VERTICAL TAIL FATIGUE MONITORING PROGRAMS. B. Aktepe ..... mode 1 strain with minimal disturbance from mode 2, a gauge is best placed on.
BUFFET LOADS AND STRUCTURAL REDUNDANCY CONSIDERATIONS IN VERTICAL TAIL FATIGUE MONITORING PROGRAMS B. Aktepe The Sir Lawrence Wackett Centre for Aerospace Design Technology Royal Melbourne Institute of Technology University GPO Box 2476V, Melbourne VIC 3001, Australia L. Molent, A.D. Graham & D. P. Conser Aeronautical and Maritime Research Laboratory Defence Science and Technology Organisation PO Box 4331, Melbourne VIC 3001, Australia International Aerospace Congress, Adelaide, Sept. 1999
ABSTRACT Vertical tail fatigue monitoring has traditionally not received much attention due partly to conservative design of the tail, and partly to the difficulty inherent in the task. More recently, fatigue significant buffeting of the F/A-18 empennage has led to an investigation of vertical tail fatigue monitoring methods. Although advances have been made in fatigue monitoring equipment, the purpose of this paper is to highlight the potential difficulties associated with using strain gauges for fatigue monitoring on redundant structures that experience buffet loads. The paper presents the results of preliminary studies carried out on strain gauges attached to the vertical tail fuselage attachment stubs of the F/A-18. Their response to structural dynamic modes, the variability between stub strains at different fuselage locations and optimal gauge location for loads monitoring are discussed. 1
INTRODUCTION
Aircraft fatigue monitoring has traditionally concentrated on the wing and centre fuselage section of an aircraft due to the relative simplicity in measuring centre of gravity acceleration and relating this to stresses in critical wing root locations. However, with the more sophisticated methods that have been developed for modern fighter aircraft that utilize flight parameters and strain gauges recorded on fast data acquisition systems, strain gauges can be strategically placed to monitor stresses at any location on the aircraft. However there are a number of difficulties associated with their use. Although strain gauges have been fitted to several locations on the empennage of all F/A-18 fleet aircraft for fatigue monitoring the vertical tail (VT) and horizontal stabilator, there has been difficulty in developing a reliable calibration procedure for them. This is crucial to enable fleet measured strains to be related to fatigue
test results. Simple ground based calibration of VT gauges on fleet aircraft [1], carried out to validate analytical calibration approaches [2][3], has shown variations of up to 40% from aircraft to aircraft, and one of the contributing factors for this is thought to be the highly redundant structure of the VT. Six load paths transfer manouevre and high amplitude buffet loading from the VT into the aft fuselage of the aircraft. As with other modern highly agile aircraft, the VT of the F/A-18 is exposed to a very severe buffet environment induced by vortices that are generated upstream from the wing leading edge extension (LEX)[4]. The vortex breakdown energises the resonant frequencies of the VT, creating significant multi-modal stresses at the VT root and contributing to the accumulation of fatigue damage [5]. Hence this is an area of concern to the fleet fatigue managers and the consequences of this loading complexity must be considered in developing an effective fatigue monitoring strategy for the VT. Furthermore, if static calibration identifies significant variation between aircraft, then additional calibration methods may be required to deal with dynamic loading. The problems associated with using strain gauges for fatigue monitoring of a redundant structure subjected to dynamic loading is studied here using data from the VT of the Australian F/A-18 International Follow-On Structural Test Project (IFOSTP) aft-fuselage and empennage fatigue test article FT46 [6][7]. The FT46 test article is subjected to dynamic loads replicating the flight buffet environment. A preliminary analysis of dynamic data from strain gauges attached to the VT attachment stubs from FT46 was undertaken to gain an understanding of the strain distributions present in the VT root structure under dynamic loading conditions. This paper highlights the issues associated with judicious placement of strain gauges for fatigue tracking purposes when a redundant structure is subjected to multiple loading actions, and also advances a possible fatigue monitoring alternative. Although the F/A-18 is used in this study, the issues are relevant to any modern aircraft experiencing similar flight load environments. 2
THE F/A-18 VERTICAL TAIL STRUCTURE
The structure of the two VT’s on the F/A-18 comprises a swept 8-spar main structural box that extends from the 22.5% (chord) spar to the 77.5% spar and from the lower closing rib to the upper closing rib. The root chord is 113 inches long, and the span is 96 inches. The thickness of the spar flanges and webs varies between the spars and along the spars. Each of the first six spars is connected to the rear fuselage by six attachment stubs as shown in Figure 1. The six attachment stub frames are at fuselage stations Y557, Y566, Y574, Y580, Y590 and Y598. The redundancy in the load path from the VT into the rear fuselage is evident. Although the stub
attachment frames are all of similar “I beam” configuration, with large inner and outer flanges connected by a web, each is of a slightly different dimension. The aft three stub frames are reinforced by cleats attached to their inboard flanges.
