hard grounding of large sailing...

HARD GROUNDING OF LARGE SAILING YACHTS

H. Conradi, Hamburg University of Technology, Germany, [email protected]

H. Hoffmeister, DNV GL, Germany, [email protected]

S. Ehlers, Hamburg University of Technology, Germany, [email protected]

Running aground with a sailing yacht can result in significant damage to the structures of a yacht;

particularly at high speeds. The prediction of grounding forces and an associated maximum speed is

therefore an important precaution measure to remain inside the structural design envelope; to obtain

robust structures at predictable grounding scenarios. Since current classification rules do not explicitly

regard the vessel’s speed, an improved method of predicting grounding forces is desired. The aim of

this work is to investigate the forces that act on a sailing yacht during grounding. A nonlinear finite

element analysis of a large sailing yacht is performed, that accounts for plastic deformation of the keel

bulb, elastic deformation of the keel fin and hull as well as hydrostatic restoring forces. The influence

of the yacht’s initial velocity, draft and ballast/displacement ratio is investigated. Furthermore, the

grounding force experienced by a yacht with a crashworthy bulb is examined.

NOMENCLATURE

Symbol Definition Unit

Length between perpendiculars (m)

Draft (m)

Displacement, fully loaded (t)

Density (

Torsional spring stiffness ( )

Pitch Angle (deg)

Initial velocity ( )

Von Mises Stress ( )

Yield Stress ( )

Engineering Stress ( )

True Stress ( )

Engineering Strain ( )

True Strain ( )

Strain Rate ( )

1 INTRODUCTION

Grounding of a sailing yacht can differ significantly from

other vessels, due to the yachts unique geometry. The

keel, in which the ballast of the yacht is concentrated,

extends significantly below the actual draft of the hull.

Modern yachts often feature a “T-keel”, that consists of a

keel fin with a foil shaped cross section and a lead bulb,

attached to the bottom end of the fin. Grounding of a T-

keel yacht is likely is likely to happen as a collision

between bulb and ground. A severe case of grounding is

“horizontal grounding”, in which the bulb hits a rigid

wall, frontally. The collision forces act in the opposite

direction of the vessels forward movement, resulting in

large reaction forces due to inertia and most likely in

damage to keel and hull structure. A worst case of

damage is the total loss of the bulb, resulting in a

dramatic reduction of hull stability.

In grounding, several factors influence the magnitude

of the forces, that act on the keel and hull. The vessel’s

mass and velocity prior to grounding determine its

kinetic energy and impulse. Plastic deformation of the

bulb prolongs the collision time, thus reducing the

maximum force and absorbing kinetic energy. Elastic

deformation of the keel fin and hull structure from the

imposed bending moment result in further reduction of

the maximum force. The yacht’s pitch motion causes

waves, which absorb some energy, while hydrostatic

restoring forces increase the moment acting around the

transverse axis at the keel root.

More recently, initiatives are undertaken to improve

structural predictability and thus to reduce pecuniary

risks associated with grounding. Keel manufacturer

APM-Keels performed a finite-element analysis (FEA)

on a “crash-save bulb”, that relied on slits to reduce its

stiffness. In the simulation, the bulb impacted into a rigid

plate at an initial velocity of 12 m/s. The slits helped to

prolong the collision time between bulb and plate and

reduced the reaction force by 40%, compared to a regular

bulb [1].

References on conventional ship grounding or collision

analysis and consequence assessments can be found, i.e.,

in ICCGS [2]. Extensive research like this is not

available for sailing yachts. An assessment of the

grounding force by estimating the deceleration of the

The Fourth International Conference on Innovation in High Performance Sailing Yachts, Lorient, France

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yacht has been presented in literature [3]. In their rules

for yachts, classification society DNV GL provides the

following equation to assess the horizontal grounding

loads [4]:

( 1 )

With being the horizontal grounding force,

gravitational acceleration, displacement of fully

loaded vessel and bulb mass. This formula neglects

some aspects that are believed to be influencing the

magnitude of the grounding force, such as initial velocity

and stiffness of the yacht’s structure and bulb. Today,

owners and operators desire a clearer correlation between

boat speed and possible damage. Therefore, the present

research project is initiated by DNV GL to improve

methodologies and refine scantling rules.

