Servo-hydraulic System for Controlled Velocity WaterImpact of Marine Sandwich Panels

M. Battley & T. Allen

Received: 10 April 2011 /Accepted: 29 August 2011 /Published online: 6 October 2011# Society for Experimental Mechanics 2011

Abstract Slamming, the impact between a marine craft’shull and the water surface is a critical load case forstructural design of marine vessels. The importance of hullslamming has led to a significant body of work tounderstand, predict and model these impacts. There ishowever, a lack of experimental data for validation,particularly for deformable panels and sandwich structures.This paper describes a high-velocity panel slamming testsystem that enables the generation of comprehensive andreliable experimental data on slamming impacts for bothrigid and flexible panel structures. The pressure magni-tudes, time-histories and spatial distributions resulting fromtesting of a nominally rigid panel have been compared withprevious analytical, semi-empirical and experimental studies.Slamming impacts of a deformable sandwich panel areshown to cause different pressures to those from a rigidpanel impact, resulting in increased transverse shear loading atthe panel edge.

Keywords Hull slamming .Marine craft . Sandwichcomposites . Dynamic testing . Hydrodynamic loading .

Failure

Introduction

Slamming, the impact between a marine vessel’s hull andthe water surface is one of the critical load cases consideredfor hull structural requirements. These impacts can result in

highly dynamic loads, where the load profile is a peakpressure followed by a lower residual pressure. Thispressure distribution travels across the panel width duringthe immersion. The pressure pulse magnitude and propa-gation speed are primarily dependent on the impact velocityand deadrise angle of the impacting body. A schematiccross-section of a hull undergoing slamming is given inFig. 1. The circled area indicates the area of interest whichis represented by the half-wedge specimens investigated inthis study.

The significance of hull slamming in the design of marinecraft has led to a significant body of work trying to understand,predict and model these impacts. The research has progressedfrom idealised analytical theories, to experimental drop tests,to numerical modelling of the situation.

By considering the conservation of momentum, ananalytical expression for the pressure distribution appliedduring a slamming event can be obtained.Wagner [1] and vonKarman [2] were two of the early researchers to use thisapproach. Since their initial developments several moreanalytical expressions for both peak pressure and pressuredistribution have been developed by subsequent researchers[3–5]. Payne [4] produced a critical review of many of themodifications made to von Karman’s [2] initial derivations.His key criticism was of Wagner’s splash up factor of π/2,which he suggests there is little evidence for, and that itresults in higher pressure than experimentation and alterna-tive analytical solutions. Payne also highlighted the numberof subsequent researchers that have continued to use thisfactor, needing additional modifications to achieve goodagreement with experimental results. A set of equations wasproposed by Payne to determine both maximum pressure andthe applied force. The differences in the various analyticalsolutions derived demonstrate the complexity of the physicalphenomena involved in slamming.

M. Battley (*) : T. AllenCentre for Advanced Composite Materials, Departmentof Mechanical Engineering, The University of Auckland,Auckland, New Zealande-mail: m.battley@auckland.ac.nz

Experimental Mechanics (2012) 52:95–106DOI 10.1007/s11340-011-9543-7

Experimental work has been undertaken to attempt tocharacterise the impact of structures into water and validateanalytical theories and numerical models. The majority ofthe experimental work has been based on drop tests ofnominally rigid structures [6–9]. Chuang [6] produced acomprehensive experimental characterisation of the pres-sures applied to rigid bodies undergoing slamming. Droptests were undertaken on 500×675 mm specimens, manu-factured from half inch steel plate with 12.7 by 76 mm steelflat-bar stiffeners, at 1, 3, 6, 10 and 15° deadrise anglesdropped from heights of 75, 115, 150 and 190 mm. Thefluid domain was 7.6 by 4.5 m and 2.5 m deep for the testsand the flow was constrained by two vertical steel platewalls extending the depth of the tank. The plates wereseparated by a distance equal to the length of the modelsplus a small amount of clearance. Chuang’s experiments [6]also investigated the effect of entrapped air in wedgeimpacts at low deadrise angles. During the impacts theelectrical resistance was measured across the surface of thewedge to determine if water or air was in contact with thesurface. He also used underwater high speed cameras(5000fps) to film the air entrapment and concluded that inimpacts with more than 3° of deadrise most of the airescapes at the instant of impact.