Figure 1: F/A-18 empennage structure The configuration of the aft three stub frames (minus cleats) is shown in Figure 3 3
F/A-18 EMPENNAGE CHARACTERISTICS
BUFFET
ENVIRONMENT
AND
DYNAMIC
Because the F/A-18 VT buffet is a consequence of the LEX vortices, the aircraft’s angle of attack (AoA) and dynamic pressure (Q) are the two primary drivers of the buffet characteristics. As the AoA increases, the buffet centre of pressure sweeps up the VT, changing the dynamic loading on the structure. A number of dynamic structural modes of the VT’s are excited by this unsteady pressure. At about 20° AoA a primary torsion mode occurring at 45 Hz (and termed mode 2) begins to be excited, with its amplitude peaking between 28° to 32° AoA. A primary bending mode occurring at about 15 Hz (termed mode 1) is also excited, with its amplitude peaking around 32° to 36° AoA. This environment has been well documented previously [5]. These modes are shown in Figure 2. The rudder rotation mode and a secondary bending mode occur at 48Hz and 72 Hz respectively but they have traditionally been considered insignificant in their contribution to the fatigue damage, hence are not considered in this paper. To assess the dynamic response of both VT’s in the buffet environment, accelerometers were fitted to the aft tips of both VT’s on several flight test aircraft [8]. These accelerometers were designated KS16 and KT16 for the port and
starboard VT’s respectively. Acceleration levels of ± 120g for mode 1 and ± 600g for mode 2 were measured. Mode 1
node line
Mode 2
node line
Figure 2: Mode shapes of the vertical tails This buffet environment has been duplicated in the FT46 fatigue test, in which VT loads for various AoA-Q combinations are applied to the test article to simulate real flight conditions experienced when an aircraft completes a large range of different flight manoeuvres within a typical flight. Previous flight test strain data analysis [5] has shown mode 1 is particularly dominant at the high AoA and low Q regions of flight, where the Root Mean Square (RMS) stub strain values have been measured at more than 20 times those of mode 2. At lower AoA from about 20° to 30°, mode 1 is still dominant at low Q up to about 225 psf1. It is only in the low AoA, low Q region that mode 1 and mode 2 contribute approximately equal amounts to the total strain, however the level of VT buffeting in this region is small. The analysis of FT46 test data examines this effect on an array of strain gauges, including those representing fleet fatigue tracking sensors described below. 4
F/A-18 VERTICAL TAIL FATIGUE MONITORING INSTRUMENTATION
Each F/A-18 in the Royal Australian Air Force (RAAF) fleet is fitted with two fatigue monitoring systems for potential individual aircraft tracking purposes. The Maintenance Signal Data Recording System (MSDRS) is the more complex of the two, recording both flight parameters and strains at a number of locations including two on the VT stub frames. It collects raw data from the strain gauges at a sampling rate of up to 10 Hz and from flight parameters at lower frequencies. The second system, the Australian developed Aircraft Fatigue Data Analysis System, AFDAS [6], although not currently fully operational, collects range pair data from two VT stub strain gauges. Unfortunately the monitoring systems and gauge locations were installed prior to the characterization of empennage buffet; 1
Imperial units used here to enable simple comparison with published results from earlier fatigue tests and flight trials.
