Aalborg Universitet

The behaviour of the stiff monopile foundation subjected to the lateral loads

Lada, Aleksandra; Ibsen, Lars Bo; Nicolai, Giulio

Publication date:2014

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Citation for published version (APA):Lada, A., Ibsen, L. B., & Nicolai, G. (2014). The behaviour of the stiff monopile foundation subjected to thelateral loads. Department of Civil Engineering, Aalborg University. DCE Technical Memorandum, No. 47

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ISSN 1901-7278

DCE Technical Memorandum No. 47

The behaviour of the stiff monopile foundation subjected to the lateral loads

Aleksandra Lada

Lars Bo Ibsen

Giulio Nicolai

DCE Technical Memorandum No. 47

The behaviour of the stiff monopile foundation subjected to the lateral loads

by

Aleksandra Lada Lars Bo Ibsen Giulio Nicolai

June 2014

© Aalborg University

Aalborg University Department of Civil Engineering

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Published 2014 by Aalborg University Department of Civil Engineering Sohngaardsholmsvej 57, DK-9000 Aalborg, Denmark Printed in Aalborg at Aalborg University ISSN 1901-7278 DCE Technical Memorandum No. 47

The behaviour of the stiff monopile foundation subjected tothe lateral loads

Aleksandra Lada Lars Bo Ibsen Giulio Nicolai

Aalborg University

Department of Civil Engineering

Abstract

The article focuses on the analysis of a large-diameter monopile foundation for offshore windturbine based on the numerical model results. The case describes the behaviour of a monopile insand subjected to a lateral loading conditions. The effects of the pile diameter, the length andthe load eccentricity on the results are investigated. Additionally, the cyclic nature of load isalso taken into consideration. The cyclic loading leads to an accumulated rotation, thus the SLSdesign must ensure that the maximum permanent rotation does not exceed the limit value fora given wind turbine. The approach analyzed by LeBlanc et al. [2010] and Lada et al. [2014] isdiscussed in the way of its use as a preliminary design.

1 Introduction

The monopile foundation is the most oftenused concept for wind turbines in the off-shore conditions. The greater demand for therenewable energy is the reason why biggerand more effective turbines are introducedin the offshore field. This has a significantinfluence on the foundation design. Not onlythe vertical load from the turbine itself arebigger, but also the overturning moment work-ing cyclically on the pile increases. As thewind working on the rotor has its center ata greater height, the arm for the load increases.

Nowadays a large diameter monopiles areused for offshore wind turbines and they areconsidered highly successful. Nevertheless,the current design of monopile embedded insand leaves a lot to be desired. The most often

used standards, DNV [2011] and API [2005]recommend the p-y curve design, which de-scribes the non-linear relationship between thesoil resistance and the lateral pile deflectionat each depth below the seabed. The methodis based on the full-scale tests of slender pilesand it found its confirmation to a greater orlesser extent by other full-scale tests. However,the increase of the diameter was found byother researchers to change the behaviour ofpile under lateral load and therefore the resultsof pile deflection do not fit the p-y curve anylonger.Currently, the attention of many engineers isdevoted to the consideration of the cyclic loadsin the design standards. Both, API [2005] andDNV [2011] recommend only the decrease ofthe ultimate soil resistance, when accountingfor a cyclic loading. Nor the load characteris-tics, neither the number of cycles are included.

5

THE STATE OF ART

The comparison between the p-y method anda fine element method for the response of alarge-diameter pile subjected to a lateral loadwas performed by Lesny et al. [2007]. Themodels of two piles, D= 1 m and D= 6 m, em-bedded in non-cohesive soil were investigated.An elasto-plastic material behavior assumingthe Coulomb friction law was the basis forthe analysis. Both piles were designed byproviding sufficient length for a rigid fixationand the limited pile head rotation, αlim = 0.7◦.As long as for smaller pile both results arequite comparable, for the large-diameter pilethe results varies significantly. The p-y curveapproach overestimates the pile-soil stiffnessof a larger pile at a great depth, questioningthe use of this method. Both piles were charac-terized by different diameters and embeddedlengths, therefore it seems to be reasonable tostate, that both these parameters influence theresponse of horizontally loadded piles.