Stavovy and Chuang [10] utilised the data collected byChuang [6–8] to derive a semi-empirical formula for themaximum pressure during impact. This semi-empiricalformula was derived by curving fitting of a three-dimensional analysis combined with a comparison of theinitial wedge and cone impact results. The resulting formulahad good agreement with the experimental work on scalemodels undertaken by Stavovy and Chuang [6, 10]. Paynealso compared the solution he had developed for maximumpressure [4] with Chuang’s experimental results [8], showinggood correlation. Both solutions had much better correlationwith the experimental data than Wagner’s solution for peakpressure [1], especially at lower deadrise angles.

The major limitation of drop tests is the deceleration thestructure experiences during the impact. Research under-taken using constant velocity impact systems has beenlimited to velocities below 1.5 m/s [25]. Only a very limitedamount of experimental work has focused on flexiblestructures [11–15]. The major difference between flexibleand rigid structures is the presence of hydroelasticity, which

is the coupled fluid structure interaction between theflexible structure and the water. As the structure deformsduring the impact the fluid boundary conditions changeresulting in different applied pressures which in turn affectsthe deformation.

Numerical modelling of slamming has been undertakenin several other studies [16–18]. Limitations in the prevail-ing methods have been discussed further by Rosén et al[19]. The current focus of many researchers is on thedevelopment of more accurate numerical methods formodelling slamming impacts of wedge impacts with fewersimplifications and the inclusion of hydroelasticity. In orderto validate the coupled models, accurate and comprehensiveexperimental data is required.

This paper describes a servo-hydraulic testing systemthat has been developed for the characterisation of theresponse of hull panels to slamming impacts. The aim ofthe system is to enable the generation of comprehensiveand reliable experimental data on slamming impacts forboth rigid and flexible panel structures for characterisationof slamming impacts as well as validation of numericalmodels. The configuration of the system hardware, softwareand instrumentation is described, and pressures resultingfrom water impact of a nominally rigid sandwich panelpresented and compared to previous research. The pressuresand structural responses are presented and discussed for adeformable foam-core sandwich panel.

Experimental Procedure

The testing facility used in this study is known as theServo-hydraulic Slam Testing System (SSTS). This wascustom designed and developed by Industrial ResearchLimited, Auckland, NZ and the University of Auckland andutilises a sophisticated high-speed servo-hydraulic systemto control the motion of a panel structure during impactwith the water. It can be programmed to have a constantvelocity throughout the slam event, or to follow a changingvelocity profile such as may be measured on a real vessel orpredicted by motion prediction software.

Battley and Stenius [18] conducted a series of constantvelocity slamming experiments on a single curvaturesandwich core panel using the SSTS. A modal transientFinite Element Analysis (FEA) simulation was comparedwith the experiments. The FEA simulation used a simpli-fied pressure distribution with the maximum pressurecalculated from Stavovy and Chuang’s semi-empirical peakpressure formula [10], and the residual pressure based onscaled Payne’s mean pressure calculations [4]. The rise andfall times were the only variables obtained from theexperimental data. The results showed good correlationbetween experimental and FEA solutions for all values

vDeadrise Angle, β

Fig. 1 Cross-section of marine vessel undergoing hull slamming

96 Exp Mech (2012) 52:95–106

except strain, which the authors attribute to rigid bodyinertial motions that the FEA solution did not account for.

Breder [20] conducted a methodical set of experimentscharacterising the pressures exerted on a rigid half wedgeimpacting at constant velocity using the SSTS. His workdid not include an evaluation of the system’s performanceor provide comparisons to analytical solutions and existingexperimental results, as given in this paper.

For determination of mechanical properties of corematerials used in slamming applications Det Norske Veritas(DNV) [21] recommends testing using standard four pointbeam tests at high stress rates. Downs-Honey, Edinger andBattley [22] used the SSTS to investigate how wellstandardised mechanical property tests, such as 4-pointbending and plate shear tests (ASTM C273 and ASTMC393), predicted the properties that are exhibited duringslamming. Their results showed that some cores, such asNomexTM honeycomb and balsa-wood performed asexpected based on static material properties. However bothof the PVC foam cores significantly exceeded the predictedperformance which the authors believed was due toincreased energy absorption at higher strain rates.

Following on from Downs-Honey et al. [22], Battley andLake [23] experimentally investigated how the response ofa sandwich panel subjected to slamming loads varies fromthe same panel configuration subjected to a uniform staticload, as is often used in design specifications and scantlingcodes. Their work concluded that slamming induced muchhigher shear loading and deflections than predicted by staticpanel theory. Further research by Battley and Allen [24–28]extended this study to additional material systems with afocus on flexible panels, and the experimental methodology.