they had been chosen to monitor manoeuvre loads. Hence the sampling rate of neither system was high enough to capture all the strain cycles from mode 1 and mode 2 dynamic loading. However the MSDRS and AFDAS gauges are considered in this paper to demonstrate the importance of prudent gauge location. It was impractical to strain gauge every VT stub attachment frame, therefore a primary and a backup gauge were fitted to the inner faces of the outboard flanges of the Y598 attachment stub frame only, for each VT for the MSDRS system. On RAAF aircraft, the starboard VT backup MSDRS gauge was wired into the AFDAS system and an additional AFDAS gauge was fitted to the Y566 frame. Hence a total of five VT gauges, listed in Table 1, were fitted to each aircraft. Table 1: Location of VT AFDAS and MSDRS gauges Frame System Side Gauge# Position on Number flange face Y598 MSDRS Port M33797 Forward of web Y598 MSDRS Starboard M33798 Forward of web Y598 MSDRS Port M33799* Aft of web Y598 AFDAS & MSDRS Starboard A/M33800* Aft of web Y566 AFDAS Starboard A33792 Forward of web # All gauges listed in this table are on the inner face of the outboard flange. * indicate MSDRS backup gauges; all others are primary gauges. The Y598 gauges locations are shown in Figure 3. The Y566 AFDAS gauge is in an equivalent location but forward of the web. FORWARD Z
33800
Z
1.62” 33798
FS598.0 FS590.5
0.200” 0.030”
FS580.5
View Z-Z Looking Outboard at Y598 Starboard Stub Outboard
Forward
Figure 3: Location of VT AFDAS/MSDRS gauge 33800 and MSDRS gauge 33798. (Gauges 33799 and 33797 are in mirror locations on the port stub).
5
ANALYSIS OF FT46 DYNAMIC TEST DATA
To investigate the complexity of the dynamic strain distribution in the VT attachment stubs, continuously monitored strain and acceleration data from block 5 of the FT46 test article were studied. Each test block represents a sequence of 250 flights of RAAF operation, with the dynamic loads developed to simulate the buffet in each AoA and Q flight combination appearing in the flight sequence. For this investigation the strain and acceleration test data were separated into separate AoA/Q bins. Power Spectral Density (PSD) analysis was performed on the strain data to assess the relative magnitude of each dynamic mode. A sample PSD plot for the MSDRS gauge A33800, Figure 4, shows that for the three AoA-Q bins chosen, the mode 2 cycles are negligibly small compared to the mode 1 cycles. This demonstrates that the Y598 A33800 gauge responds almost entirely to mode 1, although significant differences in the mode 1 response can be seen from the three AoA-Q data samples considered. Root-mean-square strains for all stub strain gauges and RMS acceleration at the aft fin tip (KS16 and KT16) for each AoA/Q bin were calculated for each VT mode. 50,000 45,000
28-32°, 225-300 psf
PSD [Mag2 / Hz]
40,000
39°+, 75-125 psf 24-28°+, 300-350 psf
35,000 30,000 25,000 20,000 15,000 10,000 5,000 0,000 5
10
15
20
25
30
35
40
45
50
55
60
Frequency [Hz]
Figure 4: PSD for A33800 from FT46 Block 5 data To relate stub strain to VT response, RMS cross-plots of strain versus acceleration response for each AoA/Q bin for the two fundamental frequencies were plotted for each VT AFDAS and MSDRS gauge. The plot for the A33972 gauge is shown in Figure 5. It shows that the strain per g values are constant across AoA/Q bins. This also holds true for all other VT stub gauges. This is an important result since it shows that the strain response at any particular location is directly proportional to the acceleration at the aft tip of the VT. Therefore, from a fatigue monitoring perspective, it should be possible to monitor
strain and hence fatigue accrual by monitoring acceleration directly on a fleet aircraft. This result also has the potential to be used in cases where the strain gauge becomes unserviceable but its data can be filled-in using accelerometer data.