It is assumed, that the bulb experiences large

deformations during grounding. Thus, it is deemed that

FE calculations will have to be performed as non-linear

type. To reduce computational effort, some

simplifications are made. It is assumed, that hydrostatic

restoring forces are greater than hydrodynamic forces,

acting on the hull due to the pitch movement of the

yacht. Therefore, hydrostatic pressure will be

implemented into the finite element (FE) simulation,

while dynamic effects are neglected. The elasticity of

keel fin and hull structure determine the overall stiffness

of the system and influence the maximum grounding

force. Based on the assumption that keel fin and hull

structure deform elastically, their stiffness can be

combined as a rotational spring that connects hull and

keel.

2 MODELLING

2.1 LS-DYNA

LS-DYNA is a keyword based, explicit finite element

solver, well suited for simulations that involve large and

plastic deformations. Simulations are performed on the

HPC cluster of the TUHH, using LS-DYNA version

7.1.2.

Meshing has been performed with ANSYS Workbench,

where the use of tetrahedron elements delivered the most

satisfactory results. Tetrahedron elements can experience

volumetric locking, which leads to an overestimated

stiffness, compared to hexahedron elements. To

overcome this, averaged nodal pressure (ANP)

tetrahedrons with four nodes and one integration point

are used. Strains and stresses are calculated as in

standard linear tetrahedrons, with additional averaging of

nodal pressures to reduce volumetric locking [5]. ANP

tetrahedrons are implemented into LS-DYNA as element

type 13 and are described to be well suited for

applications including incompressible material behaviour

of ductile metals with isochoric plastic deformations [6].

Plastic deformation of the bulb will be caused by

compression forces. Since most materials have different

a compressive strength than tensile strength, a material

model is chosen that differentiates between the two. The

material that is implemented as keyword *MAT_124, is

an isotropic elastic-plastic material. Two load curves

and are defined, representing yield stress

versus plastic strain for tension and compression [7]. The

sign of the mean stress is used for differentiation between

the two. A positive mean stress represents tension. Strain

rate effects are regarded by multiplying the stress-strain

curve with a strain rate dependent factor. This factor is

determined using the Cowper-Symonds model, that

calculates the strain rate dependent strengthening of the

material as follows:

( 2 )

In which is the strain rate, and are constants and

is the ratio between dynamic and static stress.

2.2 PROPERTIES OF LEAD-ANTIMONY ALLOYS

Lead’s high density of and

availability, compared to denser materials, make it the

ideal ballast for sailing yachts. Nonetheless, production

yachts often feature cheaper keels of cast steel. Pure Lead

has a low tensile strength of 17 MPa, compared to 400-

550 MPa of common steel. However, the addition of

alloying elements, such as antimony, strengthen the

material significantly. An antimony content of 3.5%

doubles the lead alloy’s tensile strength, whereas 8%

antimony alloys possess three times the strength of soft

lead. Since lead creeps under static load, it has no

definite yield strength. In yacht design, the antimony

content of the keel bulb must be balanced to be large

enough to support the bulbs own weight, while

maximizing the density of the ballast. Generally, more

performance oriented yachts feature lower contents of

antimony. Since the application of lead as a structural

material is rare, few sources about its mechanical

properties are available. While manufacturers provide

ultimate tensile strength and elongation [8], stress-strain

curves, influence of strain rate effects and compressive

strength were found in older publications [9]. Therefore,

material tests were conducted to validate the material

model used in the simulation.