In the research of Achmus et al. [2009] thenumerical model was used in the analysisof laterally-loaded monopile. The results ofdrained cyclic triaxial tests on cohesionless soilwas applied in the model, so the cyclic condi-tion could be reconstructed. As the increaseof plastic strain with the number of cycles wasobserved, it was interpreted as a decrease insoil secant stiffness. Therefore the model wasnamed the degradation stiffness model. Thesoil was modeled as en elasto-plastic with aMohr-Coloumb failure criteria and both, denseand medium dense sand were analyzed.Due to an investigation of different pile lengths,the difference in the behaviour of a stiff anda slender pile was observed. Due to appliedcycles the increase of deformation and thelowering of the rotation point of a pile wasnoticed. From the results it was stated thatthe pile performance is very much dependenton the embedded pile length, whereas theincrease of the pile diameter from 5 m to 7.5

m decreases the accumulated displacementof the pile only slightly. The design criterionconsidering the zero-toe-kick in order to min-imize the risk of accumulated deformationunder cyclic loading was claimed to be in-adequate for the stiff piles. The piles withthe same length, but different diameters wereinvestigated and the stiffer pile that behavesmore rigid had a smaller deformation on themudline. Additionally, a strong dependencebetween the magnitude of the applied loadand the accumulated displacement was found.

Many others publications describe the nu-merical models used in analyzing the responseof laterally loaded piles. Some of them sug-gest the new expressions for p-y curves forthe large diameter piles. Generally, a morerigid behaviour for less slender piles is proved.This results in the overestimation of the p-ycurve stiffness on the great depth and causealso the underestimation of the rotation onseabed, [Abbas et al., 2008],[Augustesen et al.,2009],[Onofrei & Ibsen, 2010].

SUBJECTS OF INTEREST

The numerical model has been made and theresults of performed simulations are used toanalyze the behaviour of a large diameter pile.The strong attention is also paid on the depen-dency of pile properties, like the diameter andthe length and also the load eccentricity on thepile design.In order to account for a cyclic loading a newapproach based on LeBlanc et al. [2010] andLada et al. [2014] accompanied with a numeri-cal calculation is presented.

6

2 The numerical model ofmonopile

For numerical calculation a geotechnical pro-gram PLAXIS 3D is used, as it is able tosimulate the soil condition with a good accu-racy. The analysis is focused on the laterallyloaded monopile embedded in a drained sand.

From a different available material modelsin PLAXIS 3D, the Soil Hardening Small Strainmodel is chosen, as a one of more advanced.The model assume the hardening of soil dueto the plastic strains. The hardening due tocompression and the hardening due to devia-toric loads (shear hardening) are distinguishedin the model. Those two loads and also reload-ing of soil are controlled by different stiffnessparameters, E50, Eoed and Eur. The relation be-tween the axial strain and the deviatoric stressis modelled to be hyperbolic. Additionally themodel adopts the dependency between thestiffness and the stress level, by the parameterm.In contrast to the basic Hardeing Soil model,a non-linear decrease of the stiffness due tothe plastic strains is included. This change iscontrolled with the small-strain shear mod-ulus, G0, and the parameter providing thelevel of shear strain at which the reductionof the secant shear modulus to about 70 % isobserved, γ0.7. [Brinkgreve et al., 2012]

MATERIAL PARAMETERS

The advanced model requires a large numberof input soil parameters in order to provide arealistic prediction of the soil behaviour. Threedifferent state of sand are used in simulations:dense, medium dense and loose sand. The rel-ative density, ID, are chosen to be 85 %, 40 %and 5 % respectively. The required soil param-eters are obtained from the data of the coneresistance, qc. There is no specific location cho-sen and no data from CPT available, therefore

the data of qc are evaluated based on ID, equa-tion (1) [Lunne et al., 2007]. The atmosphericpressure, pa, is set to be 100 kPa.

ID =1

2.96ln

qcpa

24.94 ·(

σ′v0·(

1+2·K03

)pa

)0.46

(1)

The consolidation of the sand is assessed inorder to provide a reliable formulations for therest of the parameters. The friction angle isobtained from relation in equation (2). Thepore pressure is neglected, so the total coneresistance, qt, is equal to qc.

