 9.4.3.1. 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 minimal 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 225 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 movable rigid plane (Figure 9-22). 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. 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 discretiza- tion error. For the larger tires, a 2 in. particle size was found to have less than a 4% error for the predicted drag and verti- cal loads. For smaller tires, the particle size was reduced to 1 in. to maintain similar size proportionality. Based on the parti- cle and bed size variations, a typical bed model had nominally 45K particles. 76 Arrestor and Tire Model Uses Half- Symmetry Small Pieces of Glass 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 9-21. Model of combined tire and glass foam arrestor system.

9.4.3.2. Tire Models The tire models were fully deformable finite element models (FEM), as discussed in Appendix F. Table 9-4 summarizes the tire models developed. 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 during 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. 9.4.3.3. Sequencing of Simulations Because the tires in the LS-DYNA simulations were deformable and were allowed to spin freely, a sequencing method was required to create stable, fast-running simula- tions. These two factors were additional complications that were not present in the EDEM software package 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 measured 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 distance was required for the loads and rotation to reach steady-state con- ditions 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 begin- ning the forward motion, and • Applying the forward motion before or after making contact 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 repeatedly in large batches, it was important to develop a sequencing method- ology that would produce efficient simulation run times. Multiple methods were attempted through experimentation before settling on the approach illustrated by Figure 9-23. The tire was first pressed downward into the material to a prescribed depth. Then the tire was accelerated to the desired forward 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, allowing 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. 77 36-inch Deep Bed Rigid Plane at 18- inch Depth Only Upper Part of Material Involved in Compression Figure 9-22. Adjustable height of glass 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 9-4. FEM tire library for glass foam arrestor models.

9.4.4. 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 assembled into “metamodels” for uploading and use by the APC. 9.4.4.1. 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 metamodel data fit.); • Bed depth, in incremental depths from 3 to 36 in. (Fig- ure 9-24); and • Penetration into the bed, from 10% to 100% of maximum penetration depth. Due to the two sources of compression (arrestor material and tire), the definition for penetration depth was more complex than for the rigid tire approach used for the aggregate arrestor models. Two conditions defined the maximum penetration depth: 78 Tire presses down into arrestor Tire accelerates forward and is given an initial spin rate Tire continues forward and is allowed to f reely spin Steady-state vertical and drag loads are measured Tire begins above arrestor Figure 9-23. Sequencing method for glass foam arrestor model. Bed Depth Bottoming Depth Rut DepthPenetration Depth Tire Deflection Figure 9-24. Depth definitions for glass foam bed models.

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. (This depth issue ultimately proved irrelevant, since most functional arrestor bed designs generated with the APC did not use bed depths that were greater than the tire diameter.) The large batch simulations were conducted using LS-OPT. Based on the initial model files, LS-OPT generated permuta- tions with various speeds, bed depths, and penetration 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 itera- tions were added to improve accuracy as needed. 9.4.4.2. 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. Metamodeling is analogous to fitting a curve through experimental 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 radial basis function (RBF) networks, which can effectively capture non-linear behaviors including multiple concavity changes across the data set. Table 9-5 summarizes the fit quality for the metamodels. The root-mean-squared (RMS) error was typically below 5%, and the R-squared value was typically above 0.98, indicating good fit quality with minimal noise. 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 penetra- tion depths. 9.4.4.3. Parameter Sensitivities Using the metamodels, it is possible to review how sensitive the landing gear loads are to the different variables of speed, bed depth, and penetration percentage. Figure 9-25 shows a surface plot for the 44.5-in. main-gear tire of the B737 in an 18-in. deep glass foam arrestor bed. Stronger drag loads (shown as the lower, more negative values) occur when the penetration ratio increases, up to the maximum of 1.0 (or 100%). The loading is, therefore, strongly dependent on the depth of penetration into the bed, as would be expected. By contrast, the variation with speed shows very little change between 10 and 70 knots. The loading is fairly insen- sitive to speed, reflecting the low rate-sensitivity that was exhibited during the small-scale lab testing. Practically speak- ing, this means that the glass foam system will exert nearly the same deceleration 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 EMAS material. 9.4.4.4. 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. 79 Tire SimulationsConducted Points Used Response RMS Error R 2 Drag 2.35% 0.999 H49 49 49 Vertical 5.25% 0.988 Drag 2.78% 0.999 H44 70 67 Vertical 1.08% 0.999 Drag 2.78% 0.998 H29 60 60 Vertical 4.88% 0.986 Drag 3.21% 0.998 R27 60 47 Vertical 5.06% 0.982 Drag 3.74% 0.995 H18 100 86 Vertical 5.83% 0.969 Table 9-5. Metamodel accuracy summary for glass foam arrestor bed.

