Schedule for Presentation Meeting of Master-Degree Thesis in 2019
International Course of Maritime & Urban Engineering
Date :
Venue :
Name

July 31st (Wednesday) starting from 09:30
S1-312 Lecture Room

No.

Time

Supervisor

Title of thesis

1

09:30 – 09:55

Win Naing Htay

Toda

A CFD-based Comparison Study of Conventional Rudder and Rudder with Bulb-FinsSystem of KVLCC2 in Waves

Sanada2

09:55 – 10:20

Aye Nyein Mon

Aoki

An Experimental Study on the Stability of Filter Units on Rubble Mound Breakwater

Sanada3

10:20 – 10:45

SAMDANI H M
GolamSanada4

10:50 – 11:15

Phyo Myat Kyaw

Osawa

5

11:15 – 11:40

HONG Yang

Kashiwagi

6

13:00 – 13:25

MANIYAPPAN
SreenathUmeda

7

13:25 – 13:50

WANG Yueyi

Irie

8

13:50 – 14:15

KALAMBAY
Ilunga PanickDoi

9

14:30 – 14:55

REN XuToda

10

14:55 – 15:20

YAO Yujia

SuzukiStudy on Dolphin Swimming Performance by CFD

Suzuki11

15:20 – 15:45

ABDUL HALIM
Azhan Bin

Osawa

Study on Measurement Method of Under-Film Corrosion’s Incubation Life Using Fe (II)
Fluorescent Indicator

Suzuki

16:00 –

at S1-323

Experimental Study on a Strengthening Technique by Wing Walls for Flat Plate-Column
Connections with Low-Strength Concrete
Development of an Influence Coefficient Database for Evaluating Mixed-mode SIFs andApplication to Inclined Surface Cracks in Semi-infinite Bodies
Numerical Calculation Method for Hydroelastic Response of a Ship with Forward Speedin Waves Using Orthogonal Mathematical Mode Functions
Lunch Break
Effectiveness and Mechanism of Broaching-To Prevention Using Global Optimal Control(CMA-ES)A numerical study on water environment prediction in Lake Kasumigaura using
unstructured grid three-dimensional flow model
Safety Performance Analysis of Unsignalized mid-block Crosswalks: An Agent-based
Approach for Pedestrian-vehicle Interactions
Break
Full Scale resistance prediction of paint surface -Proposal of new approximate formulafor friction with roughness

Meeting for Evaluating Mater-thesis Presentations, and Steering Committee Meeting

Meeting for Evaluation: From 16:00 at S1-323 Meeting Room
Members are Steering committee members and the supervisors

Chairman

SanadaIijimaIijima

Irie
Irie
Irie

SuzukiInternational Course of Maritime & Urban Engineering / Abstract of Master Thesis

July 26, 2019

A CFD-Based Comparison Study of Conventional Rudder and Rudder
with Bulb-Fins System of KVLCC2 in Waves
Win Naing Htay
Hull Form Design Sub-Area, Department of Naval Architecture and Ocean Engineering
Key Words: Energy-Saving Device, Rudder Bulb-Fins System (RBFS), CFD, Propulsion Performance, Head Waves

Introduction

Recently pressure with respect to greenhouse gas reduction on
ship-designers and yards is increasing considerably due toenvironmental concerns. As a result, International Maritime
Organization (IMO) published a new mandatory concept named
as Energy Efficiency Design Index (EEDI), in order to improveenergy efficiency and eliminate inefficient ships from the market.
An industry driven by eco-friendliness, which signifies thatdeveloping core technologies to cut greenhouse gas emission ofships will be increasingly essential. Climbing oil prices also raise
the necessity to develop Energy-Saving Devices (ESDs) which
minimize the energy loss during sail or partially recover the lostenergy. Rudder Bulb-Fins System (RBFS) is one of the ESDs tosolve this problems. These RBFS can recover rotational flow
behind the propeller.
The performance of RBFS in calm water by installing the finswith angle of attack on both sides and finding out the best angleof attack which gives the most efficient fin have been investigatedboth experimentally and by using Computational Fluid Dynamics(CFD) method1). It must be emphasized that design is aimed atgood performance not only for the model test or trial condition incalm seas but also for the service conditions with wind and waves.
Therefore, the objective of this research is to investigate theperformance of RBFS and to compare with conventional rudder
in different head waves conditions.

2.

Methodology

The prediction of self-propulsion factors and flow filed
parameters of KRISO Very Larger Crude Carrier (KVLCC2) oflength between perpendiculars (Lpp) 3.2m with Froude numberof 0.142 and two degree of freedom (heave and pitch) conditions
in regular head waves has been done by using CFDSHIP-IOWA2).
It is a general-purpose unsteady RANS or DES code coupling withmotion solver. The k-ω SST turbulence model was employed. A
single phase level-set method has been used to model the freesurface and two degree of freedom (DOF) motions are considered.
The grids comprising of 15 blocks for RBFS and 11 blocks forconventional rudder have been generated using Gridgen.
Thereafter SUGGAR library has been used for the overset of grids.
The computational domain and boundary condition is shown in
Fig. 1. A no slip boundary condition with u=0 and (v and w) are
specified based on propeller angular velocity; 2rhubn is imposedon the hub surface; rhub is the radius at the grid point on the hub
surface and n is the propeller revolution rate. u,v,w are the velocity
component in x,y,z direction. x,y,z are longitudinal direction,
starboard direction and vertical upward direction respectively. Allvariable are non-dimensionalized by ship speed U, ship Length L,
density  and their combinations.

Fig. 1 Computational domain and boundary conditions.

3.

Simulation Results

3.1 Ship motions
The time history comparison of heave and pitch motions
between RBFS and conventional rudder in wave of λ/L = 1.1 withamplitude of 3cm for one encounter period has been shown in
Figure. 2 together with waves at bow and propeller plane. The
time history of wave elevations at bow and propeller section arecalculated by the equation of ζ = A cos (kx – ωet). The amplitude
and phase of heave and pitch motion in RBFS are almost samewith that in conventional rudder. However, the average values ofship motions in RBFS seems to be more negative (larger sinkage
and bow down trim) as compared to conventional rudder.
Wave at bow

Wave at prop

Heave (RBFS)

Heave (Conventional)

Pitch (RBFS)

Pitch (Conventional)

3
wave (cm), heave (cm), pith (degree)

1.

2
1
0
0

0.25

0.5

0.75

-1
-2
-3

t/Te

Fig. 2 Comparison of ship motion between RBFS andconventional rudder in λ/L =1.1 for one encounter period.

1

3.2 Flow field analysis

3.3 Self-propulsion

Figure 3 plots the axial velocity contour (u/U) and cross flow
vectors (v, w) at x/L = 1, AP position across the rudder, bulb andfins, for λ/L = 1.1 at t/Te=0, wave crest at bow. The hub vortex
attaches on the portside surface of the rudder and a downward
flow is caused on the starboard side surface in both rudders. As
compared to conventional rudder, the hub vortex location islowered due to rudder bulb and fins and hub vortex area is smallerin RBFS.
To understand the vortex structure, Q-Criterion is used. In
figure 4, Q value of 500 is used for visualizing vortex for λ/L= 1.1
at t/Te = 0. The region where the selected Q value is matched, isosurface is created and colored with axial velocity contour (u/U).
The effect of rudder bulb and fins and the way of how they deformthe hub vortex are obvious in that figure. The propeller tip vortex
and hub vortex can be seen clearly and the hub vortex in port sideof rudder is stronger than the starboard side’s one in both rudders.
This is because of the clockwise rotation of propeller.
Additionally, the rudder fin tip vortex is occurred in the vortex
visualization of RBFS.

Table 1 shows the summery of mean values of thrust, wakefraction and hull efficiency between RBFS and conventionalrudder to know the improvement by the RBFS. Rudder bulb andfins can reduce the resistance because the shape of bulb and finsmakes the ship stern flow smooth and reduce the vorticity.
However, the thrust is lower slightly in RBFS only in λ/L= 1.6,
long wave conditions. In the result of effective wake, it is reduced
a little in all wave conditions except in λ/L = 1.6, long wave
condition. As KVLCC2 is tanker shape and the wake field is wide
so that the effective wake factor is lower in all conditions.
Regarding to that table, hull efficiency, ηH is improved by rudder
blub and fins in all wave conditions except long wave, λ/L = 1.6.
From these comparison, RBFS does not have very bad effect inwaves and some improvement in short and medium wave length.
Table 1 Summery of self-propulsion factors.

(a)

4.

(b)

Fig. 3 Axial velocity section in one encounter period for λ/L= 1.1
in x/L= 1 at t/Te=0 (a) conventional rudder (b) RBFS.

Conclusions

There is no significant difference in ship motions (heave andpitch) of between RBFS and conventional rudder.
In flow field analysis, the strength of hub vortex decreases dueto the rudder bulb. Its area and velocity around core is smaller in
RBFS than in normal rudder. The interaction of fin and ship
motion can be observed in waves.
As compared to the conventional rudder, the resistancedecreases on RBFS in all wave conditions. On the other hand,thrust is improved by rudder fins. Increase of hull efficiency anddecrease of open water efficiency are found in self-propulsion
analysis. Generally, overall quasi-propulsive efficiency improves
a little bit because of the rudder bulb and fins.
For the future work, it is necessary to validate the CFD resultswith experimental results. That’s why experiment will be
conducted using the KVLCC2 ship model with the same
conditions as CFD case in towing tank of Osaka University.

References
1)Fig. 4 Illustration of Q-criterion for Q=500 colored by the axial
velocity contours for λ/L=1.1 at t/Te =0(Left: conventional rudder, Right: RBFS).