ϕ′ = 17.6 + 11 · log

qtpa(

σ′v0pa

)0.5

(2)

The stiffness parameters are given in the de-pendency on the oedometer stiffness, Eoed, ob-tained from equation (3) for normally consol-idated sand and from equation (4) for over-consolidated sand. The secant stiffness E50 isrelated to the oedometer stiffness through thePoisson’s ratio as it is indicated in equation (5)and the unloading-reloading stiffness, Eur, canbe approximated with equation (6).

Eoed = qc · 101.09−0.0075·ID (3)

Eoed = qc · 101.78−0.0122·ID (4)

where ID is in %.

E50 = Eoed ·1− ν− 2ν2

1− ν(5)

Eur = 3 · E50 (6)

The shear modulus obtained from equation (7)is set as the small-strain shear modulus. γ0.7is calculated according to recommendation forthe program [Brinkgreve et al., 2012].

G = 1634 · q0.25c · σ′0.375

v0 (7)

In order to avoid numerical instabilities, thevalue of cohesion is set to be 0.1 kPa. The listof soil parameters used in the programm are

7

presented in Table 2.1. The reference values forstiffness are chosen as the ones correspondingto σ′v0 = 100 kPa.

Table 2.1: Soil parameters used in PLAXIS

ID [%] 5 40 85ϕ [o] 32.56 37.55 44.49e [-] 0.843 0.734 0.595γ [kN/m3] 18.90 19.46 20.28OCR [-] 0.67 1.43 3.98K0 [-] 0.462 0.485 0.788Ere f

50 [MPa] 18.30 103.520 142.60Ere f

oed [MPa] 25.84 132.61 150.00Ere f

ur [MPa] 54.91 310.57 427.80νur [-] 0.21 0.21 0.21Gre f

0 [MPa] 69.04 93.25 177.00γ0.7 [-] 2.14e-7 1.71e-7 1.44e-7Rinter 0.57 0.61 0.71

For the monopile a steel material is used. Thestiffness is set to 210 GPa and the Poisson’sratio to 0.3. The wall thickness is chosen tobe 0.06 m. The additional extension of pileis made in order to apply the horizontal loadon the sufficient arm, creating the appropriatemoment on the seabed. The stiffness of thepile extension is chosen to be much larger andits unit weight is chosen to be small, so therewill be no effect of the additional part on thepile-soil interaction.

PLAXIS MODEL

The model boundaries was chosen based onthe numerical model used by Abbas et al.[2008]. Chosen boundaries can be seen in Fig-ure 2.1.

The model is divided into number of ele-ments, so that the finite element calculationscan be performed. Two different kinds of ele-ments are used in the program to simulate anaccurate soil behaviour and a pile behaviour.The third kind of elements are used for the

interface between pile and soil, what is reallyuseful for investigating the soil-pile interac-tion. The parameters for soil on the interfaceare obtained by using the reduction factorRinter = tan(δ)/tanϕ, where δ is the angle ofsoil friction on the pile wall.

Figure 2.1: The model boundaries with the meshexample

The overall mesh is chosen to be coarse. Asthe most interesting results are expected nearthe pile, therefore the mesh is refined in thesurroundings of the pile. There are two chosenregion for mesh refining. The region closer tothe pile is more refined. After performing aconvergence analysis an appropriate Fineness-Factor is chosen for the soil volumes close tothe pile and the pile surface. An example ofmeshed model is presented in Figure 2.1.

CALCULATION PHASES

The calculation for a laterally loaded monopileare divided into five calculation phases. Theanalysis is initiated by the initial phase, wherethe initial soil stress are obtained. In the phasea K0 procedure is used. This phase is followedby the nil-step phase, where no changes inactivating parts is made, but the plastic calcu-lations are used. The phase has no physicalmeaning and it is used only in order to ob-

8

tain an equilibrium of stress before applyingthe load. In the third phase, the constructionphase, the plates are activated and in the nextphase the vertical load is applied in order tosimulate the self-weight of the wind turbine.In both phases the plastic calculation are cho-sen. In the final loading phase the horizontalload is additionally activated. The load isapplied as a point load.

The determination of exact failure in the nu-merical calculation is found as difficult. There-fore the failure is assessed for a 4 o of the pilerotation on the seabed. In case, where the cal-culation fails before obtaining the sufficientlylarge displacement, the tolerable numericalerror is slightly increased.