2.2.1 Material Test Setup

Material Tests were performed with lead specimens of

1.5% antimony content. The dimensions of the

specimens were defined to satisfy DIN standards for

tensile [10] and compression tests [11]. Compression

tests were performed with cylindrical specimens,

featuring the same height as their diameter of

. The cylindrical specimens were placed

between two horizontal and parallel bearing surfaces.


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Testing was performed displacement controlled, using

hydraulic cylinders.

Tensile tests were performed with flat specimens

(Figure 2.1), that were fitted in serrated grips. The

dimensions were determined to avoid plastic deformation

within the grip area, by satisfying the following

condition:

( 3 )

With and being the cross-section areas of

gage and grip and and being yield strength and

ultimate strength. Since the yield strength is unknown

and strain rate effects may account for additional

ultimate strength, a ratio of of 0.2 is chosen.

The dimensions of the grip are determined by the testing

machine, which features a width of ,

resulting in a gage width of . The tensile

specimen feature a reduced section of with

a thickness of . The grip has a height of

. The total length of a specimen is

. Specimen were cast in an open, preheated

mould.

Figure 2.1: Flat specimen according to DIN 6892-1

2.2.2 Material Test Results

Tensile tests were performed in the strength laboratory of

the Institute M-10 at Hamburg University of Technology

at a strain displacement of , ruling out

strain rate dependency. The engineering stresses and

strains of individual specimens up to breaking varied

significantly (see Figure 2.2). Several probes failed due

to blowholes, enduring only 1/5 of the expected stress

and 1/10 of the elongation. Other specimens endured

twice the stresses and strains expected from literature

values. Since casting of the specimens took place over an

extended period of time, the molten material might have

segregated. Antimony may have deposited in the upper

part of the melting pot, due to its 40% lower density. The

increased antimony concentration could have resulted in

specimen with higher strength.

Figure 2.2: Tensile test results. Black, dashed curve

represents material curve selected for the FEA.

Based on the observations made above, the tensile test

that matches literature values best was chosen for the

stress-strain curve to be used in the simulation. The curve

has an engineering stress of 24 MPa and reaches an

engineering strain of 15% at breaking. These properties

align well with specifications by manufacturer Vulcan,

who states a tensile strength of 26 MPa at 16%

elongation [8].

Compression tests were performed at different strain

displacements, to evaluate strain rate effects.

Compression tests delivered much more consistent

results and showed a logarithmic influence of strain rate

effects (Figure 2.3). Chosen displacements are

, and .

When compared to the tensile test results, it stands out

that yield strength under compression is more than twice

of that in tension.


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Figure 2.3: Compression test results (previous page).

Black, dashed curve represents material curve selected

for FE. Yield stress versus Displacement (top). Black,

dashed line represents logarithmic trendline.

The significantly higher strength of lead under

compression than in tension, is important to note for the

simulation, since the stiffness of the bulb will be mostly

determined by its compression strength. For the material

curve to be used in the simulation, measured engineering

stresses and strains are converted into true stresses and

strains (see Figure 2.4):

( 4 )

With true strain , engineering strain , true stress

and engineering stress .

Figure 2.4: True stresses and strains. Negative values

represent compression.

The stresses and strains from the material tests were

then converted to be used in LS-DYNA. Discretizing a

geometry results in stiffness changes, based on mesh

resolution and used element type. Therefore, the material

curve obtained by material testing needed to be

calibrated to match the element size and type, that will be

used in the grounding simulation. This was realized by

analysing the material tests in a FEA. The element size of

the simulated specimen was chosen to be as coarse as

possible, since the same element size must be used in the

later grounding simulation. Figure 2.5 explains the

process of aligning the FE material curve with the

stresses and strains obtained by material testing. True

stress-strain curve from the material test is used as input

data for the first iteration of the remodelled material test.

The simulation result is compared to the result of the

actual material tests. The stress-strain curve is then

modified to improve agreement of the FEA and actual

material test results. This process is repeated, until the

simulation result matches the material test result.