A33972 RMS strain [microstrain]
250 200 A33972 = 4.2244 * KT16 2 (R = 0.9944)
150
A33972 = 1.8661 * KT16 2 (R = 0.9964)
100 Mode 1
50
Mode 2 Linear line of best fit (mode 2) Linear line of best fit (mode 1)
0 0
20
40
60
80
100
120
140
KT16 RMS Acceleration [g]
Figure 5: Sample RMS cross-plot for A33972 from FT46 Block 5 Strain per g values from cross plots for four of the MSDRS and AFDAS gauges are shown in Table 2. Table 2: Mode 1 and 2 strain per g responses FS
Gauge
Y598 Y598 Y598 Y566
M33797 M33798 M/A33800 A33972
Mode 1 strain per g (R2) 10.786 (0.997) 9.595 (0.997) 14.576 (0.997) 4.224 (0.995)
Mode 2 Strain per g (R2) 1.249 (0.996) 1.110 (0.995) 1.163 (0.988) 1.866 (0.997)
The results show that: •
for Y598 gauges, the strain per g values are nearly an order of magnitude higher for mode 1 than for mode 2;
•
there is substantial variation in the response between gauges on the same Y598 VT stub (50% between M33798 and M33800) for mode 1, but smaller variation (8%) for mode 2.
5.1
Strain distribution across frames and across stubs
To investigate the strain distribution on each stub and across all twelve stubs, FT46 Dynamic Strain Survey (DSS) data prior to block 6 sampled at 606.06Hz was analyzed. Dynamic loads for various combinations of AoA and Q were
applied to FT46 during the DSS and numerous passes were made to allow all strain gauges to be collected. The same load sequence was applied during each pass; hence the RMS accelerations for each mode were constant across passes. Data from 28 gauges on the stub pairs at each frame station, covering both the inboard and outboard faces of each stub flange, as shown in Figure 6 and Figure 7, were analyzed. The gauges were located in nominally the same relative position on each stub, subject to gauge placement accuracy, stub configuration differences and gauge similarity. PORT
13 & 17
5 & 21
FORWARD
1
33
15 & 19
7 & 23
3
9
10
31
32
11
12
STARBOARD
2
6 & 22
14 & 18
34
4
8 & 24
16 & 20
CL Figure 6: A simplified plan view of FT-46 strain gauges for each of the six frames. Note that gauges 5 & 13 are opposite to one another, as too are 17 & 21, 6 & 14, etc. The complete VT gauge numbering is given by a five digit number, where the first three digits indicate the frame no (15=Y598, 20=Y557) and the last two the location on that frame. The approximate location of the MSDRS and AFDAS gauges in relation to the other FT-46 gauges is listed in Table 3. There are no gauges on the inboard face of the inboard flange for the aft 3 stubs, due to cleats at those locations. For mode 1, plots of strain at each fuselage station, Figure 8, compare the strain distribution across all 12 stubs for nominally the same gauge locations. Considerable difference exists in the strain per g value between gauges on opposing flanges of the same stub. Gauges on the outboard and inboard flanges in nominally equivalent locations showed different trends from the foremost stub to the rearmost stub. Reasonable agreement between port and starboard side gauges is seen and it was also found that gauges on the inboard faces of the outboard flange followed a similar trend to their mirror gauges on the outboard face of the same flange. Gauge 1 at the Y590 stub appears anomalous.