3 The analysis of results

TEST PROGRAM

Different aspects of the pile behaviour are as-sessed and therefore different geometries andmodel properties are chosen for simulations.The test programs can be seen in Table 3.1

Table 3.1: Test program for PLAXIS simulations

Model D [m] L [m] e [m] ID [%]1 1 11 1.2L 852 5 11 1.2L 854 6 25 1.2L 855 7 25 1.2L 856 5 30 1.2L 857 5 35 1.2L 858 5 25 1.0L 859 5 25 1.4L 8510 5 25 1.2L 4011 5 25 1.2L 5

PILE BEHAVIOUR

In order to assess the difference in the be-haviour of the stiff and the slender pile theresults of model 1 and 2 are investigated. Inboth cases the same length of the pile waschosen, but there is a difference in the pilediameter. The horizontal load is applied at thesame height. The dense state of sand is chosen.The lengths of piles were chosen based oncriterion proposed by Poulus and Hull (1989),where the pile behaviour is assessed as a rigidor flexible according to the soil and the pilestiffness and the pile geometry.

The ULS was found for both situations. Thedifference in the behaviour is analyzed when30 % of the ultimate moment capacity is ap-plied to the piles. The different patern ofdeformation is presented in Figure 3.1.

Figure 3.1: The different pattern of behaviour forthe slender (model 1) and the stiff pile (model 2)

Clearly, the slender pile gives the flexible re-sponse, whereas the rigid pile gives more stiffresponse. For the pile with the large-diameterthere is a visible deformation on the pile topresulting in additional shear stresses. Eventhough the much larger force is applied forthe stiff pile, the displacement on the seabedare significantly smaller. It can be concludedthat the large diameter piles have a better lat-

9

eral performance in comparison to the slenderpiles.

DEPENDENCY OF PILE PROPERTIES

The diameter of monopile is changed whilekeeping the others properties fixed. The effectsof the diameter on the ultimate moment capac-ity on the seabed are presented in Table 3.2.The ultimate soil-pile resistance is importantfor the design, as it is used for the p-y curve.As the rotation of the pile on the seabed isoften assessed for SLS, the change in the mo-ment capacity on the seabed are interesting toinspect.It can be clearly stated that the diameter hasa significant effect on the moment capacity onthe seabed. The diameter increase of 20 %gives a rise of more than 42 % in the momentcapacity.Additionally, the displacement of the pile,while applying 30 % of ULS from the mainmodel (model no 3) are inspected in Figure3.2.

Figure 3.2: Pile deformations for varrying D,L = 25m, e = 1.2L - H = 19.3 MN

It can be seen that the pile behaves more rigidlyfor all cases and the increase in the diametergives a better pile performance.

The same analysis is performed for differentpile lengths, while keeping other propertiesfixed. The results of the ultimate momentcapacity can be seen in Table 3.3 and the defor-mation for the same applied horizontal forceare plotted in Figure 3.3.

Figure 3.3: Pile deformations for varrying L,D = 5m, e = 1.2L - H = 19.3 MN

The increase in the length has an influence onthe ultimate capacity by increasing it. The 20% increase of length gives a 13 % increase ofthe moment capacity. It can be concluded thatthe change in the diameter is more effectivefor increasing the ultimate capacity. From thedeformation graph it can be concluded thatfor applied 30 % of the ultimate moment, thedifference in the displacement on the seabeddoes not varry significantly.

The load eccentricity is also inspected in orderto assess its influence on the moment capac-ity. The results can be seen in Table 3.4. Thechange in the load eccentricity have also aneffect on the ultimate capacity of the soil-pilesystem, however this parameter among othersis assessed to be the least meaningful.On the other hand, where the deformationof the pile for the same horizontal load isapplied on the pile, it can be seen that the

10

load eccentricity has a bigger influence on thedisplacement on the seabed in comparisonwith the change of the length, Figure 3.4.

Figure 3.4: Pile deformations for varrying e,D = 5m, L = 25 m - H = 19.3 MN

As the change in the ultimate capacity issmaller, but the change in the deformation ofthe pile is more significant, the explanation forthis is searched somewhere else. After inspect-ing the stiffness no change in the initial partis observed when the pile length is increased.When the load eccentricity is increased, thedecrease in the rotational stiffness in its initialpart is revealed. The results are presented inFigure 3.5.