Figure 2.5: Procedure of aligning the stress-strain curve

used in FE with the material test

After modifying the stress-strain curve under tension,

the elongation in the FEA reached 87% of the elongation

from the material test, compared to 82% of the

unmodified stress-strain curve. Due to a lack of necking

in the FEA, further accordance with the material test

could not be achieved. Under compression, the

elongation at yield strength was overestimated by 30%

and was reduced to 5%, by modifying the material curve.

In each of the figures below (Figure 2.6), three stress-

strain curves are shown. One resembles the result from

the material tests, the other two are the material curve

used in the last FEA iteration and the respective

simulation result. The stress-strain curve used in the

material definition of the last simulation will be applied

in the grounding simulation.


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Figure 2.6: Comparison of material test result with FE

input and FE output of last iteration for tension (top) and

compression (bottom)

2.2.3 Mesh Refinement Study

To obtain accurate results in the full-scale grounding

simulation, an element size of 5 mm, which also has been

used in the FE material tests, was applied to the bulb. To

reduce the number of elements needed, the mesh

resolution is finest within 500 mm of the contact area. In

greater distance of the contact, the mesh resolution is

gradually reduced (see Figure 2.7). Results of

simulations with element sizes of 5 mm, 10 mm, 20 mm

and 50 mm have been compared. The contact force of the

coarsest mesh is 3% greater, while the keel root force is

the same as in the simulation with 5 mm elements. An

overestimation of 3% is deemed acceptable, compared to

a 98% reduction of elements and the corresponding

increase in performance.

Figure 2.7: Mesh refinement of the bulb

3 GROUNDING SIMULATION

3.1 EXAMPLE OF BALTIC 175 “PINK GIN VI”

The example, chosen for the grounding simulation, is

175 feet (53.9 m) “Pink Gin VI”, that is being built at

Baltic Yachts, in Finland. The Baltic 175 is

representative for a modern super-yacht, that reaches

high speeds, thanks to a comparatively low displacement.

“Pink Gin VI” has a lifting keel that can reduce its draft

from 7 m to 4.5 m. The lifting keel is mounted in a keel

trunk, which extends to the deck. The trunk is supported

by a bulkhead, resulting in a high local stiffness, in the

keel area. It is assumed that the speed and structural

stiffness of the Baltic 175 can result in large loads in case

of grounding, which makes her a good example for this

study.

3.2 FINITE ELEMENT MODEL

The FE model of the Baltic 175 consists of one

deformable and four rigid bodies (Figure 3.1). The front

part of the bulb is defined as plastically deformable,

whereas the remaining part of the bulb, keel fin, hull and

the plate that resembles the obstacle are defined as rigid.

The rigid back of the bulb is simplified as a box. Mass

and inertia is assigned to the hull and keel parts.

Figure 3.1: Part definition for the grounding simulation


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3.2.1 Keel Fin and Hull Stiffness

It is expected that elastic deformation of keel fin and hull

prolongs the contact time during collision and therefore

reduces the peak force. Even though modelled as rigid

bodies, the elastic stiffness of keel fin and hull is

regarded by implementing a rotational spring, that

connects hull and keel. The combined rotational spring

stiffness of hull and keel is:

( 5 )

Hull stiffness is derived from a FE simulation

performed by GURIT, the structural designer of “Pink

Gin VI” [12]. In the simulation, the DNV GL grounding

loads were applied on the bulb tip of the keel, which was

modelled from overly stiff, solid elements. The

simulation also included rig forces in dock condition.

The hull stiffness is derived as deflection per force from

the simulation. Similarly, keel stiffness is derived from

an FE simulation, performed on behalf of APM, who

manufactured the keel [13]. As in the simulation

described above, the grounding force is applied to the

COG of the keel bulb. The keel fin is constrained by

contact elements at the bearings.