Table 3: Location of AFDAS and MSDRS gauges in relation to other FT-46 gauges Gauge Approximate Location M33797 Between gauges 15005 and 150021 M33798 Between gauges 15006 and 150022 M33799 Between gauges 15007 and 150023 A/M33800 Between gauges 15008 and 150024 A33972 Between gauges 19006 and 190022 Note: All VT AFDAS and MSDRS gauges are on the inboard face of the outboard flanges
Figure 7: Gauges on outboard face of outboard flange of the Y598 port side stub. Note the last two numbers (ie xxo17) identify locations common to each stub frame. For mode 2 (Figure 9), it was again seen that the strain per g distribution between port and starboard stubs is reasonably symmetrical. The inboard flanges were observed to experience a different strain distribution to the outboard flanges. Once again significant variation between the stub frames is noted. The maximum mode 1 strain per g was experienced by Y598 outboard flange gauges, while the maximum mode 2 strain per g was experienced by the Y580 outboard flange gauges. Data from gauge 1 on the Y590 frame (i.e. gauge 16001) appears anomalous. The trends for the forward inboard gauges 1 and 2 are very similar for mode 1 and mode 2, suggesting that the mode 1/ mode 2 ratio for this location is similar at all fuselage stations.
Mode 1 Strain per g
25
cleated stubs
20
13
14
15
16
17
18
33
34
15
10
FWD
5
13 17 33
14 18 34
15
16
0 Y557
Y566
Y574
Y580
Y590
Y598
Fuselage Station
(a)
cleated stubs
14
Mode 1 Strain per g
12 10 8 6
1
2
3
4
4 FWD
1
2
3
4
2 0 Y557
Y566
Y574
Y580
Y590
Y598
Fuselage Station
(b) Figure 8: Pre-block 6 DSS mode 1 strain per g for gauges on outboard faces of (a) outboard flanges and (b) inboard flanges
cleated stubs
2.6
Mode 2 Strain per g
2.4 2.2 2.0 1.8 1.6 1.4
13
14
15
16
17
18
33
34
FWD
13 17 33 15 19
14 18 34 16 20
1.2 Y557
Y566
Y574
Y580
Y590
Y598
Fuselage Station
(a)
2.0
cleated stubs
Mode 2 Strain per g
1.8 1.6 1.4 1.2 1.0 0.8 0.6 0.4 0.2
1
2
3
4
FWD
1
2
3
4
0.0 Y557
Y566
Y574
Y580
Y590
Y598
Fuselage Station
(b) Figure 9: Pre-block 6 DSS mode 2 strain per g for gauges on outboard faces of (a) outboard flanges and (b) inboard flanges
In order to determine if a location could be found to respond predominantly to a single mode, the ratio of mode 1 RMS strain to mode 2 RMS strain, ρ, was determined from the DSS data, since this data was readily available from the previous analysis. It should be noted, as shown in Figure 5, that strain per g is a quantity independent of AoA and Q. Thus the ratio of mode 1 strain per g to mode 2 strain per g, υ, would be invariant with test data used for calculating the ratio. However ρ is directly proportional to υ. A plot of ρ versus FT46 gauge number using Block 6 DSS data for each of the 6 frames, Figure 10, shows that for almost every gauge, the RMS strain for mode 1 was higher than that for mode 2, i.e. ρ greater than 1. It was also found that the frame with consistently high ρ is Y598. Therefore, in order to capture high mode 1 strain with minimal disturbance from mode 2, a gauge is best placed on the outboard flange of the Y598 frame stub. The MSDRS gauges on the Y598 are therefore reasonably well placed (but not optimal) on the outboard flange of the Y598 stub to monitor mode 1 loading. Y580
Y566
Y590
Y574
Y598 M33797 M33799
RMS strain ratio
8.0
Y557
A/M33800
9.6
6.4
4.8 A33972 3.2
1.6 1.0 Points joined for illustraion only 0.0 0
2
4
6
8
10
12
14
16
18
20
22
24
26
28
30
32
34
36
Gauge number
Figure 10: Ratio of Mode 1 RMS strain to Mode 2 RMS strain for FT46 VT stub frame gauges The ratios from gauges 3 and 4 appear anomalous, although there is a similar but lesser trend on the Y590 stub.