System Overview

Figure 2 shows the overall testing system which consists ofa polyethylene tank measuring 3.5 m in diameter filled withwater to 1.4 m deep, giving a fluid domain of 13,500 L. Aservo-hydraulic system moves the specimen and supportfixture vertically into the water surface during the impacts.The specimen fixture (Fig. 3) runs on linear bearingsbetween flow restriction panels and is supported by analuminium frame as shown in Figs. 4 and 5. There are threeflow restriction panels, on either side of the specimen andone at the rear. The rear panel is to create a symmetryboundary allowing for comparisons with full wedges or hullforms. The control and data acquisition systems are situatedoutside of the tank, while the hydraulic ram system islocated above the tank. The system is designed to maintaina constant impact velocity during the course of theslamming events. During the impacts the applied load andcorresponding panel responses are measured and recordedusing a variety of load cells, strain gauges, displacement

transducers and pressure transducers with a high speed dataacquisition system.

Servo-hydraulic Impacting System

The specimen fixture, along with the attached panel andinstrumentation, is forced vertically downwards using a1.4 m travel hydraulic ram, supplied by two pre-charged(140kPa) 20 L accumulators. The rate of descent iscontrolled using a servo valve (Model: D664-3033 K byMOOG GmgH) with feedback from a linear ram displace-ment transducer and an accelerometer. Rigid steel piping isused between the manifold and the ram to maximise theflow rates and the accuracy of the velocity control.

Fig. 2 Servo-hydraulic slam testing system

Fig. 3 Specimen fixture with specimen in place

Exp Mech (2012) 52:95–106 97

Control System

The closed loop control system is based on a MOOG™MSC 3000 electronic controller and custom developedsoftware within the Moog™ Axis Control System (MACS)programming environment. The control parameters can bedefined in different ways depending on the complexity ofthe desired motion. In the simplest case, for a constantvelocity the primary inputs are initial acceleration from rest,target velocity for the constant speed stage, then a de-acceleration rate at the end of the test. Additional inputparameters define the duration of each phase of motion and

filtering settings for the feedback signals. For morecomplex velocity profiles the command signal can bedefined by a table of time-displacement inputs. This enablesan arbitrary motion to be defined, such as increasing ordecreasing velocity during the test, or the input of a velocityprofile measured from an actual vessel in sea-trials.

At higher velocities and low deadrise angles theslamming impact can occur in a shorter time than theservovalve and control system are able to adequatelyrespond. To improve the performance of the system forthese situations the control system enables a feed-forwardcommand to be defined, input as a table of time versusadditional valve opening commands. This has the effect ofstarting to open the servovalve just before the slammingimpact occurs, so that the required valve opening and henceapplied force can be achieved within the duration of theslamming event. The correct settings for the feed-forwardcommand are determined by iterative testing.

Specimens and Fixtures

The system has undergone several redesigns of thespecimen mounting fixture. The current fixture (Fig. 3) ismade primarily of carbon fibre skin PVC foam corecomposite sandwich and can be adjusted to deadrise anglesof 0, 10, 20, 30 and 40°. Specimens have typical externaldimensions of 550±50 mm by 1025±5 mm, with anunsupported region between simply supported edges of500x1000 mm. The panel specimen is held against the edgesupports by fabric straps at each end of the panel as shownin Fig. 3. The panel outer edge is constrained from inplanemovement by a metallic fitting at each of the strappositions. Other size and shape specimens can be fittedwith appropriate modifications to the specimens, such asplywood supports bolted directly to the main specimen forthe curved panel shown in Fig. 6.

1

2

3

4

Fig. 5 Schematic of components inside the SSTS: 1-SpecimenFixture, 2-Linear bearings and tracks, 3-Aluminium support frame,4-Flow restriction panels

1

2 34

5 6

3

3

Fig. 4 Plan of the SSTS: 1-Tank, 2-Support columns, 3-Flow restrictionpanels, 4-Specimen fixture, 5-Hydraulic ram, 6-Main support beam

Fig. 6 A curved specimen in the SSTS fixture with plywoodsupports. The 5 bolt holes fix the plywood directly to the side ofmain specimen fixture

98 Exp Mech (2012) 52:95–106

Instrumentation

A modular National Instruments™ compact data acquisi-tion system (NI cDAQ 9178) and National InstrumentsLabVIEW software is used to record data from a set oftransducers during each test. The parameters measured aretypically total applied load, ram displacement, ram accel-eration, panel strains, panel displacements and waterpressure. Specifications of the transducers used are givenin Table 1. Typically data is sampled at rates of 10–30 kHz,dependant on the impact speed.