A MATLAB conversion program was written to map this data into multi-dimensional matrix form that could be quickly accessed by the APC. 9.5. Arrestor Performance Predictions 9.5.1. 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, • 70-knot starting speed for the aircraft, • No reverse thrust, • Braking factor of 0.25 before and within the bed, • Cover layer had negligible effect, • Material had no seams (Section 9.4.3), and • Arrestor bed loads based on interaction with tires, neglecting 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 glass foam arrestor design functioned best as a partially recessed bed, such that the fully crushed material depth was level with the runway. This left 85% of the bed thickness above grade, and 15% below grade. This approach produced the smoothest landing gear loads by limiting the effective step-up or step- down that the aircraft experienced upon entering the bed. Two design variables were considered for the aircraft: the bed depth and the material strength. The material strength was adjusted by applying a scale factor to the metamodel loading data in the APC during a simulation. It was considered as an 80 Figure 9-25. Metamodel drag load surface plot for 44.5-in. tire in an 18-in. deep arrestor/turf bed. D ra g Fo rc e (lb f) Speed (knots) Penetration Ratio

open variable because the glass foam can be manufactured at a variety of strength levels (varying density). 9.5.2. Performance for Test Aircraft Table 9-6 lists best-case arrestor designs for each aircraft taken individually. Each arrestor bed listed uses a different material strength and depth that are optimized for the design aircraft. Generally, a range of acceptable strength and depth combinations was available. Compared with the similar EMAS design cases on the right (provided by ESCO), the distances are comparable if the ultimate loading criterion is used. Table 9-7 shows the compromise design case with the best arrestor design for all three aircraft. The CRJ-200 limits the bed depth in this case, while the B737 controls the bed length. With the material strength and depth as specified, the B747 would require 840 ft to decelerate from 70 knots. Since this is longer than a standard RSA, it is highly impractical. 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 assume a 400 ft length, which is sufficient to arrest the two smaller aircraft with 70-knot exit speeds. At this length, the B747-400 would have a maximum exit speed of approximately 46 knots, which satisfies the requirements of AC 150-5220-22a. 9.5.3. General Observations The overall deceleration and loading trends on the three aircraft showed several common characteristics. Figure 9-26 and Figure 9-27 illustrate sample deceleration and nose-gear load plots generated by the APC for the B737 aircraft. Figure 9-26 represents the arrestor bed using the limit load design criterion, while Figure 9-27 is based on the ultimate load criterion. The overall bed lengths and depths reflect the values of Table 9-6. 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 15% of the bed thickness, as discussed in Section 9.5.1. 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 70 feet. After the initial transition, the deceleration settles to a fairly constant value in both cases, at about 0.6 g for the limit design and about 0.8 g for the ulti- mate design. The relatively constant deceleration value is pos- sible because the glass foam material shows only mild rate dependence. An ideal arrestor bed would provide a perfectly constant deceleration; the glass foam bed performance approx- imates the ideal bed fairly well. 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 50 and 125 ft, which is after the main-gear tires have entered the bed. When the main-gear tires enter, the deceleration increases substantially, 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. 9.5.4. 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 81 Nose-Gear Limit Load Criterion Nose-Gear Ultimate Load Criterion Current EMAS, Optimal Designs Aircraft Strength (psi) Depth (in.) Bed Length (ft) Strength (psi) Depth (in.) Bed Length (ft) Depth (in.) Bed Length (ft) CRJ-200 17 20.0 294 25 19.9 243 22 258 B737-800 15 25.6 388 20 30.0 302 22 287 B747-400 47 36.0 409 53 36.0 406 28 495 Table 9-6. Individual aircraft 70-knot arrestor beds for glass foam arrestor system. Nose-Gear Ultimate Load Criterion Bed Dimensions 26 psi material 19.1 in. depth 400 ft long Aircraft Exit Speed (knot) Stopping Distance (ft) CRJ-200 70+ 244 B737-800 70+ 338 B747-400 46 400 Table 9-7. Fleet design arrestor bed for glass foam arrestor system.