3.2.2 Hydrostatic Restoring Forces

Hydrostatic pressure is applied to the hull by using the

keyword *LOAD_SEGMENT_SET. The magnitude of

the pressure is determined by a user defined function,

using the keyword *DEFINE_FUNCTION. The function

calculates the hydrostatic pressure based on the

segment’s z-position and sets it to zero, if the z-position

is greater than zero. This is implemented as follows:

*DEFINE_FUNCTION

$# fid heading

5hpress

$# function

float hpress(float z, float z0)

{

float fac1, refz, rho, grav;

refz = 0.; rho = 1.025e-9; grav = 9810;

fac1 = 1.0;

if(z>refz) fac1 = 0.0;

return fac1*rho*grav*(refz-z);

}

4 RESULTS

Grounding simulations were performed at five different

initial velocities, ranging from to

, or an equivalent of Froude numbers from

to .

Figure 4.1 shows the contact and keel root force

progression for an initial velocity of . The

cross-section at the centreline of the bulb at four different

time-steps is presented below. At the beginning of the

collision, the reaction force of the plate progresses

steeply, up to a peak at (1). The bulb bounces

back from the plate, forcing the keel to rotate backwards,

reducing the contact force to ¼ of the first peak (2). The

stiffness of the torsional spring builds up resistance

against the backward rotation of the keel, which

transmits forces between the keel and the so far

unaffected hull, resulting in the maximum force (3). The

pitch-motion of the yacht rotates the keel back from the

plate, which ends the contact and causes the contact force

to drop to zero (4).

Figure 4.1: Contact force and keel root force progression

(top) and bulb deformation (bottom)

Up to its first peak (1), the force vs. time curve is

identical to that of a simulation excluding the hull.

Whereas the sole keel would rebound from the plate after

the first peak (2), a simulation involving the hull leads

the force vs. time curve to increase, again. Thus, the

second peak (3) in the force progression is a result of the

hull’s inertia. This is reflected in the keel root force,

which is zero during the first peak (1), but rises to its

maximum with the second peak (3).


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4.1 INFLUENCE OF HYDROSTATIC FORCES

Figure 4.2 shows a comparison between a simulation

with implemented hydrostatic forces and a simulation

without. The movement of the yacht over the course of

the collision is displayed above a graph with the contact

force and pitch angle progression. It stands out, that at

(3), when the force reaches its maximum, the

pitch angle is small, at 0.5°. The pitch angle progression

curves of the two simulations separate only when the

force is already declining, rendering the influence of

hydrostatic forces insignificant.

Figure 4.2: Yacht’s movement (top) at four different

time steps and contact force progression on primary axis,

with pitch angle on secondary axis, for a simulation with

and without hydrostatic pressure (bottom)

4.2 INITIAL VELOCITY

Figure 4.3 displays the maximum forces over initial

velocity, that occur over the course of the collision. The

upper curve resembles the maximum contact force,

acting between bulb and obstacle and the lower curve

being the keel root force, acting between keel fin and

hull. The forces increase linearly with initial velocity,

with the keel root force being about 22% lower than the

contact force.

Figure 4.3: Maximum contact force and keel root force

over initial velocity

4.3 ANTIMONY CONTENT IN LEAD ALLOY

Simulations with four different material curves were

performed, representing lead alloys containing 0%, 1.5%,

4% and 8% antimony. The simulations were performed

at an initial velocity of . As expected, higher

amounts of antimony result in greater stiffness and

consequently higher grounding forces. Figure 4.4 shows

the maximum contact and keel root force over the

antimony content. The force increase from pure lead to

an alloy with 8% antimony is 35%, while the ratio

between contact force and keel root remains constant.