Results and Discussion

System Performance

Figure 7 shows the performance of the system during a3 m/s, 10° impact. The target acceleration and decelerationis 320 m/s2 and 23 m/s2 respectively. A detailed descriptionof the event by time is given in Table 2. The key period ofinterest is during the impact with the free surface of thewater, denoted on the graph (Fig. 7) as being between theplotted water level and panel submerged lines.

Figure 8 shows the correlation between the targetvelocity and the measured average velocity as well as themaximum and minimum velocities during the period of

impact for impacts with 10° deadrise angle. The increasedload at lower angles means the variation is the mostsignificant at 10°. Impacts with larger deadrise angles haveless variation between target and average velocities. Thedifference between the target and average velocities for 10°impacts remains less than 10% for speeds up to 6 m/s. Themaximum difference is 18.5% for the impact at 7.5 m/s.

Resulting Pressures

Flexible and rigid sandwich composite panels have beenconstructed and tested in the SSTS to establish the pressureloads applied during impacts. The key consideration for therigid panel selection was ensuring that the response of thepanel was small enough to have negligible influence on thefluid pressure. The onset of hydroelasticity has been linkedto the ratio between the first natural period of the panel andthe loading period [11]. The ratio, R, between the firstnatural frequency of a panel and the loading period of aslamming impact is presented in equation (1) where β=thedeadrise angle, D=panel flexural rigidity, V=vertical

-10

0

10

20

30

40

-1000

0

1000

2000

3000

4000

0 100 200 300 400 500

Lo

ad [k

N]

Vel

oci

ty [

mm

/s],

Po

sitio

n [

mm

]

Time [ms]

Position [mm] Load [kN] water levelpanel submerged Velocity [mm/s]

Fig. 7 System performance during a 3 m/s, 10° impact

Table 1 Transducer specifications

Parameter Manufacturer Model Signal Conditioning

Total applied load Precision Transducers LPC10t NI 9237 cDAQ Module

Ram displacement Vishay 139 L PCB-483B03 24v DC Supply

Ram acceleration PCB Piezotronics Series 308B PCB-483B03 24v DC Supply

Panel strains Measurements Group Various, typically 6 mm gauge length NI 9237 cDAQ Module

Panel displacements Schaevitz LCIT-1000 DC Supply ±12vLCIT-2000

Fluid pressure PCB Piezotronics W113A26/003C20 NI 9234 cDAQ Module

Table 2 Summary of the motion of the 3 m/s, 10° impact shown inFig. 7

Time Description

0–70 Pre-trigger recording before the start of the testwhile the rig and specimen are at the start of itsrun and not moving

70–140 Acceleration of the fixture and specimen up tothe target impact velocity

140–216 Travelling at approximately constant velocitythrough air prior to impact with the free surfaceof the water

216–251 Period of impact, from the point at which thelower edge of the panel touches the water tothe point at which the entire panel is below theinitial free surface level

251–380 Panel is completely submerged and travelling atconstant velocity through the water

380–480 The rig and specimen are decelerated andbrought to rest ending the test

Exp Mech (2012) 52:95–106 99

impact velocity, a=width of panel, b=length of panel andρw=density of water. The equation is based on the formulafor free vibration of a panel by Vinson [29], modified toconsider the additional added mass of the water, asformulated for a beam by Stenius [16]. It has been proposedthat hydroelasticity will occur when R<1 [11, 20].

R ¼ 4 tan bV

ffiffiffiffiffiffiffiffiffi

Db

prw

s

1

a2þ 1

b2

� �

ð1Þ

The nominally rigid panel was constructed using a vacuumassisted resin transfer moulding (VARTM) process with thematerial specifications shown in Table 3. Based on thepanel’s construction and the ratio presented, hydroelasticitywould only be expected above 33 m/s for 10° impactswhich is beyond the testing capabilities of the system in itscurrent form. Alternatively it has been suggested by Breder[20] that the deflection experienced by a panel can be usedas a criteria to identify the likelihood of hydroelasticeffects. The maximum deflection for the rigid panel withimpact conditions of 6 m/s and 10° was estimated to be3 mm. The reduction in local velocity based on thisdeflection was calculated to be 0.5 m/s resulting in anapproximate peak pressure reduction of 16%.