For fatigue monitoring purposes, in addition to assessing ρ, to determine the dominance of the mode of interest, the absolute magnitude of the response at the location chosen should also be considered. It should be noted that the maximum VT tip acceleration response for mode 2 is considerably higher than for mode 1, so although the strain per g values for mode 1 are higher than mode 2, the actual strain may still be quite significant. Most of the gauges positioned on each of the stubs were seen to respond predominantly to mode 1, but with some influence from mode 2. While the influence of mode 2 is significant for a large number of gauges, no gauge could be said to respond predominantly to mode 2. This is probably due to the VT root region not being influenced significantly by mode 2 or it may be due to the sensing axis of all the gauges being aligned for VT bending and not for VT torsion. The most appropriate location to place a strain gauge to monitor both mode 1 and mode 2 appears to be the forward outboard flange of the Y580 stub, gauge locations 13 or 14, where both modes are reasonably significant. The forward stub frame at Y557 appears to be the best location for monitoring mode 2 with the minimum influence from mode 1. A fatigue monitoring system must be capable of capturing all significant strain cycles, so in this example, it must have a sampling frequency of at least 500Hz to monitor mode 2. In this case, mode 1 data will also be captured, so it is possible to monitor all loading by only one strain gauge located to respond to both modes. A detailed dynamic strain survey or flight test program would be necessary to relate this single gauge response to strain at other locations on the structure. Strain gauge calibration then becomes a significant issue. Slight differences in dynamic structural response between VTs have the potential to make strain comparison between aircraft and reference fatigue test article very difficult. Although mode 2 does not appear to be as significant at the VT attachment stubs, it may to some extent be due to gauge orientation. Where multiple modes are present it is recommended that strain rosettes be installed to identify the optimal orientation of gauges for all dynamic modes. 6
NON-STRAIN GAUGE BASED OPTIONS FOR VERTICAL TAIL FATIGUE MONITORING
Non-strain gauge methods, using unit damage matrices, to quantify the fatigue effects of buffeting using MSDRS data have been reviewed in [9] and [10]. In another VT fatigue monitoring approach, the F/A-18 Structural Appraisal of Fatigue Effects (SAFE) software in the latest version, V200, collects usage time spent in AoA and Q regimes, then generates dynamic usage spectra for each vibration mode and combines these with a fixed manoeuvre spectrum. A crack initiation model is then used to calculate the fatigue life expended and severity
ratio for each of the two dynamic modes [8]. The premise in SAFE V200 is that certain regions of the VT are affected by particular dynamic modes. However a review of the SAFE V200 has identified several anomalies in the program. A comparison of SAFE V200 results with simply the time an aircraft spends above 10º AoA to predict VT damage suggests that this parametric approach may be just as valid. A parametric approach similar to the methodology used by SAFE V200 may in future be suitable for routine fatigue monitoring. However the variability in acceleration response identified in IFOSTP flight-testing suggests that the VT buffet loading mechanism is not sufficiently understood at this stage for an accurate parametric method to be developed for fatigue tracking. Alternatively, any comprehensive VT tail fatigue monitoring system should include the monitoring of accelerometers, since the tip accelerations of the VT provide a good indication of the level of buffet strain induced in the structure and hence the amount of damage accrued. A detailed dynamic strain survey or flight test program would also be required to relate fleet acceleration to multiple locations on the structure. The use of accelerometers is an option that should be studied for future fatigue monitoring. 7
CONCLUSIONS
A preliminary analysis of data from strain gauges fitted to the VT attachment structure on the F/A-18 has been conducted to identify some of the complexities involved with using strain gauges for fatigue monitoring purposes. The study has found that comprehensive strain data are required to determine the optimum location at which to place a gauge in order for it to respond either purely to one mode or all modes. The variability between port and starboard gauges was small. Significant variability exists in the response between gauges on the same flange and on opposing flanges of the same stub, so gauge placement for fatigue tracking purposes is not intuitive. It was shown that the rear most frame Y598 was the most suitable location for detection of primary bending mode (mode 1) strain. The primary torsion mode (mode 2) was harder to detect on its own without significant influence from mode 1, but the frame with the lowest mode 1 to mode 2 strain ratio was the Y557 frame. The existing MSDRS and AFDAS gauges were found to be at locations which responded reasonably well to mode 1. However they are not optimally placed to pick up both modes, and could be better placed if only mode 1 response was required.