Figure 3.5: The inspection of the initial rotationalstiffness

Table 3.2: Results of the ultimate moment capacityfor different pile diameters

Pile diameter, Moment capacity,D [m] MR [MNm]

5 1934.496 2754.507 3395.91

Table 3.3: Results of the ultimate moment capacityfor different pile lengths

Pile length, Moment capacity,L [m] MR [MNm]

25 1934.4930 2183.7835 2262.05

Table 3.4: Results of the ultimate moment capacityfor different load eccentricity

Load eccentricity, Moment capacity,e [m] MR [MNm]1.0 L 1835.381.2 L 1934.491.4 L 2008.36

Finally the difference in the soil density on theultimate capacity are inspected. The resultsfor all three chosen relative densities are indi-cated in Table 3.5. The soil density is foundto be the factor affecting the ultimate capacitysignificantly.

Table 3.5: Results of the ultimate moment capacityfor different soil densities

Relative density, Moment capacity,ID [%] MR [MNm]

5 467.2440 1106.2085 1934.49

11

4 The cyclic lateral load

LeBlanc et al. [2010] proposed a new methodfor determining an accumulated rotation for along-term cyclic loaded monopile. Based onthis research, the monopile response for lat-eral load in medium-dense and loose sand canbe obtained. The approach was modified byLada et al. [2014] for a dense sand. In bothresearches the cyclic test are described withthe loading characteristics ζb (ζb = Mmax/MR)and ζc (ζc = Mmin/Mmax).

ACCUMULATED ROTATION DUE TO THECYCLIC LOADING

For the accumulated rotation an exponentialdependency on the number of cycles was re-vealed. The displacement due to the cyclicloads can be approximated from equation (8).

∆θ(N)

θs=

θN − θ0

θs= Tb · Tc · Nn (8)

For medium-dense and loose state of sand theparameters n was chosen to be 0.31. For densesand the value was found to be smaller and itdiffers between tests. However, the mean valuecan be chosen, n = 0.14. The dimensionlessfunctions can be predicted based on Figure 4.1and Figure 4.2.

Figure 4.1: The dimensionless function Tb

Figure 4.2: The dimensionless function Tc

The approach can be used as a preliminarydesign, however it requires the information ofthe static corresponding rotation.

The ULS for three different state of sand isfound in PLAXIS. The pile of D = 5 m andL = 25 m is chosen as an example. The hor-izontal load is applied on the arm of 1.2 · L.The FLS is assessed for 30 % of ULS, there-fore corresponding load magnitude is appliedand the rotation is obtained. The results areindicated in Table 4.1.

Table 4.1: The static rotation corresponding to 30 %of ULS

ID [%] Static rotation [o]5 0.446

40 0.60385 0.837

EXAMPLE

The load characteristics of ζc = -0.9 is cho-sen and the number of cycles is predicted tobe 107. The pile is embedded in dense sand.The accumulated rotation , ∆θ(N), can becalculated as following:

∆θ(N) = Tb · Tc · Nn · θs =

0.56 · 0.4 · (107)0.14 · 0.837 = 1.79o

12

The most often the tolerable criteria for thepermanent rotation is set to be 0.2o, so the ob-tained value exceeds this limit. However, inthis case all cycles are working in the samedirection, and as the multi directional load-ing is rather expected, the permanent rotationwould be smaller. Nevertheless, when ζc is setto be ≈ −0.4÷−0.6 the accumulated rotationwould be much larger.

5 Conclusion

A number of simulations for the laterallyloaded monopiles embedded in drained sandhave been performed. Different pile geome-tries and model properties have been used inorder to analyze the changes in the ultimatesoil resistance. Currently, the most often usedapproach for the design of offshore monopileis the p-y curve approach. The relation be-tween the pile lateral deformation and the soilresistance is based inter alia on the ultimatesoil capacity. According to the standards, onlythe pile diameter and the soil properties havethe influence on the value of the ultimate resis-tance. Moreover, the p-y curve was formed fora slender pile. Currently the large diameterpiles that behave more rigidly are used in theoffshore sector.