Figure 4.4: Maximum Contact force and keel root force

versus antimony content of lead alloy


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4.4 STRUCTURAL STIFFNESS

The effect of the elasticity of keel fin and hull structure

in the keel area is assessed by varying the stiffness of

the rotational spring that connects hull and keel by

factors of 0.5 and 2. Figure 4.5 displays the progression

of the contact forces versus time for an initial velocity of

. A 50% lower stiffness reduces the

grounding force by 17%. A 100% greater stiffness

increases the grounding force by 26%. In both cases, the

keel root force is about 25% smaller than the contact

force.

Figure 4.5: Contact force progression for different

structural stiffness factors.

This shows that a lower structural stiffness prolongs

the collision time effectively and has a similar impact on

the grounding forces as the stiffness of the bulb material.

4.5 DRAFT

The draft of the Baltic 175 is varied by a factor of 0.5

and 1.5. The resulting maximum contact and keel root

force versus draft is shown in Figure 4.6. While the

maximum keel root force decreases uniformly with

greater drafts, the maximum contact force plateaus

between draft factor 1 and 1.5. This is due to a similar

force progression as in Figure 4.5, where the first peak in

the force progression is its maximum.

Figure 4.6: Contact force progression (top) and

maximum contact and keel root force (bottom) for

different drafts

4.6 BALLAST/DISPLACEMENT RATIO

The ratio between ballast and hull mass is varied while

keeping the displacement constant. Figure 4.7 shows the

maximum contact and keel root force over the

ballast/displacement ratio. An increase in

ballast/displacement ratio results in a larger contact force

and smaller keel root force.

Figure 4.7: Maximum contact and keel root force versus

ballast/displacement ratio.

4.7 CRASHWORTHY BULB

There are different concepts to reduce the grounding

force, that generally aim to prolong the collision time

while reducing the maximum force. One concept of keel

manufacturer APM is the implementation of vertical slits

into the bulb, that run longitudinally through the front

part. Those have the advantage of removing little

material from the bulb and therefore conserve its density,

while maintaining section modulus against bending

under own weight.


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For this study, a simulation with five longitudinal slits

has been performed. The slits extend 1.2 m backwards

from the tip of the bulb. The slits are 100 mm apart and

feature a width of 10 mm, reducing the total volume of

the bulb by 0.5%. The simulation has been performed at

an initial velocity of .

Figure 4.8 shows the deformation of the slitted bulb as

well as the force progression versus time. Compared to a

regular bulb, the first force peak is reduced by 22%,

whereas the second force peak is reduced by 12%. The

longitudinal force in the keel root is reduced by 14%,

compared to a regular bulb.

Figure 4.8: Deformation of a crashworthy bulb (left

column) and force progression in contact and keel root,

compared to a conventional bulb (top)

The arrangement of the slits has not been optimized,

thus it is assumed that the reduction in grounding force

can potentially be reduced further. In the present

simulation, the root of the outer webs is significantly

wider than that of the inner ones. Their higher resistance

against bending locks the inner webs, which results in

higher stiffness. Additional slits or a more even spacing

of the existing ones may lead to further reduction in

stiffness and grounding force.

5 CONCLUSIONS

A method of assessing the grounding forces experienced

by a sailing yacht has been successfully demonstrated. A

nonlinear FEA was performed to evaluate the influence

of plastic deformation of the keel bulb, while the

simplified implementation of the yacht’s structural

stiffness as a rotational spring reduced computational

effort. The investigation delivered valuable insights, such

as the negligible influence of hydrostatic restoring forces.

Material tests have shown, that lead features a more than

two times greater yield strength under compression, than

under tension. Further material tests that aim at

compression strength for different lead alloys are

recommended for future research projects.

The DNV GL rules predict a grounding load of 3.2 MN

for the Baltic 175. Based on the research of this paper, it

can be assumed that this force acts on the hull structure

when grounding occurs at a velocity of 4 knots. Since the

Baltic 175 is capable of much higher speeds, the actual

grounding loads might exceed the ones predicted by the

rules.