2)

Truong, T. Q., Wu, P-C., Aoyagi, K., Koike, K., Akiyama,
Y. and Toda, Y. (2017) “The EFD and CFD Study of RudderBulb-Fin System in Ship and Propeller Wake Field of
KVLCC2 Tanker in Calm Water”,27th ISOPE conference,
California, USA.
Paterson, G., Wilson, W. and Stern, F.: "General-purpose
parallel Unsteady RANS Ship Hydrodynamics Code."
CFDSHIP-IOWA, IIHR Report No.432, pp.2-21, 2003.

International Course of Maritime & Urban Engineering / Abstract of Master Thesis

July 26, 2019

An Experimental Study on the Stability of Filter Units on Rubble
Mound Breakwater
Aye Nyein Mon
Coastal Engineering Laboratory, Department of Civil Engineering
Key Words: filter unit, FU damage, stability coefficient, Hudson formula1. Introduction

2. Outline of Hydraulic Model Experiments

Filter unit (FU) is a stone-filled net made of synthetic fiber.
The FU was used first in rivers in order to prevent scouring ofriver banks. Any type of solid materials such as stones, pebbles,
crushed concrete, etc. can be used as the filling material. The FU
can be placed on rougher surfaces because of its flexibility
compared with conventional materials1). Therefore, the FU has
recently been used also in coastal works.
When FUs are used in coastal engineering work, it has
identified that the filling stone movement generated by wave
action causes deformation and displacement of the FU. Further,
violent collision of the filling stones with the synthetic fiber
would damage the net. To overcome these problems, restraining
filling material need to be done which can not only increase thehydraulic stability but also durability of filter net. Therefore, arope restraining method was adopted1). The FU, that is tied with aconstraining rope from top to bottom, is referred to as the
restricted type (S-type FU) and the conventional unit that is
widely used for rivers is referred to as the normal type (N-typeFU). In addition, there is a newly developed FU by a JapaneseCompany, which is called “Rock Unit (RU)” (see fig. 1).

2.1 FU Models
Two groups of restricted type (S-type FU), by changing thevolume ratio of filter net to filling material and constraining rope
length and normal type (N-type FU) as well as Rock Unit (RU)
were used in the first experiment. All these FUs have the same
mass, 186.5 g, and also same volume (approximately) of fillingmaterial, but different volumes of filter net.
The particulars of S-type FU models for the single-FUexperiment, that are considered based on the FUs applied inseveral practical cases, are as shown in Table 1 & 2.
Table 1 S-type FUs changing in filter net size.
Case A
A1
A2
A3
A4
A5

Width x Length (mm x mm)
of Filter Net
200 x 120240 x 120
260 x120
280 x 120
320 x 120

Constraining Rope
Length (mm)
28
28
28
28
28

Table 2 S-type FUs changing in constraint degree.
Case B
B1
B2
B3

Fig. 1 (A) N-type, (B) S-type & (C) RU.
Previous studies highlighted that S-type FU is quite difficult to
deform compared with N-typed FU, and stacking arrangement of
S-type FUs is more stable than single layer layout 2). However,N-type FU can adapt perfectly into any space where theconstruction is required.
In this study, we have carried out two types of hydraulic modelexperiments: “single-FU” experiment and “multiple-FU”
experiment. In the former experiment, various types of FUs aretested one by one with the scale of 1/35 to study sliding
mechanism of each FU, and to compare stability of the FU. Basedon the experimental results, an optimum FU is selected for
multiple-FU experiment. The multiple-FU experiment (scale
1/40) is then carried out on the sloping breakwater in planararrangement in order to evaluate the stability coefficient of
optimum FU.

Width x Length (mm x mm)
of Filter Net
260 x120
260 x 120
260 x120

Constraining Rope
Length (mm)
18
28
38

2.2 Experimental Set-up and Test Conditions
All tests were conducted in a 30 m-long, 0.71 m-wide, and 1
m-deep wave flume of Osaka University. The tests in bothexperiments were started with small amplitude waves which did
not cause any movement of FU, and the wave height was
gradually increased until the FU show movement due to thestriking wave. The height of the striking wave was increased
gradually without rearranging the FU and without rebuilding the
rubble mound slope.
2.2.1 Single-FU Experiment
The FU movement was defined as a damage when a portion of
FU has moved in to the lee side of the rubble mound slope. Theseexperimental cases were conducted under regular waves on thesurface of fixed rubble mound. The test procedure was repeatedthree times to confirm critical wave height of each FU. The
experimental set-up in the wave flume is given in Fig. 2.

Critical Wave Height (cm)

9.5
9
8.5
8
7.5

Fig. 2 Experimental set-up of single-FU experiment.

7
6.5
6

2.2.2 Multiple-FU Experiment

5.5

In these test cases, FU damage was defined when a FU had
moved more than half of its diameter from its initial position. The
experimental set-up of these experiment cases are as shown in Fig.
3 and the test conditions are given in Table 3.

V of f ilter net/V of f illing stones

5
0

A1

A2

A3

A4

A5

Fig. 4 Comparison of critical wave heights in Case A.

Critical Wave Height (cm)

9.5
9

8.5
8
7.5
7

Fig. 3 Experimental set-up of multiple-FU experiment.

6.5
6Constraining Rope Le ngth (mm)Table 3 Test conditions of multiple-FU experiment.

5.5
8

18

28

38

Scale

1/40

Water depth

45 cm

Fig. 5 Comparison of critical wave heights in Case B.

Irregular Waves
Wave period T1/3
Wave height H1/3
Number of waves

Modified Bretschneider Spectrum
1 sec, 1.5 sec, 2 sec
6cm-10cm
1000

FU model: mass

125 g

The critical wave heights of N-type FU and the Rock Unit werelower than that of Case A3/B2. Thus it was selected as the
optimum FU for the next experiment.
The sliding mechanism of the Rock Unit followed a linear
motion, and N-type FU moved by deforming. Some of S-type
FUs turned over during the wave attack.

3. Hudson’s Formula
The stability of a particular shore protection unit can be
described by the Hudson’s Formula (1),

Where, KD and M are stability coefficient and mass of armor
unit; ρr and ρw are density of armor unit and density of water; H isdesign wave height; and ѳ is angle of structure slope measured
from horizontal in degrees. This formula was developed throughthe results of extensive small scale model testing and somepreliminary verifications of large scale model testing to determine
the stability of armor units on rubble structure.

4. Experimental Results

4.2 Multiple-FU Experiment
The optimum FU, selected from the first experiment, wassubjected to irregular waves in these experiment cases. The
number of FUs, which had moved more than half of its diameter,
was recorded with the significant wave height for all tests of wave
periods, 1 sec, 1.5 sec and 2 sec. The structure damaged severely
for waves with longer periods.

5. Conclusions
From the first experiment, it had found that the critical wave
height (stability) of S-type FUs depends on the volume ratio of
the filter net to filling stones as well as on the constraint degree,
and the Case A3/B2 is the most effective FU structure among all
FUs tested. The stability coefficient of the FU is quite large
compared with other armor units, which indicates that an FU issuitable in practical use from the viewpoint of structure stability.

4.1 Single-FU Experiment

References

The movement of all FUs was measured in both horizontal andvertical plane, and the critical wave heights that caused thedamage as per the definition were investigated. The critical wave
height results in both cases A & B are as shown in Fig. 4 & 5.

1) Shin-ichi Kubota et al, Stability and Durability of Filter Unit
for Protecting Rubble Mound of Composite Breakwater,
Coastal Structures 2007, Vol.2.
2) Yuki Saito, The Stability of Filter Units on a Slope ofComposite Breakwater, 2017

International Course of Maritime & Urban Engineering / Abstract of Master Thesis

July 26, 2019

Experimental Study on a Strengthening Technique by Wing Walls for
Flat Plate-Column Connections with Low-Strength Concrete
Samdani H M Golam
Concrete Structure Laboratory, Department of Architectural Engineering
Key Words: Brick chips, Half scale test, Punching shear failure, Retrofitting, Wing walls

1. Introduction
A reinforced concrete (RC) flat plate structural system is a
structurally economic and feasible system. However, it is prone toshow brittle punching shear failure at the flat plate-column
connections, which may lead in the progressive collapse of abuilding. This study focuses on the strengthening of flat platecolumn connections in existing structures made with low-strength
concrete built in Bangladesh. Because of unexpected increase inapplied loads, lack of consideration of seismic effects, andinappropriate management during design and/or construction, asignificant number of existing flat plate structures are currently
required to be strengthened against punching shear failure.
Therefore, the seismic evaluation and retrofit methods arenecessary to be developed. This study proposes a newstrengthening technique using RC wing walls and describes a seriesof experimental investigations on the behavior of half-scaled flat
plate-column connections.

direction for specimen FP-W2.
The specimen details are shown in Fig. 1. The plate dimensions
were 1500 x 1100 x 75 mm in length x width x thickness,
respectively. The dimensions of wing walls were 175 x 100 mm inlength x thickness, respectively. Moreover, two steel plates (PL-6
x 80, SS400) were embedded in both sides of the slab, mainly torealize the punching shear failure. Thus, the design ultimate
flexural strengths of the plate and column were sufficiently larger
than the punching shear strength. The material properties ofconcrete and reinforcements are shown in Table 1 and Table 2,
respectively. A cement and CaCO3: sand: brick chips volumetricratio was designed as 1 (0.7+0.3): 2: 4, where 30% of cement was
replaced by CaCO3 mainly to reduce the concrete strength. A ratioof water to the sum of cement and CaCO3 was 60%.
Table 1 Material properties of concreteType