82 -10 0 10 20 30 40 50 60 70 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 0.1 0.2 0.3 0.4 0.5 0.6 0.7 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 Decleration 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 9-26. Limit criterion glass foam arrestor design plots for B737-800 showing speed (top), deceleration (middle) and nose-gear drag load (bottom).

83 Aircraft Speed Upper Surface of Arrestor Lower Surface of Arrestor Aircraft Deceleration Nose Gear Drag Nose Gear Limit Load Nose Gear Ultimate Load -10 0 10 20 30 40 50 60 70 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 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 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 Decleration 0 5 10 15 20 25 30 35 40 Fo rc e (ki p) Landing Gear Forces - NOSE STRUT Figure 9-27. Ultimate criterion glass foam arrestor design plots for B737-800 showing speed (top), deceleration (middle) and nose-gear drag load (bottom).

main-gear tires was due to the braking plus the arrestor resis- tance. As an approximate solution, this approach worked well and was appropriate. However, a small side study conducted during the meta- modeling 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 would not be captured. For a bottomed tire, little increase is feasible and this effect would likely not occur. For a non-bottomed tire, however, 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 deceler- ation process while not affecting the nose gear. Stopping dis- tances could be somewhat reduced by this effect, though it is unclear by what amount or in what limited subset of arrestor design cases. 9.5.5. Effect of Material Seams As previously discussed, seams in the glass foam material produced notable pulses in the pendulum test strut loading (Section 9.3.4.4). However, the metamodels for the material assume a monolithic arrestor bed without seams. If seams were present in the material, such as in the design concept featuring separate blocks (Section 9.1.1), they could impart pulsed drag loads to the nose gear. Pulsed loads would require a reduction in the bed depth to ensure that the peak loads did not exceed the limit or ultimate thresholds for the nose gear. Consequently, the arrestment distances would increase. Since the effect of the material seams in a fielded system has not been established, the caveats of Section 9.3.4.4 apply to these observations. 9.6. Estimated System Cost and Upkeep 9.6.1. Installation Process The arrestor beds would likely use sloped entry and terraced sides regions, with a slight vertical depression below the runway level (described further in Section 9.5). For the block variant of the system, the glass foam would be manufactured and assembled into the larger blocks while at the production facility. These blocks would then be transported to the airport and laid in place in the RSA. Because of the lower density of the material (∼6 pcf), a 4-ft block may only weigh 200 lbs; this could permit placement of the blocks using pallet jacks or dolly-type devices rather than forklifts. After placement, the seams would be sealed using joint tape or caulk. For the monolithic variant, the glass foam would be shipped to the site without prior assembly. The small blocks would then be hand-laid in rows and layers, with glue applied in between. After the monolithic bed was completed, a roll-on or spray- on polymer would be applied to the surface to seal the bed. The hand-laying process would require more time than plac- ing larger blocks; however, the final bed would have reduced maintenance needs thereafter due to the lack of joints. 9.6.2. Cost to Establish System A preliminary estimate was made for the cost to establish a glass foam arrestor system. It must be noted that the cost esti- mate from this section is only a basic approximation for the purposes of comparing the different arrestor alternatives. 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 installation effort. Where possible, the methodologies used were consistent with the prior survey information collected regarding the existing EMAS (Section 3.5). The glass foam arrestor concepts would require site prepa- ration and paving similar to the current EMAS. It was assumed that the costs for this preparation would be identical to the site preparation cost for an EMAS. In terms of manufacturing cost, lower density versions of the material are more expensive than the commercially avail- able insulating product. To provide a conservative cost estimate, the upper bound for the possible cost range was assumed. The installation cost estimate was separated into specific materials and general installation labor needs. Because these costs were specific to the glass foam arrestor concepts, they do not have a direct connection to any prior EMAS data. Discus- sions with a potential manufacturer produced cost estimates for the foam blocks, adhesive to join the blocks, and a polymer top coating material. The labor costs were estimated based on similar labor needs for installation of glass foam in roof applications. 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 9-8 and Table 9-9, respectively. Using the survey cost assumptions of Table 9-8, a 300-ft arrestor bed would cost approximately 7% less than the cur- rent EMAS design. If the Order 5200.9 costs are assumed, the glass foam system would be nominally 37% more expen- sive (Table 9-9). Due to the preliminary nature of the esti- mated system cost, it is feasible that actual costs for the system would be higher based on unforeseen elements. A conservative approach would conclude that the cost for this concept will probably be similar to that of EMAS. In addition to the tables in this section, longer-term life-cycle issues could also be considered. FAA Order 5200.9 includes a 84