Figure 9 shows the correlation between the peak pressureresults obtained from the SSTS using the nominally rigidpanel and the expected peak pressure based on Stavovy &Chuang’s semi-empirical method [10] and Wagner’s analyticaltheory [1]. The SSTS experimental data is higher than thatpredicted using Stavovy and Chuang’s method [10] and lowerthan that predicted by Wagner’s theory [1]. The higherpressures than Stavovy and Chuang may be related toimproved measurement and acquisition devices used in currentwork, and also because the velocity in Chuang’s originalexperiments was measured at the point of initial impact [7]and is hence likely to have reduced by the time that thepressure pulse reached the transducer. The variation betweenthe experimental data and Wagner’s solution [1] can in part beattributed to the simplifications made in the solution, asdiscussed by Payne [4] who states that Wagner used anincorrect wetted width of πy for a wedge, where 2y is thewetted width assuming no pile-up. He suggests that this givesresults which are (π/2)2 too large. The predicted peakpressures shown in Fig. 9 are different by a factor closer toπ/2 with an average difference of 37% to the experimentalresults. Figure 9 also shows that the peak pressures measuredexperimentally are higher than those predicted by Payne’sown formulation [4].

0

2

4

6

8

10

12

0 2 4 6 8 10

Ach

ieve

d V

elo

city

[m

/s]

Target Velocity [m/s]

Maximum velocity Average velocity Minimum velocity

Fig. 8 Achieved impact velocity during slam event compared totarget velocity

Table 3 Rigid panel specifications

Materials

Lamina Material Carbon Fibre / Epoxy

Lamina Orientation 0/903,±45, 0/902Core Material Balsa, Baltek C100

Dimensions

Skin Thickness, tf 3 mm

Core Thickness, tc 50 mm

Panel Dimensions 1030×605 mm

Calculated Properties

Flexural Rigidity, D 281,000 Nm

Shear Stiffness, S 8,900,000 N/m

0

200

400

600

800

1000

1200

1400

1600

1800

2000

0 1 2 3 4 5 6 7

Pre

ssur

e [k

Pa]

Velocity [m/s]

Stavovy & Chuang

Wagner

Payne

Experimental Data

Fig. 9 Comparison of experimental and analytical results for 10 degrigid panel impacts

0

50

100

150

200

250

300

350

400

0 5 10 15 20 25

Pre

ssu

re [

kPa]

Time [ms]

Wagner Experimental Data

Fig. 10 Experimental pressure time histories for 3 m/s 10° impact ofrigid panel compared with Wagner’s solution

100 Exp Mech (2012) 52:95–106

While peak pressures are a commonly used parameter tocompare different impact situations, they are actually verydifficult to measure experimentally and there is limitedagreement between theoretical predictions. For experimentallyobtained peak pressure values a finite number of data points iscollected by the data acquisition systemwhichmay not capturethe actual peak value. The maximum reduction of the peakpressure compared with Wagner’s [1] predicted pressuredistribution is estimated to be 4.3% when sampling at20 kHz. The peak pressure can also be affected by the finitesensor area, which averages the actual pressure. An analysisof the influence of this effect on the peak pressure has shownthat for the 5.53 mm diameter transducer used in this study adecrease of 10.5% can be expected for all velocitiescompared with Wagner’s [1] predicted pressure distribution.

The time history of the pressures may also becompared. Figure 10 shows a comparison betweenWagner’s theory and the experimental data obtained atthe centre of the rigid panel. There is a significantdifference in magnitude for both the peak and residualpressures. Figure 11 compares the experimental data withWagner’s theory. For this comparison the theoretical peak

pressure has been converted to a single point and the risein pressure has been taken as linear from zero pressure tothe peak pressure. The magnitude has been scaled linearlywith the scaling factor determined by minimising the sumof the residual square differences between the theory andexperimental data. This resulted in a scaling factor of 0.53and coefficient of determination of 0.937. The comparisonin Fig. 11 shows that the SSTS produces a similarlyshaped pressure time history to that predicted by Wagner’sanalytical solution.

Spatial distribution of pressures for a rigid panel

Characterisation of the spatial distribution has been under-taken on a rigid panel [20] to determine if the flow issufficiently constrained by the flow restriction panels.Ideally the pressure would be equal at all locations alongthe length of the panel, only varying across the panel widthin the direction of the flow. In order to characterise thespatial distribution the pressures along the length and widthof the panel have been compared, with pressure transducerspositioned as shown in Fig. 12.