2. Experimental Program

FP-1
FP-W1
FP-W2
Wing walls

2.1 Material properties
Experimental research on a series of three half-scaled interiorslab-column connections was conducted mainly to investigate theeffectiveness of the proposed retrofit technique using wing walls
and the punching shear capacity of the connections with low
strength concrete. Two specimens, FP-W1 and FP-W2 wereretrofitted by wing walls. Another specimen (specimen FP-11)) wastested without wing walls and served as a control specimen to makecomparison between the cases with and without retrofitting.
Wing walls were attached along the longitudinal direction for
specimen FP-W1, while the direction was along with the transverse
1500
175

487.5

662.5

175

Elastic
modulus(N/mm2)
8,450
8,672
8,741
30,929Strain atcompressive
strength (μ)
1,829
1,694
1,586
1,956

Table 2 Material properties of reinforcements

D6 (SD295A)
D16 (SD345)
1500
175

175

487.5

Yield
stress(N/mm2)381
369

Elastic
modulus(N/mm2)
191,000
175,000
1500
175

662.5

Tensilestrength(N/mm2)
505
536

662.5

75

662.5

Compressive
strength(N/mm2)
6.93
7.16
7.20
37.04

Column mainbars 8-D16

Slab Reinforcement
2-D6@75(1.12%)
Hoop D6@50(0.72%)

Insert pipe

Horizontal bars
D6@150Wing-walls verticalbars 4-D6(0.70%)

Bottom stub mainbars 10 D-16
Hoop D-6

Steel plate

Strain gauges

Strain gauges

Strain gauges

175

175

100

175

175

Critical Section(a) Specimen FP-1

175

175

175

175

100

(b) Specimen FP-W1
Fig. 1 Details of specimens (Unit mm)(c) Specimen FP-W2

Mo M f Ms Mt

2.2 Loading and measurement method
Fig. 2 shows an experiment set-up. The specimens were
subjected to static cyclic vertical loads by two hydraulic jacks. The
loading was controlled by a drift angle, which was the ratio ofvertical displacement from the application point of load to thecenter of column.

where symbols and details of the above equations (1) and (2)can be referred to the reference2,3).
Table 3 Experimental and calculated strengths
Strength
Mexp (kNm)
My (kNm)Mf (kNm)Ms (kNm)
Mt (kNm)
Mo (kNm)

Max. capacity
Flexural strength
Positive

Negative

Negative

Positive

Punching shear
strength

Hy draulic Jack

Colum nChannel Section
125x65x12x21

(2)

FP-1

FP-W1

FP-W2

7.3
25.1
3.56
2.71
3.66
9.93

19.9
25.1
3.48
6.9
9.84
20.2

14.7
25.1
8.72
6.92
3.73
19.4

The maximum capacity of FP-W1 was 2% lower than thecalculated result of the punching shear strength, whereas 27%
lower for FP-1 and 25% lower for FP-W2. In particular, theequations for punching shear strength provided by the AIJ
standards overestimated for all specimens, which might be caused
by application of low-strength concrete.

Flat Plate

Bottom Stub

Fig. 2 Schematic view of experiment set-up

5.

The nodal moment vs vertical drift angle relationships of thespecimens are shown in Fig. 3. Specimen FP-1 was tested as acontrol specimen. Shear cracks extended radially from the column
and the surface concrete peeled up due to punching shear failure.
Punching was fully visible at R of 3.34 x10-2 rad.

From the experimental results explained above, the following
conclusions were obtained:
1) The maximum strengths of both strengthened specimens weresignificantly increased compared to that of the control
specimen. Moreover, the ductility of the specimen FP-W2 wasNodal moment (kNm)

3. Experimental Results and Discussion

Specimen FP-1

20
10

Mmax =+7.29 kNm-2

Specimen FP-W1

20
9.93

10

Rmax=2.35x10 rad

0Conclusions

20.2

Mmax =+19.9 kNm-2

Rmax=1.35x10 rad

0

-10

9.93

-20
-9

-6

moment FP-1
using AIJ std.
initial crack
max. strength
punching failure

-3
0
3
-2
Drift angle (x10 rad)6

Specimen FP-W2

10

Mmax =+14.7 kNm

19.4

-2

Rmax=4.82x10 rad

0moment FP-W1
using AIJ std.
initial crack
max. strength
punching failure

-10
-20

9

20

20.2

-9

-6

-3
0
3
-2
Drift angle (x10 rad)

6

-10
-20

9

19.4

-9

-6

moment FP-W2
using AIJ std.
initial crack
max. strength
punching failure

-3
0
3
-2
Drift angle (x10 rad)6

9

Fig. 3 Nodal moment vs drift angle relationshipSpecimen FP-W1 showed higher strength but did not experiencehigher deformation capacity. The strength decreased with raising
up of the surface concrete due to punching shear failure. The
punching shear failure was visible at the peak loads of both positive
and negative loading at R of 1.35 x10-2 radSpecimen FP-W2 showed higher strength and deformation
capacity than that of the control specimen. The strength decreased
with the increase of drift after the peak load. Two slab
reinforcement in the connection between slab and wing walls were
yielded at peak load. Punching shear failure was fully visible at Rof -6.0 x10-2 rad on the bottom surface of the slab.

4. Ultimate Strength Evaluation
The experimental ultimate strengths are compared to the design
calculations, as shown in Table 3. The flexural strength of the plate
was calculated by the following equation3):

My 0.9at1σy1d10.9at2σy2d2(1)

The punching shear strength M0 was calculated according to thefollowing equations by the AIJ standards for RC structure3):

2)

3)
4)

larger than that of the specimen FP-W1.
Punching shear failure occurred in the specimens. Shear
cracks extended radially from the column or wing walls
surfaces and the slab surface concrete peeled up due topunching shear failure.
The punching shear strengths calculated based on the AIJ
standards overestimated for all specimens.
It was experimentally verified that installing RC wing walls
was a feasible method of upgrading flat plate-column
connections under seismic loading conditions.

References
1) H M Golam Samdani, Seolmi Kim, Susumu Takahashi,
Suguru Suzuki, Yasushi Sanada: AIJ Annual Convention,
Vol.4, pp.445-446, 2018.
2) Seolmi Kim, H M Golam Samdani, Susumu Takahashi,
Suguru Suzuki, Yasushi Sanada: AIJ Annual Convention,
Vol.4, pp.447-448, 2018.
3) Architectural Institute of Japan: AIJ Standard for Structural
Calculation of Reinforced Concrete Structures revised 2010,2014.9 (in Japanese).

International Course of Maritime & Urban Engineering / Abstract of Master Thesis

July 26, 2019

Development of an Influence Coefficient Database for Evaluating
Mixed-mode SIFs and Application to Inclined Surface Cracks in Semiinfinite Bodies
PHYO MYAT KYAW
Ocean Material Engineering Lab., Naval Architecture and Ocean Engineering Department
Key Words: influence function method, mixed-mode SIFs, crack face traction, inclined surface cracks, FEM

1. Introduction
Accurate estimation of stress intensity factors (SIFs) is a criticalfactor in improving the reliability of evaluated fatigue crack
propagation rates for ship and offshore structures. By using theinfluence function method (IFM)1), SIFs for cracks under arbitrarystress fields can be easily calculated. Previous studies of IFM
described the evaluation of mode-I SIFs and mixed-mode SIFs
(MM-SIFs) 1)-4) without crack face traction (CFT) which can
improve the accuracy of SIF solutions.
In this study, an effective technique is proposed to compute MMSIFs for inclined surface cracks in semi-infinite bodies underarbitrary stress fields by means of IFM considering the CFT5)-7).
The accuracy and effectiveness of the developed influence
coefficient database (ICDB) and proposed technique is verified
with well-established reference solutions7)-9). The effect ofPoisson’s ratio on MM-SIFs is also discussed based on theevaluated MM-SIF solutions.

2. Theoretical background
The IFM proposed by Shiratori et al.1) is based on superposition
principle and finite element analysis (FEA). The six componentsof stresses, 𝜎𝑖𝑗,𝑃 , at P-th node which is located on the crack faceof the non-cracked medium are evaluated under given externalloading by means of FEA.
𝑖𝑗,𝑃𝑄
𝑖𝑗,𝑃𝑄
MM-SIFs at the crack front node, Q-th node, 𝐾𝐼, 𝐾𝐼𝐼 and𝑖𝑗,𝑃𝑄
𝐾𝐼𝐼𝐼 , due to unit distributed load (UDL) at the P-th node are thenevaluated. The interaction integral method (IIM) considering the
CFT integral5)-7) is employed in the evaluation of MM-SIFs due tothe UDL. The IIM given in WARP3D code is modified so that six
components of traction stresses can be applied at each node on theelement face of the crack face. The calculated MM-SIFs, are called
influence coefficients (IC). The values of IC for each node on thecrack face are then saved to a database, called ICDB.
According to the IFM, MM-SIFs at the Q-th node on the crack
𝑄
𝑄
𝑄
front, 𝐾𝐼 , 𝐾𝐼𝐼 and 𝐾𝐼𝐼𝐼 , can be calculated as in Eq. (1).
𝑖𝑗,𝑃𝑄

𝑄

𝐾𝐼 = ∑ 𝐾𝐼
𝑃

𝑖𝑗,𝑃𝑄

𝑄

𝐾𝐼𝐼 = ∑ 𝐾𝐼𝐼
𝑃

𝜎𝑖𝑗,𝑃

𝑖𝑗,𝑃𝑄

𝑄

{

𝜎𝑖𝑗,𝑃

𝐾𝐼𝐼𝐼 = ∑ 𝐾𝐼𝐼𝐼
𝑃

(1)