standard 10-year replacement interval for an EMAS, which translates into present-value life-cycle costs. Such a replace- ment could arguably be unnecessary for this arrestor concept (Section 10.6.4). Eliminating the assumed 10-year replacement could effectively trim about $2.6M of present-value life cycle costs (based on the EMAS replacement cost estimates of the survey). 9.6.3. Maintenance The glass foam material is expected to have a superior durability when compared with cellular cement. Glass foam does not exhibit the tendency to crumble during handling that 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 that long service life is possible with little degradation as long as mild protective measures are taken. In the monolithic version of the concept, joint tape/caulk maintenance would not be required. Annual maintenance could be limited to upkeep of the required yellow chevron markings on the arrestor. 9.6.4. Replacement and Overhaul It is not anticipated that replacement of the bed would be required after 10 years, as anticipated for an EMAS in FAA Order 5200.9. Based on other fielded applications for glass foam material, life cycles of 20 years should be realistic, with some past applications indicating that it can last as long as 50 years. This effectively reduces the life-cycle cost of the concept. Since the installation costs are expected to be similar to the existing EMAS, the elimination of a 10-year bed replacement would effectively trim $3.5M of present-value life-cycle costs. However, the top cover layer applied to the material must also be considered. Polymer coatings and plastic lids can degrade when exposed to ultraviolet radiation and varying weather conditions. Overhaul may be required to replace or repair the top material periodically. 9.6.5. Repair Repairs after an overrun would parallel those of the current EMAS design. Damaged sections of glass foam material would require removal and replacement after aircraft extraction. 9.7. Transition to a Fielded System In order to transition the glass foam concept to a fielded system, the following additional development steps may be advisable. 9.7.1. Low Material Density Calibration As shown in the example arrestor beds of Table 9-6 and Table 9-7, the compressive strengths for optimal designs ranged from 17 to 53 psi. Since the material tested and cali- brated during the research had a nominal strength of 55 psi with a 6-pcf density, the fielded material will be softer and lighter. To estimate the needed density and strength, the metamodel data was simply scaled, which was an appropriate simplification. However, in migrating to a fielded system, several standard densities would probably be selected to provide flexibility dur- ing the design of actual arrestors (currently done with EMAS). For each density, some testing would be required to generate calibration data. The foam glass material model would require minor adjustments to compensate for the alterations. Fresh metamodel data should then be generated such that data scaling would no longer be required in the arrestor design process. 9.7.2. Cover Layer Design Two major cover-layer design concepts are possible: 1. Plastic lids, applied to the block version of the concept, or 2. Spray-on or roll-on polymer coating, applied to the mono- lithic concept. 85 Cost Category Glass Foam System Current EMAS Site Preparation $ 2.17 $ 2.17 Installation $ 5.49 $ 6.03 Cost to Establish $ 7.65 $ 8.19 Percent of EMAS 93% Table 9-8. Estimated costs to establish glass foam arrestor, 150 x 300 ft, assuming survey average costs for current EMAS, units of millions USD. Cost Category Glass Foam System Current EMAS Site Preparation $ 0.68 $ 0.68 Installation $ 5.49 $ 3.83 Cost to Establish $ 6.17 $ 4.50 Percent of EMAS 137% Table 9-9. Estimated costs to establish glass foam arrestor, 150 x 300 ft, assuming Order 5200.9 costs for current EMAS, units of millions USD.