The results of PLAXIS simulations confirmthat there is a difference in the deformationpattern between the slender and the stiff pile.The former behave flexible under the lateralloads, whereas the latter show a more rigidresponse. The results of different simulationsreveal that the ultimate moment capacity isdependent not only on the pile diameter andsoil properties, as it is stated in DNV [2011]and API [2005]. The length and the load ec-centricity was found to influence the bearingresistance as well. However, the pile diameterand the change in the soil density are foundto be the most significant, wheres the load ec-

centricity is considered the least contributing.Nevertheless, when the deformation of the pilefor applied 30 % of ULS are inspected, the loadeccentricity was found to be more meaningfullthan the length of the pile. This is explainedwith the changes in the initial stiffness whenchangning the load eccentricity.A method for designing the offshore monopileshould be improved, so that all the parametersinfluencing the pile displacements will be in-cluded.

Also the cyclic response for the monopileshould be accompanied by the new calculationapproach. The proposition given by LeBlancet al. [2010] should be still investigated for itsvalidation and confirm by the full-scale results.

References

Abbas, Jasim M, Chik, Zamri Hj, & Taha,Mohd Raihan. 2008. Single Pile Simulationand Analysis Subjected to Lateral Load. Elec-tronic Journal of Geotechnical Engineering, 13-E,1–15.

Achmus, M., Kuo, Y., & Abdel-Rahman, K.2009. Behavior of monopile foundations un-der cyclic lateral load. Computers and Geotech-nics, 36, 725–735.

API, American Petroleum Insitute. 2005. Recom-mended practice for planning, designing and con-structing fixed offshore platform. Downloaded:10-10-2013.

Augustesen, A.H., Brødbæk, K.T., Møller, M.,Sørensen, S.P.H., Ibsen, L.B., Pedersen, T.S.,& Andersen, L. 2009. Numerical Modellingof Large-Diameter Steel Piles at Horns Rev.Proceedings of the Twelfth International Con-ference on Civil, Structural and EnvironmentalEngineering Computing, Paper 239.

Brinkgreve, R.B.J., Engin, E., & Swolfs, W.M.2012. Manual for PLAXIS 3D 2012. Down-loadet: 02-05-2014.

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DNV, Det Norske Veritas. 2011. Design of Off-shore Wind Turbine Structures. Downloaded:10-10-2013.

Lada, A., Gres, S., Nicolai, G., & Ibsen, L. Bo.2014. Response of a stiff monopile for a long-term cyclic loading. Department of Civil Engi-neering, Aalborg University, part of a Masterthesis.

LeBlanc, C., Houlsby, G.T., & Byrne, B.W. 2010.Response of stiff piles in sand to long-termcyclic lateral loading. Geotechnique 60, No.2,79–90.

Lesny, K., Paikowsky, S.G., & Gurbuz, A. 2007.Scale effects in lateral load response of largediameter monopiles. Contemporary Issues inDeep Foundations, Geotechnical Special Publica-tion, 158.

Lunne, T., Robertson, P.K., & Powell, J.J.M.2007. Cone penetration testing in geotechnicalpractice. Handbook, nos. ISBN: 0–419–23750–X. E & FN SPON - an Imprint of Routledge.

Onofrei, M.G., & Ibsen, L.B. 2010. Numericalanalysis of lateral deflection for monopilesin drained sand. Department of Cuvil Engi-neering, Aalborg University.

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Recent publications in the DCE Technical Memorandum Series

DCE Technical Memorandum no. 45, Response of a stiff monopile for a long-term cyclic

loading, Aleksandra Lada, Szymon Gres, Giulio Nicolai and Lars Bo Ibsen, June 2014.

DCE Technical Memorandum no. 46, The stiffness change and the increase in the ultimate

capacity for a stiff pile resulting from a cyclic loading, Aleksandra Lada, Lars Bo Ibsen and

Giulio Nicolai, June 2014.

DCE Technical Memorandum no. 47, The behaviour of the stiff monopile foundation

subjected to the lateral loads, Aleksandra Lada, Lars Bo Ibsen and Giulio Nicolai, June 2014.

ISSN 1901-7278

DCE Technical Memorandum No. 47