𝜎𝑖𝑗,𝑃

3. Proposed Technique
By Saint-Venant’s principle (SVP), when a body is large enough,the MM-SIFs for a non-inclined crack subject to this multiaxial
CFT stresses are nearly equal to those of the inclined crack subject

to the remote uniaxial stress. These lead to an assumption that the
MM-SIFs of a non-inclined crack in a large body subject to themultiaxial CFT stresses, which occur on the inclined crack faceunder the uniaxial loading, should be the same as those of an
inclined crack in a semi-infinite body subject to the remote uniaxial
stress. Based on this assumption, a technique is proposed to
calculate MM-SIFs accurately for surface cracks in semi-infinite
bodies subject to a remote stress with different crack orientationsin short time and with no effort for generating many finite element(FE) models.
The traction stresses induced on the crack face of the inclined
crack model can produce MM-SIFs. The applied traction stresseson the inclined crack face were transformed in accordance with thecrack angle.
The traction stresses for any crack angle can be readilytransformed, for example, by using the Microsoft spread worksheet.
By employing the proposed technique, only one FE model with a
non-inclined crack can be used for calculating MM-SIFs for any
crack angle based on the transformed traction stresses. Based onthat, MM-SIFs for cracks with different angles in semi-infinite
bodies and subject to arbitrary stress distributions can be readily
calculated using a non-inclined cracked model by means of the
proposed technique. This technique efficiently reduces the manhours needed to generate FE model and traction stresses as well as
it accurately evaluates MM-SIFs for different crack angles.

4. Numerical Examples
Four different numerical analyses are conducted for this study.
The first two analyses validate the adequacy of the IFM based MMSIF calculation system which is to be employed in the proposed
technique. The proposed technique is then validated. Finally, MMSIFs for different crack angles and Poisson’s ratios are evaluated.
4.1 Model definition
The target FE models used in the numerical examples are chosen
to be large enough so that they can be treated as the semi-infinite
bodies based on the assumption of SVP. The semi-infinite bodies
are needed to be employed in the proposed technique so thatdifferent load directions and crack angles can be included in theanalyses. In the validation of ICDB for semi-infinite bodies, theflat plate surface-cracked models which are subjected to external
uni-axial tensile loading for mode-I case and pure shear loading formixed-mode case are employed.
To validate the proposed technique described in section 3, semiinfinite flat plate models with inclined and non-inclined surface
cracked models are used. And then, by means of proposed
technique, MM-SIFs for different crack angles are evaluated byusing the non-inclined cracked model. Only one FE cracked modelwith a non-inclined crack is employed for all the applied crack
angles. Fig. 1 shows the geometric configuration of one of the flat
plate surface-cracked models used in this study.

0.4
0.35

FII

0.3
0.25
0.2
0.15
0.1

0.05
0
-0.05 0

10

20

10

20

30

40

50

60

70

80

90

30

40

50

60

70

80

90

Location, φ (deg.)
(b)

0.3
0.25

FIII

0.2

Fig. 1 A semi-circular surface crack in a semi-infinite flat platebody. (a) geometry of the crack face and (b) geometry of thecracked body.

0.1
0.05

4.2 Results and discussion
The calculated SIFs for the above-mentioned FE models arecompared with well-established reference solutions7)-9) where an
excellent agreement was achieved. Fig. 2 shows the calculated
MM-SIFs for different crack angles by means of the proposed
technique in comparison with the reference solutions9). The
solutions obtained by means of the IFM are described as ‘IFM’.
The reference analytical solutions9) are denoted as ‘Noda’.
𝑄
The calculated MM-SIFs are normalized as: 𝐹𝐼 = 𝐾𝐼 ⁄𝜎√𝜋𝑎,
𝑄
𝑄
𝐹𝐼𝐼 = 𝐾𝐼𝐼 ⁄𝜎√𝜋𝑎 and 𝐹𝐼𝐼𝐼 = 𝐾𝐼𝐼𝐼 ⁄𝜎√𝜋𝑎 , where 𝐹𝐼 , 𝐹𝐼𝐼 and𝑄
𝑄
𝐹𝐼𝐼𝐼 denote normalized mode-I,-II and -III SIFs while 𝐾𝐼 , 𝐾𝐼𝐼
𝑄
and 𝐾𝐼𝐼𝐼 are calculated mode-I,-II and -III SIFs. The locations
along the crack front are represented by φ (in degrees) in which φ= 0° denotes the crack mouth and φ = 90° represents the crack
deepest point.
0.75
0.65
0.55
0.45

FI

0.15

0.35

0
0

-0.05

Location, φ (deg.)(c)Fig. 2 Validation of MM-SIF solutions for different crack angles,
a/c = 1.0, ν = 0.3. (a) 𝐹𝐼 , (b) 𝐹𝐼𝐼 and (c) 𝐹𝐼𝐼𝐼 .

5. Conclusions
The following facts can be concluded for the present study.
1)ICDB of cracks in semi-infinite bodies is constructed andvalidated by using FE models of cracked bodies which are large
enough compared to the crack size based on the assumption of
SVP.
2)The new proposed technique works well to evaluate the MMSIFs for different crack angle with single FE model. It also
reveals the effectiveness and handiness for numerical analyses ofdifferent crack angles without consuming time for FE modelling.
3)In this study, wide range of MM-SIFs for different aspect ratios
and Poisson’s ratios which are not described in previous studies
is also calculated and presented.

0.25

References

0.15

1)

0.05
-0.05 0

10

20

30

40

50

60

70

80

Noda_θ=0°
Noda_θ=45°
Noda_θ=90°
IFM_θ=30°
IFM_θ=60°

Noda_θ=30°
Noda_θ=60°
IFM_θ=0°
IFM_θ=45°
IFM_θ=90°

(a)

90

2)

Location, φ (deg.)

3)
4)
5)
6)
7)
8)
9)

Shiratori M et al.: T of JSME, Vol.52, pg. 390–8, 1986. (inJapanese)
Shiratori M et al.: T of JSME, Vol.56, pg. 265–71, 1990. (inJapanese)
Shiratori M et al.: ASME, Vol. 6, pg. 929–39, 2011.
Osawa N et al.: J of JASNAOE, Vol.12, pg. 201–8, 2010. (inJapanese)
Gadallah R et al.: J of Ocean Eng., Vol.138, pg. 123–39, 2017.
Gadallah R et al.: J of EFM, Vol.197, pg. 48-65, 2018.
Gadallah R et al.: J of ASME, Vol.3, 2016,
Noda N et al.: J of IJF, Vol.75, pg. 19–48, 1996.
Noda N et al.: J of IJF, Vol.127, pg. 167–91, 2004.

International Course of Maritime & Urban Engineering / Abstract of Master Thesis

July 26, 2019

Numerical Calculation Method for
Hydroelastic Response of a Ship with Forward Speed in Waves
Using Orthogonal Mathematical Mode Function
Hong Yang

Ocean Space Development Laboratory, Department of Naval Architecture and Ocean Engineering
Key Words: Hydroelasticity, Rankine panel method, Mode functions, Orthogonal polynomials, Frequency domain

1. Introduction
In recent years, larger merchant ships are built due to theincrease in ocean transport. As the size of ships increases, ships
become more flexible and associated accident is liable to occur.
Therefore, to compute wave-induced hydroelastic responses with
sufficient accuracy becomes of engineering importance. In
particular, wave-induced vibrations, referred to as springing andwhipping, can aggravate the fatigue damage of very large ships.
In structural analysis, it is common to define a special set of
natural mode shapes, which correspond to the actual flexural
deflections of the body in a specified physical context. Sometimes
this may be difficult in hydroelastic problems because the modeshapes must satisfy the edge boundary conditions and they areaffected by the hydrodynamic pressure field and hence cannot be
specified in advance. This difficulty can be avoided if thestructural deflection is represented instead by a superposition of
simpler orthogonal mathematical mode shapes which are
sufficiently general and complete to represent the physicalmotion.
In the present study, the structural deflections of a floating
slender ship can be expressed by the free undamped ‘wet’
bending modes of the hull in the water, by the ‘dry’ modes of thesame structure in air, or more simply the orthogonal modes of a
uniform beam. An even simpler representation can be developedin terms of orthogonal polynomial mathematical functions,despite the fact that these lack a physical basis and do not satisfy
the appropriate free-edge boundary conditions1). This research is
directed to prove that a superposition of orthogonal mathematical
functions can represent the flexural deflection of a ship in waves,despite of the required boundary conditions, and the advantages
and disadvantages of orthogonal mathematical functions are
discussed.

Fig. 1 Mode shapes of Euler-Bernoulli’s beam and Legendre
polynomials in vertical beading.
2.2 Response function
The response function given from beam equation2) is shown inEq.(2).
(2)

2.3 Mass matrix and stiffness matrix
The mass matrix and stiffness matrix in above equation canbe calculated as(3)(4)

and both matrices are symmetric matrices.

3. Computational Results

2. Numerical Calculation Method

3.1 Hydrodynamic analysis based on frequency domain RPM

2.1 Mode functions
There are two mode functions applied in this research topresent the deformation of the elastic body. The first is naturalmodes using Euler-Bernoulli’s free-free beam, which satisfies
boundary conditions and is used commonly. The another isLegendre polynomial modes despite the fact that lack a physicalbasis. An especially compact expression for the Legendre
polynomials is given by Rodrigue’s formula:(1)

There two mode functions are both orthogonal, and theillustration of mode shapes of these modes are shown as Fig. 1.