This paper presents an intermediate result on the

research currently done on horizontal grounding of

sailing yachts. A parametric study, which has not been

concluded at the time of completion of this paper,

investigates the grounding forces of yachts with different


INNOV'SAIL 2017 273

displacements. The aim is to find a general approach of

assessing horizontal grounding forces on sailing yachts.

ACKNOWLEDGEMENTS

The authors would like to acknowledge the help of Baltic

Yachts, for providing the geometry of the Baltic 175 and

the permission to use the yacht as an example for this

study. The authors also thank Schumacher Metall for

providing specimens for material testing. Further

contributions were made by APM-Keels, who provided

data about crash-save bulbs, GURIT, who provided FEA

results on the Baltic 175 and judel-vrolijk, who

contributed weight calculations.

REFERENCES

[1] APM, “Impact Simulation on Bulb,” Unpublished

internal document, 2012.

[2] J. Amdahl, S. Ehlers and B. J. Leira, Collision and

Grounding of Ships and Offshore Structures, 2013.

[3] L. Larsson, R. E. Eliasson and M. Orych,

Principles of Yacht Design, 2014.

[4] DNV GL, “Chapter 7 Rudder, foundations and

appendages,” in Rules for Classification - Yachts,

2016.

[5] A. J. Burton and R. A. Clegg, “Tetrahedral

Elements for Explicit Ballastics Simulations,” in

23rd International Symposium on Ballistics, 2007.

[6] Livermore Software Technology Corporation, LS-

DYNA Keword User's Manual, 2015.

[7] Livermore Software Technology Corporation,

Theory Manual, 2006.

[8] Vulcan Global Manufacturing Solutions, “Cast

Lead - Antimony Alloys Properties,” [Online].

Available: http://vulcangms.com/cast-lead-

antimony-alloys-properties/. [Accessed 2017].

[9] W. Hofmann, Blei und Bleilegierungen, Berlin,

1941.

[10] DIN Deutsches Institut für Normung e. V.,

“Prüfverfahren bei Raumtemperatur (ISO 6892-

1:2016),” in Metallische Werkstoffe - Zugversuch,

2016.

[11] DIN Deutsches Institut für Normung e.V.,

“Druckversuch bei Raumtemperatur (DIN 50106),”

in Prüfung metallischer Werkstoffe , 2016.

[12] GURIT, “Baltic 175 Longitudinal Grounding

Analysis,” Unpublished internal document, 2016.

[13] APM, “Structural assassment of the Baltic 175'

lifting keel,” Unpublished Internal Document, 2013.

AUTHORS BIOGRAPHY

H. Conradi is graduating as M.Sc. in Naval Architecture

and Ocean Engineering from the Hamburg University of

Technology (TUHH). He currently writes his master

thesis on the topic of grounding of sailing yachts at DNV

GL.

S. Ehlers, D.Sc., is a professor for design and analysis of

ships and offshore structures and the head of the institute

for ship structural design and analysis at the Hamburg

University of Technology (TUHH). He is an expert in

consequences assessment for accidental events and

further in the field of material modelling for non-linear

finite element simulations. Furthermore, he is developing

new ice material models to assess the ice-structure

interaction and design methods for ice going vessels. He

is concerned with the overall structural response and

strength of ships subjected to extreme conditions.

Additionally, he combines optimization techniques with

extensive assessment procedures to obtain new concepts.

H. Hoffmeister has been working for Classification

Society DNVGL (former Germanischer Lloyd) since

1993, as graduated Naval Architect. Since, he has been

involved in the assessment of marine structures, yachts

and rigs, developed several standards and guidelines,

amongst which are the GL Guidelines for Structural

Design of Racing Yachts. Thus, his particular field of

expertise is the evaluation of composite structures.

Parallel to his employment at Germanischer Lloyd,

Hasso Hoffmeister was member of the design team of

“United Internet Team Germany” for the Americas Cup

2007 as the team’s rig designer.


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