0

50

100

150

200

250

0 5 10 15 20 25

Pre

ssu

re [

kPa]

Time [ms]

Experimental

Least Squares Wagner

Fig. 11 Experimental pressure time histories for 3 m/s 10° impact ofrigid panel compared with Linear scaled Wagner’s solution

P1

P2

P3

P4

P5

P6

P12

P10P9

P8

P11

P7 P13

x

keel

chine

1030

605

a

b

b

b

b

c

dd/2d/2 2d

Fig. 12 Schematic of the panelshowing positions of pressuremeasurement points, a=60,b=107.5, c=120 andd=200 mm

-200

0

200

400

600

800

1000

1200

0 5 10 15

Pre

ssur

e [k

Pa]

Time [ms]

P1P12P3P6P5P10

Fig. 13 Comparison between centreline pressures and P6, P10 & P12

Exp Mech (2012) 52:95–106 101

Figure 13 compares the centreline pressures and thepressures measured 200 mm closer to the panel edge whileFig. 14 compares the centreline pressures with measure-ments taken 400 mm away. There is good correlationbetween the timing of the pressures in both Figs. 13 & 14confirming the spatial uniformity of the flow front, apartfrom the timing between P5 and P9. The timing differencebetween the peaks for P5 and P9 is 0.45 ms. The shapes ofthe pressure pulses are also similar in both comparisonswith comparable residual pressures. The most significantdifferences between the sets of pressure measurements is inthe magnitude of some of the peak values, particularly forthe P1 vs P12 and P5 vs P9 comparisons, shown in Fig. 14.The differences in peak pressures are 33.5% (P1 vs P12),3.6% (P3 vs P6) and 10.4% (P5 vs P10) 200 mm from thecentreline and 20.5% (P1 vs P11), 5.9% (P3 vs P8) and23.3% (P5 vs P9) 400 mm from the centreline. Of note isthat the P1 pressures were 16.7% different between the twotests, for the same impact conditions. The differencebetween the peak pressure at P11 and P1 from the firsttest is 4.8%, showing that the variability in the peak

pressure measurements can be greater than the variationacross the width of the panel.

An alternative to comparing peak pressure is to comparethe impulse applied to each transducer. This is calculated byintegrating the pressure, multiplied by the area of thetransducer, with respect to time. The differences betweenimpulse measured at the centreline and those 200 mmtowards the panel edge are 10.2% (P1 vs P12), 22.7% (P3vs P6) and 6.2% (P5 vs P10) while at 400 mm from thecentreline the impulses are 23.4% (P1 vs P11), 17.0% (P3vs P8) and 16.5% (P5 vs P9) different. There are severalpossible reasons for the variations. In addition to possiblesampling rate effects, the fluid boundary conditions createdby the flow restriction panels may be such that the flow isnot completely two dimensional which would create avariation in the pressure along the length of the panel. Thevariation in the timing between P5 and P9 and the slightdifference between the timing for P3 and P8 indicates thatthat there is a small reduction in the flow front velocity

-200

0

200

400

600

800

1000

1200

0 5 10 15

Pre

ssur

e [k

Pa]

Time [ms]

P1P11P3P8P5P9

Fig. 14 Comparison between centreline pressures and P8, P9 & P11

Fig. 15 System displacement and velocity from repeated 5 m/s tests

Fig. 16 Pressure time-history from repeated 5 m/s tests

-100

0

100

200

300

400

500

600

700

800

900

0 2 4 6 8 10 12 14 16 18 20

Pre

ssu

re [

kPa]

Time [ms]

Fig. 17 Pressure time-history from repeated 5 m/s tests withcorrelated timing

102 Exp Mech (2012) 52:95–106

towards the ends of the panel. The good correlationbetween the timing of P10 and P5 and P6 and P3demonstrates that the flow front is planar within the centralregion of the panel.

It is also noted that the peak pressures are not constant atP1, P3 and P5 as would be predicted by an analytical theorysuch as Wagner [1]. The peak pressure at P1 is observed asbeing lower than P3 and P5 in all impacts. It appears that adepth of immersion must be reached before the pressurebuilds to its maximum. The reduction in pressure due to thetransducer size at P1 is 38.3%, larger than the reductionpredicted at either P3 or P5. Peak pressures at P3 and P5 aremuch closer in magnitude, however they are still not equal.These differences are unexplained at this stage, highlightingthe difficulty in measuring the peak pressure as well as thecomplexity of the loading situation.