A linear relationship was found between stub strains and the VT aft tip acceleration across a range of AoA and Q for almost every stub frame gauge. That is, strain per g is independent of AoA and Q. It is recommended that aircraft buffet acceleration data be considered for development of future fatigue tracking of aircraft structure subjected to a buffet environment. This preliminary data analysis highlights the difficulty of identifying locations on a redundant load path structure for monitoring all loading actions, and appears to raise more questions than answers. With the availability of FT46 dynamic full scale fatigue test and flight test data collected for test development, there is a great opportunity to investigate some of these issue to assist in developing more robust fatigue tracking methodologies for the future. 8
ACKNOWLEDGEMENTS
The authors are thankful to the support provided by the AMRL IFOSTP and F/A18 Life Assessment Groups. This work was undertaken under the RAAF sponsored AIR 98/092 task. 9
REFERENCES
[1]
Aktepe, B., Hewitt, K., Ogden, R., & Molent, L., Ground Calibration of RAAF F/A-18 Onboard Fatigue Strain Sensors, Defence Science and Technology Organisation, DSTO-TR-0641, Melbourne, TBP.
[2]
Molent, L., Aktepe, B. & Polanco, F., F/A-18 Airframe Fatigue Data Analysis System Validation, Defence Science and Technology Organisation, DSTO-TR-54, Melbourne, 1994.
[3]
Ooi, Y.G., Ogden, R.W. & Molent, L., Validation of an analytical Technique for Scaling of F/A-18 Vertical Stabilator Strain Sensors, Defence Science and Technology Organisation, DSTO-TR-0539, Melbourne, 1997.
[4]
Lee, B.H.K. & Brown, D., Wind Tunnel Studies of F/A-18 Tail Buffet, Proceedings of the 16th Aerodynamic Ground Testing Conference, Seattle, WA, June 18-20, 1990, AIAA Paper 90-1432, 1990.
[5]
Graham, A.D, Madley, W. B., Rider, C. K. & Waldman, W., Fatigue Analysis and Testing of Aircraft Subjected to Manoeuvre and Buffet Loads of Comparable Magnitude, Proceedings of the 24th International Committee on Aeronautical Fatigue Conference, ICAF’95, Melbourne, Australia, May 1-2, 1995.
[6]
Graham, A.D., Symons, D., Sherman, D. & Eames, T., ARL F/A-18 IFOSTP Full Scale Fatigue Test, 5th Australian Aeronautical Conference, Melbourne, 13-15 Sep 1993.
[7]
Conser, D. P., Graham, A. D., Smith, C. J. & Yule, C. L., The Application of Dynamic Loads to a Full Scale F/A-18 Fatigue Test Article, Proceedings of the Congress of the 20th International Council of the Aeronautical Sciences, Sorrento, Italy, 1996.
[8] Conser, D.P. & Keys, G.L., F/A-18 Production ASPJ Vertical Tail Dynamic Fatigue Test FT98 Test Spectra Development, McDonnell Douglas Corporation, MDC 91B0424, St. Louis, 1992. [9]
Aktepe, B., Monitoring Vertical Tail Fatigue on the F/A-18 Using the Mission Severity Monitoring Program, Defence Science and Technology Organisation, Melbourne, TBP.
[10] Chapman P., Hewitt, K.P. & Molent, L., A Critical Review of a F/A-18 Vertical Tail Fatigue Tracking Process, Defence Science and Technology Organisation, DSTO-TR-0642, Melbourne, TBP.