Fig. 2 Added mass and damping coefficient of the firstbending mode.
Fig.2 shows the computed results of the non-dimensional

added mass and damping coefficient. The differences of thesebetween alternative mode functions can be noticed but thetendencies is at the same pace.

larger than natural modes.
In this research, it is the total elastic deformation rather than
the amplitude of each bending mode that should be forced on. For
find out the convergence numbers of each mode we need, theseveral values are defined as Eq.(5).
(5)

The results of j from 1 to 4 are illustrated in Fig.5

Fig.3 Exciting force of the first and second bending mode.
The tendency of amplitude of each bending mode is at samepace but the phase could be almost same from Fig.3.
The differences between two mode functions exist surely
because the magnitude of natural modes is slightly larger than
Legendre polynomial modes shown in Fig.1.
3.3 Amplitudes of ship motions

Fig. 5 Ship motion summations of several bending mode.
From Fig.5, we can find good agreement is observed as well inthis comparison as long as enough bending modes added, despiteof the difference existed in hydrodynamic force and lacking
physical meanings.

4. Conclusions
In this paper, the hydrodynamic force is computed using
different mode functions, and there are must difference
existed.due to different magnitudes of mode functions.
The Legendre polynomial modes can replace of nature modes
using at least five bending modes for precision and in theengineering filed with more complex boundary conditions,
Legendre polynomial modes are much more easier without
considering any boundary conditions as long as enough bending
modes.

References
1) Newman J N. Wave effects on deformable bodies[J]. Applied
ocean research, 1994, 16(1): 47-59.
2) Bishop R E D, Bishop R E D, Price W G. Hydroelasticity ofships[M]. Cambridge University Press, 197

Fig.4 Ship motions of the first five bending modes.
Conversely, in Fig.4, the ship motions of the first fivebending modes using Legendre polynomial modes are much

International Course of Maritime & Urban Engineering/ Abstract of Master Thesis

July 23, 2019

Effectiveness and Mechanism of Broaching-To Prevention
Using Global Optimal Control (CMA-ES)
Sreenath Maniyappan
Ship Design Subarea, Department of Naval Architecture and Ocean Engineering
Key Words: optimal control, broaching, fixed point stability, unstable invariant manifold

1.

Introduction

Broaching-to is a phenomenon in which a ship cannot
maintain a constant course despite maximum steering
effort. Most experiments and numerical simulations carriedout related to broaching, engage PD control. However, inorder to realize the concept of the maximum steering effort
in broaching, the control system can be improved using the
optimal control strategies.
The first effort in this direction using optimal control wasdone by Maki et al.3) in 2008, using the variational method
based algorithm known as SCGRA (Sequential Conjugate
Gradient-Restoration Algorithm) method and SQP(Sequential quadratic programming) algorithm. However,these techniques were not sufficiently powerful to be usedin rugged search landscape. Further, in 2018 Maki et al. 1)
carried out a numerical simulation using a powerful
optimization algorithm known as CMA-ES (CovarianceMatrix Adaptation-Evolutional strategy)4). Both theseattempts showed the capability of optimal control inpreventing the broaching phenomenon but physicalexplanation of this mechanism remains to be investigated.
Hence, in this research we use CMA-ES for optimal
rudder control, with an improved rudder model in a 4DoF
Surge-Sway-Yaw-Roll maneuvering model. Further,simulate broaching conditions and nonlinear analysis on thesystem carried out. Subsequently, the effectiveness of
optimal control over a range of Froude number andautopilot course are investigated.
2.

Methodology

2.1

CMA-ES

(

)

(

(

)

)

(

)

(3)

Here 𝑓1 𝑡𝑜 𝑓8 are obtained from the equations of ship
motions in waves.
2.3

Optimal control formulation

The problem is treated as an optimal control case, having
specified initial condition and free final condition, with anobjective function J to be minimized.

(

)

(4)

here K, is a penalty value to prevent capsizing.
Optimal control is realized in two different ways as
follows:
Method I (Optimal 𝜹̇ CMAES): the state equation (3) is
modified by replacing the rudder differential equation withthe optimal solutions generated by CMA-ES. This meansthat steering gear dynamics are ignored.
Method II (Optimal Rudder gains CMAES-PD): The
state equation for steering gear remains the same with therudder proportional and differential gains optimized.
2.4

Stability Analysis
̅̅̅̅𝜉Firstly, the fixed points 𝐱̅ = { 𝐺 , 𝑢̅, 𝑣̅ , 𝜒,
̅ 𝑟,
̅ 𝜙̅, 𝑝̅ , 𝛿 ̅} are
𝜆obtained by solving the following equation:

It is a probabilistic model-based and comparison-basedalgorithm, which does not require differentiability of theobjective function that is suitable for simulation-based
optimization.
2.2

where, 𝜉𝐺 is the longitudinal position of CoG of thevessel from a wave trough. The dynamical system incase of PD control can be represented by the following
state equation.

(

Then evaluating the eigenvalues of Jacobian matrix

Maneuvering Simulation model

The subject ship is ITTC-A2 a 135 GT purse seiner. I use
the 4 DoF surge-sway-yaw-roll model2) having the wave
trough fixed and ship CoG fixed coordinate systems arebeing used. The system could be represented using the statevector, x, and parameter vector, p, as follows
(1)
(2)

(5)

)

(

)

(

)

(6)

If an eigenvalue of equation (6) has a positive real part,
local asymptotic behavior at 𝐱̅ is unstable and if the real partof all eigenvalues are negative then local behavior is stable.
3.

Results

Fig.1 shows the success region for the two different
optimal control methods adopted in this research.

Fig. 4 Rudder histories during broaching time
Fig.1 Effectiveness of optimal control in
avoiding broaching/capsizing

3.1

Mechanism Type 1

3.2

Mechanism Type 2

In this case the system instability is weakened by the large
differential gain in optimal PD and causes the system toremain in surf-riding for longer duration.

Here the conventional PD control results in broachingassociated with surf-riding, the optimal control and optimal
PD control does in periodic motion as shown in Fig. 2. The
slightly larger rudder action around the equilibrium point
prevents the ship to touch the unstable equilibrium and the
trajectory undergoes a qualitative change, as can be explainedusing Fig. 3&4.

Fig. 5 Eigen value of unstable fixed point under optimal gain

4.

Fig. 2 Phase portrait of surge motion

Two different mechanisms how optimal control avoids
broaching are identified. This information could be vital in
avoiding dangerous scenarios. The effectiveness of optimal
control methods and the dominant mechanisms under varying
Froude number and heading angle is also brought out.
5.

Reference

1)

Maki, A., Banno, Y, Umeda, N., Sakamoto, N., and Akimoto,Y.: On Broaching-to Phenomenon Using Optimal Control
Theory with Evolutional Strategy, Conference proceedings,
The Japan Society of Naval Architects and Ocean Engineers,
Vol.26, pp.501-506, 2018.
Umeda, N.: Nonlinear Dynamics of Ship Capsizing due to
Broaching in Following and Quartering seas, Journal of Marine
Science and Technology, Vol 4, pp.16-26, 1999
Maki, A., Umeda, N. and Ueno, S., Investigation on Broachingto Using optimal control theory, Journal of the Japan Society ofNaval Architects and Ocean Engineers (In Japanese), Vol.8,
pp.115-122, 2008
Hansen, N.: The CMA Evolution Strategy: A Tutorial. ArXiv
e-prints, arXiv:1604.00772v1, 2016

2)

3)

Fig. 3 Unstable Invariant Manifold of the
fixed point near wave crest

Conclusion

4)

International Course of Maritime & Urban Engineering / Abstract of Master Thesis

July27 ,2019

A numerical study on water environment prediction in L. Kasumigaurausing unstructured grid three-dimensional flow model
Wang Yueyi
Water Engineering Laboratory, Department of Civil Engineering
Key Words: Unstructured grid, wind driven current, dissolved oxygen, diurnal temperature stratification

1. Introduction
Due to temperature difference between inflow river waterand lake and the existing of temperature stratification in lakes,
inflow rivers can trigger density currents. The different
movement patterns of density currents in lakes will directly
affect the movement and distribution of nutrients, pollutants orsediment particles, and water quality and ecologicalenvironment，then finally has a significant influence on theformation of hypoxia. Hypoxia is not only killing the fishes andinvertebrates, but also impacted as an endocrine disruptor andpresents a serious danger to the reproduction, hence
sustainability of fish populations1). Therefore, it is important tostudy the characteristics of currents and tempo-spatial change ofdensity in temperature-stratified water for grasping theformation, development and fate processes of surface algae andbottom hypoxia in temperature stratified lakes.
It is very difficult to accurately grasp the flow of lakes only
by analyzing the observed data. Consequently, the numericalsimulation method should be adopted. Numerical simulation can
provide more comprehensive and coherent understanding of theproperties of current in the whore area, which makes it possible
to precisely analyze phenomena like temperature stratification,
hypoxia and so on. The objectives of this study are to verify theapplicability of SUNTANS, a recently developed coastal flow
model, in Lake Kasumigaura, and then simulate the influence of
external forces on variations of current to reveal formation anddispersion characteristics of hypoxia and fate of surface algae.
Particle tracking also been used to show the wind driven current
on the surface of lake.

2.