Repeatability

The repeatability of the system has also been charac-terised using the rigid panel. Figure 15 presents the

displacement and velocity profiles from two equivalenttests at a target velocity of 5.0 m/s undertaken 3 yearsapart with the same panel. The average velocity from thefirst test series (S1) was 4.74 m/s while the second testseries (S2) had an average velocity of 4.58 m/s, adifference of 3 %. Figure 16 shows the pressures recordedat transducers P1, P3, and P5 with positions as shown inFig. 12, the solid lines being from S1 and the dashed fromS2. The difference in the timing can be attributed to theslight variation in velocity profile between the two tests.Figure 17 shows the pressures with the time shifted sothat initial pressure rise coincides for comparable trans-ducers. The shapes of the pressure time-histories aresimilar for the two tests, using least-squares method thecoefficients of determination are 0.659, 0.924 and 0.865for P1, P3 and P5 respectively. The difference in impulseat each of the transducer locations is 11.3 %, 4.4% and0.8% for P1, P3 and P5 respectively. While P3 and P5have close correlation seen both from the high coefficientof determination and the low difference in impulse thereis a non-negligible difference at P1. The most significantdifferences appear in the residual pressure and an unusualstep in the pressure shortly after the peak pressure for P1.The comparison highlights the reasonable level ofrepeatability of the system.

Deformable Sandwich Panel Response

A deformable sandwich panel was fabricated and tested inthe SSTS in order to establish the difference in pressuresbetween a rigid and flexible body. The design of the panel(Table 4) was chosen to represent a realistic construction fora high speed craft. The panel was manufactured using thesame VARTM process as the rigid panel.

The deformable sandwich panel was subjected toslamming impacts at speeds of 1 to 5 m/s. The maximum

Table 4 Deformable panel specifications

Materials

Lamina Material Glass Fibre / Epoxy

Lamina Orientation (0/±45/90)4Core Material PVC, C70.140

Dimensions

Skin Thickness, tf 2.5 mm

Core Thickness, tc 15 mm

Panel Dimensions 1030×565 mm

Calculated Properties

Flexural Rigidity, D 6,730 Nm

Shear Stiffness, S 960,000 N/m

-50

0

50

100

150

200

250

300

350

400

-2

0

2

4

6

8

10

12

14

16

0 5 10 15 20 25 30 35 stra

in [

mic

rost

rain

], P

ress

ure

[kP

a],

Pan

el D

isp

lacm

ent

[mm

]

Time [ms]

Panel Displacement [mm] Pressure 1 [kPa] Pressure 3 [kPa]

Pressure 5 [kPa] Strain 3 [microstrain] Strain 5 [microstrain]

Fig. 18 Pressures, strains andpanel deformation time historiesfor slamming of deformablepanel at 3 m/s and 10°

Exp Mech (2012) 52:95–106 103

velocity used was limited by the shear strength of the corematerial. Figure 18 presents typical pressure, panel dis-placement and strain results for a test at 3 m/s. Thepositions of the transducers are as defined in Fig. 12, withpressures measured at P1, P3 and P5, panel displacement atthe centre of the panel (adjacent to P3), and strainsmeasured on the inner panel skin at the centre (Strain 3,adjacent to P3) and near to the edge of the panel (Strain 5,adjacent to P5). The strain gauges are positioned 50 mm tothe side of the pressure transducers.

Figure 18 shows that the panel does not begin to deformuntil after the pressure pulse reaches P1 and that themaximum deformation of the panel occurs at a similar timeas the pressure pulse reaches P5. This is expected, as at thisstage the entire panel is being loaded, with most of thepanel subjected to a distributed residual pressure, combinedwith the high pressure at the flow front near to the edge ofthe panel. The maximum strain at the centre of the panel(Strain 3) occurs slightly earlier than the maximumdeflection, while there are only very small strains near tothe edge of the panel (Strain 5) until quite late in theslamming event, but then this strain increases rapidly as theflow front nears the panel edge.

Figures 19, 20 and 21 show the resulting pressure time-histories from transducers P1, P3 and P5 (positions as inFig. 12) compared to the corresponding rigid panel resultsfor impact velocities of 1, 2 and 3 m/s respectively. In allcases the resulting pressure profiles at P1 are similar for therigid and deformable panels. This is expected as the

deformable panel does not deflect significantly until laterin the slamming event as shown in Fig. 18.

The pressures at P3 (panel centre) are significantlydifferent, particularly the residual pressure which is initiallyhigher for the deformable panel, then reduces to a similarlevel as the rigid panel at the end of the slamming event.The difference in the residual pressure is greater at higherimpact velocity. Multiple tests were completed at eachspeed, each showing the same pressure characteristics.