For either case, the material type and thickness of the layer must be established. The material should be selected based on ultraviolet and weather durability. A material that can be dyed to a pavement gray color would be preferable, such as is currently in use for plastic EMAS tops. Paint application for yellow chevron markers would be required. After selection of a cover layer material that meets these general requirements, a moderate to full-scale test bed should be constructed for overrun testing by a one-wheel bogy appa- ratus or an aircraft. Such testing would ensure that the cover layer does not negatively impact the mechanical deceleration performance of the bed. 9.7.3. Seam Effect Characterization The research performed to date has identified that seams between glass foam blocks should be investigated further. Seams in the material have been observed to generate pulsed loads on a rigid tire form during testing; it is not presently known how substantial such loading pulses might be on actual landing gear. One evaluation method could involve using a one-wheel bogy apparatus fitted with a full-scale pneumatic tire that is towed through the material. Direct comparisons could be made between fully glued test beds and those having un-glued seams. 9.7.4. Braking Dynamics The research performed has identified that braking in the glass foam material may require further study. The current modeling method of the APC makes simplifying assumptions regarding this phenomenon that are sufficient for a concept- level evaluation. However, to ensure accuracy of design pre- dictions, some additional tests may 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. 9.7.5. Full-Scale Testing A full-scale aircraft overrun test of the glass foam arrestor bed may or may not be critical to approve and field such a system. The mechanical nature of the system is very similar to that of the current EMAS, which leaves fewer unknown factors than for the other concepts evaluated. A one-wheel bogy test apparatus may prove sufficient to characterize the remaining unknown factors of the concept, such as the cover layer, seam effects, and braking dynamics. 9.8. Summary Of the candidate systems evaluated, the glass foam arrestor concept is most similar to the current EMAS. It uses blocks of crushable foam material that are installed on a paved surface. While the mechanical properties differ somewhat from cel- lular cement, the overall dynamic behavior proves analogous. Consequently, the glass foam concept could be considered an equivalent to the existing EMAS technology, with the caveat that bed thicknesses and material properties would not be identical. The material is closed-cell glass foam, which provides inher- ent moisture and chemical resistance and improved handling durability as compared with cellular cement. Manufacturer information regarding past use indicates that long service life is possible, potentially eliminating the standard 10-year replacement for an EMAS assumed in FAA Order 5200.9. Water immersion must be avoided, so drainage measures would still be required for the bed. Installation of the system would rival the current EMAS design in terms of time and effort required. Repair of the bed after an overrun would be analogous to the current EMAS requirements. Preparatory paving requirements would be identical. The cost to establish such a system would be nom- inally equal to, or greater than, EMAS; however, life-cycle costs could be reduced due to longer bed life. Maintenance needs could be reduced for a monolithic version of the concept, which would have no joints to seal with tape or caulk. The APC predictions for the glass foam arrestor show fairly constant deceleration throughout the arrestment with little speed dependence, which are desirable characteristics. Arrest- ing distances were comparable to the current EMAS. Of the three candidates evaluated, the multi-aircraft design case for the glass foam concept showed moderate performance. Arresting distances varied with the aircraft size. Transition to a fielded system would require selecting a covering material and finalizing a set of foam densities that could be used for different arrestor applications. Additional investigation would be advisable to determine the impact of material seams in producing pulsed landing gear loads. 86

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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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