Study Area and Observations

2.1 Study Area
Lake Kasumigaura is located in the east-central part ofJapan, about 60 km from Tokyo, as shown in Figure 1. Lake
Kasumigaura refers to a group of adjacent lakes, can be roughly
divided into three parts: Nishiura (西浦;167.63km2), Kitaura(北浦; 35.16 km2), and Sotonasakaura (外浪逆浦; 5.85 km2), as
well as rivers connecting them. The total area is approximately
220 km2, which is second largest in Japan. The average depthof the lake is quite small for the large area.

Fig. 1 Location of Lake Kasumigaura and observation stationsin the lake area

2.2 Observation DataIn Figure 1, red circles marks show the location ofobservation stations for water quality related data, while thegreen polygons indicate the location of observation stations formeteorological related data. From observation data, the annual
average COD fluctuates around 7mg/L, which is twice as highas environmental standards. At the same time, the concentrationof total nitrogen and total phosphorus is about 1.3 mg/L and 0.1
mg/L respectively. That is to say, it exceeds the limit value ofeutrophication.

3.

Calculation Methods

3.1 Flow Model
SUNTANS is an unstructured grid, non-hydrostatic, parallel
computing coastal ocean model, created by Fringer2). It has beendeveloped for non-hydrostatic simulation of flows at highresolution in estuaries and coastal sea3) 4), while most application
uses with hydrostatic condition. Compared with widely usedstructural grids, unstructured grids have better geometric fitness.
This is an important advantage for lakes with complex planar
shapes like Lake Kasumigaura. On the other hand, SUNTANS
was designed to use Cartesian coordinate system, which could
avoid horizontal gradient errors which might be sometimes
occurred in sigma coordinate system. With these superiorities,
SUNTANS has been chosen as a powerful tool to simulate andalso used in this study to analyze the feature of Lake
Kasumigaura. SUNTANS uses Boussinesq approximation andincompressibility assumption to solve the Reynolds average
Navier-Stokes equation. The horizontal and vertical eddy
viscosities are given by Smagorinsky model and MellorYamada turbulent closure model, respectively.
3.2 Model Setup
This study followed the approach of Guillaume and Wells5)
who studied the influence of a typhoon on gyre generation in
Lake Biwa, Japan. As for the grid, SMS, a software to generate
grids, was also used. Table 1 shows model setup. In order tounderstand the effects of various external forces, several
calculation cases were run with different idealized forcing
conditions, as shown in Table 2.
Table 1 Calculation condition
Horizontal Average Interval: 200m
Grid size
Vertical: 20 layers
Calculation period
2010/9/1-2010/9/31
Precipitation, atmospheric temperature,
Input datawind speed and direction, short wave, airpressure, humidity and cloud conditions.
Inflow
4 major rivers
Boundary condition
Outflow
The Hitachi tone river
Initial condition

Water temperatureTable 2 Forcing scenarios of Lake Kasumigaura simulation
Base CaseCase-1Case-2Case-3

Inflow⚪
⚪
⚪

Insolation
⚪
⚪
⚪

Wind
⚪
⚪
⚪

4. Results and Discussion
4.1 Model Verification
Figure 2 shows the evaluation of simulation results in thebase case at all four stations. It can be seen that the calculated
results have good consistency. Root mean square errors(RMSEs) of temperature were 0.51, 1.13, 0.56, and 0.91, at theKakeumaoki, Hirayama, Koshin, and Asou stations,
respectively. These results show model has enough prediction
skill and still have a potential for improvement, especially in
representation of vertical stratification and its diurnal change.

Fig. 4 Particle tracking result at about 11:00 on September 2.
The wind chart with red lines marking the time points.
By comparing the cases in Fig. 3, the temperature
differences between surface and bottom layers is greater in
Case-1, which shows the effect of wind encouraging surface
heat exchange. In the case without insolation (Case-2), there isalmost no temperature stratification, which indicates that solar
radiation plays an important role in the formation of temperature
stratification. Case-3 combined with the initial setting of inflow
water temperature can also indicate that inflows has a certain
effect on water temperature distribution, which is inversely
proportional to the distance between the location and the inflow
entrance. Figure 4shows the surface particles like algae moveto the north shoreline by the prevailing south wind, resulting ina strong correlation between surface flow and wind.

5. Conclusions and Future Plan
Fig. 2 Comparison of model and observations.
4.2 Results
Figure 3 shows the temporal change of temperature during the
calculation period in the 4 cases. Red and blue lines indicate
surface and bottom temperature, solid and dashed lines are thecalculation results and the observations, respectively. Figure 4
shows the results of particle tracking to illustrate the winddriven effect on motion of surface materials.

Fig. 3 Comparisons between calculated and observed values offour cases at the Koshin station.

The calculated results show a good performance for the
obvious diurnal stratification. With evaluation indices it could
be said that SUNTANS can maintain reproducibility in the lake
region as well as application in coastal and estuarine areas.
Through the comparison of results, we have preliminarily
grasped the driving effect of wind, the great contribution of
insolation to temperature stratification and the importance ofcorrect setting of inflow conditions. Particle tracking simulation
can provide a suitable tool for revealing current flow effect on
the material cycles.
In order to furthermore investigate the specific causes andcharacteristics of water blooms and hypoxia in this lake, it isnecessary to add the feature of aquatic plants into an ecologicalmodel, and then couple with SUNTANS to achieve higher
reproducibility in biogeochemical simulation.
References
1) Doudoroff P, Shumway D L. Dissolved oxygen
requirements of freshwater fishes[J]. 1970.
2) Fringer O B, Gerritsen M, Street R L. An unstructured-grid,
finite-volume, nonhydrostatic, parallel coastal ocean
simulator[J]. Ocean Modelling, 2006, 14(3-4): 139-173.
3) Chua VP, Fringer OB (2011) Sensitivity analysis of threedimensional salinity simulations in North San Francisco
Bay using the unstructured grid SUNTANS model. Ocean
Model 39(3):332–350
4) Rayson MD, Gross ES, Fringer OB (2015) Modeling thetidal and sub-tidal hydrodynamics in a shallow, micro-tidal
estuary. Ocean Model 89:29–44
5) Guillaume, A., and Wells. J.C. "Gyre generation after atyphoon-induced upwelling in a stratified lake."
International Symposium on Stratified Flows. Vol. 1. No. 1.
2016International Course of Maritime & Urban Engineering / Abstract of Master Thesis

July 24, 2019

Full Scale resistance Prediction of paint surface
Proposal of new approximate formula for friction with roughness
Ren XuHull Form Design Sub-Area, Department of Naval Architecture and Ocean Engineering
Key Words: Frictional resistance, surface roughness

1. Introduction
Both the environmental pollution problem and high fuel costs
demands ship and shipping industry to find a way for savingenergy. Reduction of hull resistance is one of the ways to improveenergy efficiency. As frictional resistance account for a majorpart in ship total resistance, especially for low speed large block
coefficient ship, reducing the frictional resistance of ships is avery significant way for energy saving. So many new low friction
coating have been proposed and have been used for the real new
ships and existing ships to reduce the fuel consumption. But thetesting methods of new coatings and prediction method based on
experimental results have not been fully established. Inoue et. al.
made the experiments using 3m and 5m simple models withdifferent coatings and determined the equivalent sand roughness
based on the measurement results. They predicted the full scale
frictional resistance based on estimated equivalent sand roughness.
On the contrary, Kawashima et. al. did the resistance test for 2mflat plate model with two (long and short) wave length surface
roughness plate and 14m plate with medium wave lengthroughness. They showed the resistance coefficient increase by
medium flat plate does not show the intermediate value between
two 2m plate long model with long and short roughness. They
also show the mean shear stress of plate for each speed vsresistance coefficient increase and they could explain theroughness effect. They predicted 200m ship resistance increase by
their proposed method.
In the first stage of this research, the equivalent sand roughness
hypothesis is tested using Kawashima et. al. data. The results
show the promising results.
So, in the second stage, new approximate formula of frictional
resistance is proposed based on the computed results by theextended method of Katsui for roughness surface. The equation isfunction of two Reynolds number based on plate length andequivalent sand roughness height.

Substitute function (1) into the definition of momentum
thicknessuu(1  )dy
U
U
0

 

2

3 
     1       2  2 1  A    2 
2   
 

 

(2)

u
y
, 
U


The momentum thickness can be solved by momentum integral
equation as below:d1
 2  Cf
dx
2

(3)

From equation (2) and (3), we could calculate the frictional
resistance coefficient with different surface roughness.

3. Comparison with Experiments
Inoue et.al have tested previous calculation with the experimentresults by Osaka University1) which shows that we could
determine the equivalent sand roughness by previous method. In
this study, we made a new program similar as Inoue and tested
this program with the experiment data from NMRI.
National Maritime Research Institute have done flat plateexperiment with different surface roughness. The flat plateexperiment using 2m flat plate with short and long wavelengthroughness surface, and 14m flat plate with medium wavelengthroughness surface. The original data sets were provided by NMRI
and Chugoku Marine Paints, Ltd.

2. Calculation
In this study, the frictional resistance coefficient is caluculated
by logarithmic law with Coles law of the wake following Katsui
and White function of surface roughness:u 1 u y

y 
 ln(
)  5.0  B  1  cos(  ) 

u



 
  0.62  1.21exp( 

B 

1






290

ln 1.0  0.3k 



)

(1)

Fig.1 Comparison of experiment results by NMRI of 2m flat
plate and computational calculations.
Fig.1 shows that the calculations of the 8μm and 20μm agreewith the experiment results of long and short wavelengthroughness. So we considered the equivalent sand roughness of
medium wavelength roughness should be about 15μm in the
middle of 8μm and 20μm.