These differences are believed to be related to transientkinematic behaviour, where both the impact velocity andangle spatially vary as a function of the instantaneousdeformed shape of the flexible panel during the slammingevent. Figure 18 shows that the centre of the deformablepanel deforms up to a maximum of approximately 8 mmwhen impacted at 3 m/s and 10°. This has the effect ofreducing the velocity of the centre of the panel relative tothe water to approximately 2.2 m/s, which is likely to affectthe magnitude of the peak pressure. As the panel reachesthe maximum deflection its relative velocity increases backto the velocity of the testing fixture, then increases furtheras the panel rebounds towards, and then beyond its initialposition. This is believed to be the cause of the increasedresidual pressure at the panel centre.

The deformation in the vicinity of P5 (close to the panelboundary) is also significant, reaching approximately 4.3 mmat 3 m/s impact velocity. This deformation is measuredapproximately 35 mm from the boundary support, demon-strating that there is appreciable rotation of the panel in this

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Fig. 20 Pressure comparisonbetween rigid and deformablesandwich panels for 2 m/s

104 Exp Mech (2012) 52:95–106

region, presumably partly due to transverse shear deformationof the panel. The deformation has the effect of reducing thelocal deadrise angle, which would be expected to increaselocal pressures. An increase of 13.7% was measured at 3 m/scompared to a rigid panel. Significant increases were notobserved at 1 & 2 m/s impacts most likely due to the level ofdeflection being much smaller.

The strain at P5 is expected to be proportional to thetransverse shear force [25]. Figure 18 shows that the strainincreases rapidly as the pressure peak passes, due to the highresidual pressure combining with the moving peak pressure.This is of particular concern for sandwich panels as this isthe region most highly loaded in transverse shear, and theposition where core shear failure normally occurs, as shownin Fig. 22. This supports conclusions of previous studies thatslam loaded sandwich panels experience significantly highertransverse shear forces than uniformly loaded panels [23].

Conclusions

A high-speed servo-hydraulic testing system has beendeveloped to enable water slamming impact tests of marinehull panels to be undertaken in a controlled laboratoryenvironment. Panels with an unsupported region of 1000×500 mm can be impacted into a body of water with a volumeof approximately 13,500 L at speeds of up to 10 m/s. Thedeadrise angle can be set to be from zero to 40° in 10°increments.

A closed loop control system uses displacement andacceleration feedback to achieve the targeted motion. Thecontrol parameters can be defined in different waysdepending on the complexity of the desired motion,

including nominally constant velocity or a changing velocity,such as increasing or decreasing velocity during the test, or theinput of an arbitrary velocity profile. Instrumentation enablesmeasurement of transient pressures, displacements, acceler-ations, load and strains.

The performance of the system has been characterised,and nominally rigid and deformable sandwich compositepanels tested. The pressure magnitudes and time-historiesresulting from testing of the nominally rigid panel havebeen compared with previous analytical, semi-empiricaland experimental studies demonstrating that the system cansuccessfully reproduce complex slamming impacts in alaboratory environment. The spatial distribution of pres-sures has been characterised, demonstrating reasonableuniformity of the pressure field across the impacting panel.

It has been demonstrated that slamming impacts on adeformable sandwich panel result in different peak andresidual pressures to those from a rigid panel for impactvelocities when hydroelastic effects become significant.The residual pressure at the centre increases significantlyfor the flexible panel, and there are indications ofreductions in peak pressures at the centre of the panel.Large local strains are observed at the chine edge of theflexible panel, which are believed to be due to the highresidual pressure combining with the moving peak pressure.This is of particular concern for sandwich panels as this isthe region most highly loaded in transverse shear, and theposition where core shear failure normally occurs.

Despite the complexity of the testing system and thesubstantial development effort undertaken, some differ-ences are observed in pressures between this study,alternative analytical solutions and previous experimentalstudies. This highlights the complexity of slammingphenomenon and the need for continued development ofadvanced experimental facilities to enable meaningfulvalidation of predictive models.

Future Work

Further testing will be undertaken with an increased rangeof panel stiffness and constructions. The results of these

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Fig. 22 Core shear fracture of foam core sandwich panel at mid-pointof chine edge

Exp Mech (2012) 52:95–106 105

tests along with those already completed will be used tovalidate dynamic numerical models in order to comprehen-sively characterise the impact of flexible panels underslamming loads and better understand the influence ofhydroelastic coupling on the loads and panel responses.

Acknowledgements This research has been financially supported bythe USA Office of Naval Research (Grant N00014-08-1-0136) withDr Y. Rajapakse as programme manager. The assistance of IndustrialResearch Limited in providing access to the SSTS and SP-HighModulus (NZ) in providing technical guidance, materials supply andspecimen manufacturing is also gratefully acknowledged.

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