Fig.3 shows that the estimated frictional coefficient of 15μm,
55μm and 85μm could also fit the computation line with good
accurancy although frictional resistance for those roughness are
not used to redict the simple approximate formua.
We also calculated the frictional resistance coefficient withdifferent ship speed shows as Fig.4

Fig.2 Comparison of experiment results by NMRI of 14m flat
plate and computational calculations.
Fig.2 shows that the calculation of 15μm agree with theexperiment results of medium wavelength roughness. It means theestimation given by 2m model is reasonable.
As the comparison between calculation and experiment datashows good agreement for very heigh Rn, we could predict thecoating resistance of 14m model based on the 2m modelexperiments. So, it suggest that the prediction of full scale can bedone using equivalent sand roughness concept.

4. Approximate Formula
The previous calculation of frictional coefficient is complex
which is not easy to be widely used. In this chapter, we propose
an approximate formula which could be easily calculated.
Katsui’s formula3) shows good agreement with Osaka’s newexperiment and could calculate frictional resistance coefficient forsmooth surface, the calculation data of Prof. Katsui’s formula isexpressed as CF0

CF 0 

0.0066577(log Rn  4.3762)0.042612log Rn 0.56725(4)

For the trend line by binomial expression between log10Rn and
CF(Uk/ν) *log102Rn / CF0 shows good correlation, we propose
the approximate function as below:
CF (

Uk



)  CF 0 ( A0 

A1
A2

)
log Rn log 2 Rn

(5)

A0, A1, A2 are parameters calculated from roughness Reynolds
number Uk/ν=constant (0~2000).
We calculated frictional resistance coefficient of different
equivalent sand roughness (15μm,55μm,85μm) with same ship
speed by the approximate function shows in Fig.3.

Fig.4

Comparison of calculations by program andapproximate formula with different ship speed.
In Fig.4, lines are the computational results and circles arecalculation results by approximate function which also shows
good correlation.

5. Conclusions
In this study, we calculated frictional resistance coefficient bylogarithmic law with Coles law of the wake and White function of
surface roughness, and the comparison with the experiment dataof National Maritime Research Institute shows promising resuts
to support the equivalent sand roughness concept. The
measurement of surface roughness is difficult and same
geometrical roughness with different materials shows thedifferent frictional resistance performance. So, by comparing theexperiment data with the calculation lines, the equivalent sandroughness of painting surface is estimated and full scale
performance with the tested coating can be predicted bysimple
formula. The previous approximate formula proposed by Inoue et.
al. was not accurate, So, in this study, the approximate formula by
the parameter from roughness Reynolds number Uk/ν=constant is
proposed. By comparing the estimated result with the program
computation shows good agreement for the Rn=1.0x106~1.0x1010, the proximate function could be used for calculation of
the frictional resistance coefficient for full scale ship.
The author express her gratitude to Dr. Hideki Kawashima andDr. Hirihisa Mieno for prividing the data of the collaborative
research between National Maritime Research Institute, Japan andChugoku Marine Paints, Ltd, Japan.

References

Fig.3

Comparison of calculations by programapproximate function with same ship speed.

and1) Inoue, R: A consideration on the resistance test of the modelwith paint roughness, JASNAOE Conference Proceedings 15,
pp.77-80, 2012.
2) Kawashima, H: Scale effect on frictional resistance of
painited rough surface-Comparison of frictional resistance oflong and short painted rough surface flat plate, JASNAOE
Conference Proceedings 24, pp.729-734, 2017.
3) Katsui,T: The proposal of a new friction line, Proceedings of
the 5th Osaka Colloquium, pp.76-83, 2005.

International Course of Maritime & Urban Engineering / Abstract of Master Thesis

July 25, 2019

Study of Dolphin Swimming Performance by CFD
Yujia YAO
Marin Mechanism System Engineering，Department of Naval Architectural andOcean Engineering
Key Words: Dolphin, CFD, Oscillation motion, Thrustdolphin), a single phase transient flow solver was used withmoving mesh techniques. The Reynolds Averaged Navier-Stokes
Marine life has excellent swimming performance because of(RANS) Equations can be written as below:
nature selection. Among various species, dolphin has high speed
and high efficiency when swimming and has been an area ofinterest for both biologists and engineers to develop an energy
2.2 Assumption and Turbulence Model
efficient propulsion mechanism for applications on a variety ofSince the purpose of this step is to estimate the flow behavior
vehicles ranging from underwater robots as well as surface
around a dolphin, the dolphin is assumed to be in a fresh water at
cruising robots to ultimately, ocean-going ships.
room temperature rather than in cold sea water. Therefore, theGray (1936) pointed out that the propulsion performance of akinematic viscosity and density of fluid were assumed to be thatdolphin is seven times larger than that of any land mammal. Heof fresh water. The effect of free-surface on the flow field as wellobtained this conclusion based on the assumptions that theas the dolphin body was neglected. Moreover, the fluid is fresh
maximum speed of dolphins is 20 knots (10 m/s), the friction dragwater which is incompressible and no cavitation occurred duringof the dolphin body can be estimated by assuming the boundary
the dolphin motions. Thus, an incompressible transientlayer around the body is turbulent and the propulsive efficiency issingle-phase flow solver was used. The effects of dolphin skin on
100%. In the same paper, he also pointed out that if thethe boundary layer such as skin roughness are neglected in this
propulsion performance of dolphins is similar to terrestrial
analysis. The dolphin skin is assumed to be a smooth, water-tightmammals, these very high swimming speeds are due to thebody without nostrils. The dolphin is in neutral buoyancy with nolaminar boundary layer around the dolphin body. However, after
significant gravitational, buoyant and Coriolis forces affecting its
an extensive study, Fish (1993) concluded that the dolphin
motion as well as the flow field. The inflow is a uniform flowswimming speeds are not related to the boundary layer around thewith medium turbulence intensity (5%). The body of the dolphin
dolphin.
is an isotropic material without any form of nonlinearities
Therefore, it is necessary to understand the effect of dolphin
including geometric and material non-linearities. The motion ofswimming pattern on its very high propulsive performance. Isogai
the body surface is directly proportional to that of the backbone.
(2014) carried out numerical analysis on the flow around theBased on these assumptions, the Reynolds number for thiscaudal fin of a bottlenose dolphin (Tursiops truncates) by a 3D
scenario is estimated to be approximately 1.5x107. Therefore, it isNavier-Stokes (NS) approach. In his work, computations werenecessary to carry out turbulence modelling in the computations.
conducted based on the standing swimming motion of dolphin,
Spalart-Allmaras turbulence model was chosen due to its stability
neglecting the effect of flow around the dolphin body as well as
and the lack of severe flow separation due to streamline shape of
the interaction between the body and caudal fin. However, it isthe dolphin.
necessary to include these effects in order to fully understand thedolphin swimming mechanism.
3. Simulation
The numerical simulation of dolphin swimming patterndeveloped by Xiaozhou Wang (2018) is inconsistent with the real
3.1 Definition of Fin/Body motion
situation. The thrust should be higher when using swimming
The motion algorithm for general fish swimming was
pattern which is close to real dolphin motion. Thus, in this paper,
developed based on the mathematical model proposed in
the simulation of swimming pattern which is similar to real
Wolfgang et al. (1999) with the assumption that thesituation has been developed and compared with the mode ofdisplacements of the body-surface is the same as that of the
Xiaozhou Wang. The thrust of each pattern is calculated and webackbone of fish. According to their observations, thecan see the pattern which is similar to real dolphin swimmingtime-dependent transverse motion of a fish in two dimensionalmotion can generate larger thrust. And in the future research, realreference frame can be expressed as:
dolphin swimming data should be obtained to get reliable
swimming pattern.

1. Introduction

2. Methodology
2.1 Numerical Method
In this research, numerical analyses were carried out tounderstand the flow around the body of a swimming dolphin bysolving 3D Reynolds Averaged Navier-Stokes (RANS) Equationsin an open-source finite volume framework, OpenFOAM. In
order to compute an unsteady flow around a moving body (thewhere 𝑘 is wave number of body motion, 𝜔 is the period ofbody motion, 𝐴𝑙 is amplitude of the deformation expressed in theratio of body length 𝑙. And a(x) describes the shape ofbackbone deformation which can be expressed by asixth-order polynomial function as below:
Where A, B, C, D, E, F, G are coefficients.
Fig.1 shows the pressure filed on the dolphin body, Fig.2

shows the grid around the dolphin body. Fig.3 shows the velocity
field around the flapping dolphin.

length of the body which will be described as small and middle
amplitudes. Table 1 shows the derived thrust of small amplitude
fluke motion. Table 2 shows the thrust comparison obtained fromthis research and in Xiaohou Wang’s research.
Table 1 Thrust derived under small amplitude fluke motion
U (m/s)Thrust (N)
f=1Hz
f=2Hz
f=3Hz

1

45.366

3-109.3061

172.227

3

9.396

1

854.0623

416.217

Table 2 Thrust (N) compared with WANG’s result

Fig. 1 Pressure field on the dolphin bodyWang(2018)

Present

f=1Hz

10.587845.366f=2Hz

34.6834

172.227

f=3Hz

84.1642854.0616

4. Conclusions
Fig. 2 Grid around the dolphin body

From the CFD simulation, the hydrodynamic force ofswimming dolphin is obtained. Compared to Xiaozhou Wang, themotion equation used in this paper obtained larger thrust ofswimming dolphin, this shows that the large thrust of real dolphinswimming is related to the swimming pattern.
And we can see the thrust increases as amplitude increases
and also increases as the frequency increases.
In future research, the geometry model obtained from real
dolphin should be used and dolphin swimming motion of large
amplitude and high velocity should be conducted to get reliablethrust.

Fig.3 Velocity field around the flapping dolphin

References
1)

3.2 Derivation of Hydrodynamic Force by CFD
2)

3)
4)

Fig. 4 The hydrodynamic force when frequency is equal to2Hz
The hydrodynamic force is obtained by integrating pressure
and sheer stress around the dolphin body. The hydrodynamic
force when frequency is equal to 2Hz is shown in Fig.4.
Because the dolphin conducts oscillation motion, thehydrodynamic force is also periodic. Two amplitudes ofpeak-to-peak tail motion are chosen for this study:10%,15%

5)

6)Fish, F. E.: Power output and propulsive efficiency ofswimming bottlenose dolphins (tursiops truncatus).
Journal of Experimental Biology, pp.179–193,1993.
Gray, J.: Study in animal locomotion iv-the
propulsive powers of the dolphin. J. Exp. Biol.,pp.192–199,1936.
Isogai, K.: Effect of flexibility of the caudal fin on the
propulsive performance of dolphins,2014.
Ltd, O.: Openfoam v1706: New and improved numerics,
2017.
Wolfgang, M., Anderson, J., Grosenbaugh, M., Yue, D.,
and Triantafyllou, M.: Near-body flow dynamics inswimming fish. Journal of Experimental Biology,
pp.2303–2327.38,1999.
Xiaozhou Wang, Peng Wei, Ye Yuan, Zhiguo Zhang,
Dakui Feng.: Evaluation of Dolphin Swimming Speed andThrust Based on CFD, ISOPE Journal Paper, Vol.28, No.2,
pp.120-127,2018.

International Course of Maritime & Urban Engineering / Abstract of Master Thesis

July 31, 2019

Study on Measurement Method of Under-Film Corrosion’s Incubation
Life Using Fe (II) Fluorescent Indicator
Azhan Bin Abdul Halim
Ocean Material Engineering Laboratory
Department of Naval Architecture and Ocean Engineering
Key Words: Nondestructive monitoring, under-film corrosion, water ballast tank, Fe (II), fluorescent indicator, spectrum imaging

1. Introduction
It is necessary to carry out nondestructive monitoring of underfilm corrosion of anti-corrosion coated steel panels such as waterballast tanks (WBT) and investigate the statistical nature of thecorrosion’s incubation life to improve the structural safety of
aging vessels. Hirayama et al.1) synthesized the compound (1S)3'-(diethylamino)-N,N-diethyl-3-oxo-3',9a'-dihydro-3H
spiro[isobenzofuran-1,9'-xanthen]-6'-amine-oxide or RhoNox -1
that reacts with Fe (II) to produce chelation-enhanced
fluorescence (CHEF). Osawa et al.2) prepared semi-transparent
epoxy paint added with RhoNox-1, conducted corrosion tests forseveral steel panels coated with it and confirmed the corrosion
detection of RhoNox-1. Several scribed test panels were exposed
to sodium chloride solution drops to introduce initial corrosion
and then coated, were used in this study. Because of this, Fe (II)was formed before or after the start of corrosion test and it was notpossible to determine the corrosion’s incubation life.
The purpose of this study is to develop a method to measure thecorrosion’s incubation life by detecting the forming of Fe (II)under the coating film. The anti-corrosion coating system for
actual WBTs has high corrosion resistance and it takes years togenerate detectable amount of Fe (II). In this study, it is necessary
to compare corrosion’s incubation lives by changing the steel type,
coating specification as well as corrosion environment. For thispurpose, it is important to shorten the test time by using coated test
panels whose corrosion resistance has been lowered to an extent
that Fe (II) formation could be detected in a short time. The anticorrosion performance of the paint can be controlled by mixing
multiple hardeners with different anti-corrosion performances.
In this report, several hardeners were mixed and semitransparent epoxy paints containing RhoNox-1 with different anticorrosion performance were prepared. Corrosion tests of the steel
panels coated with the RhoNox-1-containing paints wereconducted. During the corrosion tests, spectrum imaging (SI)measurements of the coating surface were performed to detect Fe(II) formation on the surface under the coating film without initialcorrosion.

Table 1

Hardener mix ratio and target DFT for test panels.

Test plate
E41, J41
E58, J58
E62, J62
E66, J66

H:L
50:50
100:0
100:0
100:0

Target DFT (µm)
50
50
100
200

3. Fluorescence measurement and test condition
RhoNox-1 was synthesized in an amount sufficient to coat thetest panels with the same material and synthesis method as
previous report2). RhoNox-1 does not produce prematurely
fluorescent CHEF by the reaction with amine, which is the main
component of the epoxy paint’s hardener, but emits fluorescenceat around 570 nm when reacting with Fe (II).
Due to the limitations of the equipment used for this study, theconfocal laser scanning microscopy (CLSM) measurement was
performed at λex=473 nm. It was observed that the reaction
between RhoNox-1 and Fe (II) can be detected by the increase inemission intensity around λem=570 nm2). The test panels were
subjected to immersion test with water at 50 °C and thefluorescence measurement of the coated surface of test panels was
conducted.

Fig. 1 Fluorescence spectrum of RhoNox-1 with or without
FSA(II) in DMSO (λex/λem=540nm/570nm).

2. Coating specification
A side of rectangular panels of high tensile strength steel for
shipbuilding, with dimensions 150mm×35mm×3.2mm, blast
grade Sa 2.5, surface roughness Rz 30μm, were coated with a clear
two-pack epoxy paint; main agent and hardener. The hardener forthe epoxy paint was prepared by mixing a PSPC-certified hardener
for WBT paints; Hardener H with high corrosion resistance andHardener L with low corrosion resistance, to adjust the anticorrosion performance to be significantly reduced compared to theactual paint. Each test plate was coated with the paint withdesignated target dry film thickness (DFT).

Fig. 2

RhoNox-1-coated panels subjected to immersion test.

4. SI measurement
The fluorescence measurement was performed according to
the following procedure. A rectangular measurement area for SI
of 2.5 mm side lengths or 800×800 pixels was set at two locations
on each test panel. The SI measurement area was divided into
rectangular areas with dimensions of 10×10 pixels (hereinafterreferred as “area elements”). λ mode measurement was performed
to acquire SI data of the whole measurement area, light emission
intensity was averaged within each surface element, and theaveraged spectrum of each surface element was calculated. The
averaged spectrum was subjected to Savitzky-Golay smoothing oforder k=2 and half band width n=3, and the maximum wavelength
λmax and the maximum emission intensity Imax of each element
were calculated.
As shown in Figure 1, RhoNox-1 used in this study emits
maximum fluorescence intensity around λem=570 nm when it
reacts with Fe (II). Therefore, if Fe (II) is formed under the coating
during immersion test, it is expected that Imax will increase andλmax near 570 nm at the anode part.

5.

Fluorescence emission characteristics

Figure 3 shows the box plot of each surface element Imax forseveral panels. The thick lines in the box plot represent the medianQ2/4, the upper and lower sides of the box represent the third
quartile point Q3/4 and the first quartile point Q1/4, and the upperand lower whiskers represent the upper and lower limits of thesection of the following equation.
Q1 4  1.5IQR, Q3 4  1.5IQR  ; IQR  Q3 4  Q1 4

(1)

The Imax contours for two panels, E58 and J58 are shown in
Figure 4 with the colour level of the contour. The test panels here
were conventional steel and corrosion-resistant steel respectively
coated with the paint that have the same corrosion resistance.

Fig. 4 Comparison of Imax ‘s contour maps of panels E58 andJ58 during the hot water immersion test.

6.

Conclusions

Several hardeners were mixed to prepare RhoNox-1containing semi-transparent epoxy paint with different anticorrosion performances. Corrosion tests for the steel panels coatedwith the prepared RhoNox-1-containing paint were conducted.
During the corrosion tests, spectrum imaging (SI) of the coatedsurface was performed to detect Fe (II) formation on the surface
under the coating film without initial corrosion. The findings
obtained in this study can be summarized as follows.
(1) Coated test panels using the RhoNox-1-containing semitransparent paint prepared in this study were prepared. By
evaluating the fluorescence maximum wavelength and maximum
emission intensity for each rectangular area of 10 pixels side
length from SI result, it is possible to estimate the spatial
distribution of Fe (II) formation and increase due to the under-film
corrosion. This indicates that it is possible to measure the underfilm corrosion’s incubation life and the degree of anodic reaction
progress under the coating film by fluorescence measurement
using the RhoNox-1-contaning semi-transparent paint developedin this study.
(2) Under the test conditions in this study, it was shown that as fortest panel with the higher mixing ratio of high corrosion resistancehardener and dry film thickness, the formation of Fe (II) was
slower and the progress of corrosion was slower. This result
indicates that the fluorescence measurement using RhoNox-1containing semi-transparent paint developed in this research may
be able to evaluate the effect of the factors such as coating
specification and steel type influencing the under-film corrosion’s
incubation life.

References1)

2)

Fig. 3Comparison of Imax‘s boxplots of different panels.

Hirayama, T., Okuda, K., Nagasawa, H. (2013) “A Highly
Selective Turn-on Fluorescent Probe for Iron(II) to Visualize
Labile Iron in Living Cells,” Journal of Royal Society ofChemistry, 4, pp. 1250-1256.
N. Osawa, A. Takada, T. Kojima, S. Takano (2015)“Development of Fe(II) Fluorescent Indicator for UnderFilm Corrosion Monitoring of Epoxy-coated Steel Panels,”
Proceedings of The Japan Society of Naval Architects andOcean Engineers, 21, pp. 93-99.