dr. ir. J.H. den Besten Delft University of Technology

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{CAPEX, OPEX} evaluation of structural element innovations in marine structures

Marine structures are typically structural member assemblies, consisting of characteristic elements in

stiffened panel and {truss, frame} setup (Fig. 1). Stiffened panels typically appear in {free,

constrained} floating marine structures like FPSO’s; {trusses, frames} are common in fixed marine

structures like jacket supported wind turbines.

Figure 1: stiffened panel and {truss, frame} elements in marine structures.

Fatigue & fracture is a governing limit state and typically confined to – welded – joints connecting the

structural members in the stiffened panel and {truss, frame} structures (Fig. 2) because of the

macroscopic stress concentrations, supporting the (welding induced) stress concentrations at {micro,

meso}-scale.

Figure 2: fatigue damage in stiffened panel (DNV report) and {truss, frame} structures (Dong et al.,

2012. Rel. Eng. and System Safety, vol. 106).

Fatigue (resistance) involves 4 interactive dimensions. The reference fatigue resistance consists of a

material and geometry contribution. The fatigue influence factors loading & response and

environment will define the actual fatigue resistance in operational conditions.

To improve the reference fatigue resistance w.r.t. material, functional grading and micro-structural

tailoring is considered to be one of the solutions. At {micro, meso} scale it means that at the material

surface the initiation resistance needs to be improved; sub-surface, improved crack growth

resistance and fracture toughness properties are required. Since the quasi-static σ-ε curve based

toughness as measure for the fracture toughness decreases with increasing {yield, ultimate} strength

whereas the fatigue initiation resistance increases with increasing strength, conflicting requirements

will be part of the game. At macro-scale, functionally grading can be applied using a composite. For a

bending dominated response a sandwich configuration seems obvious.

transverse frame

(beam)

1 bay stiffened panel

stiffener (beam)

longitudinal girder

(beam)

{shell, plate}

joints (nodes)

beams

(tubular members)

dr. ir. J.H. den Besten Delft University of Technology

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To improve the reference fatigue resistance w.r.t. geometry, replacing a conventional stiffened panel

by a sandwich (i.e. macroscopic functionally graded material) panel first of all reduces the number of

hot spots since the secondary order stiffeners are eliminated (Fig. 3). Adopting an advanced welding

technique like (stationary shoulder) friction stir welding, the joint fatigue resistance can be

significantly improved. For groove welds the weld reinforcement has been eliminated in comparison

to its conventional arc-welded equivalent. The fillet weld geometry is fully controlled and contains a

relatively large notch radius reducing the weld notch stress concentration (i.e. fatigue sensitivity).

Because the heat input for friction stir welding is relatively small, the residual {stress, deformation}

level is quite small and the {number, size} of welding induced defects is significantly reduced,

meaning the relative contribution of crack initiation to the total fatigue resistance has significantly

improved. Even the sandwich face sheets can be functionally graded at {micro, meso} scale to

improve the fatigue resistance (and provide sufficient weldability).

Figure 3: Sandwich panel and characteristic friction stir welded {butt, T-} joints.

For {truss, frame} structures like wind turbine jackets the (tubular) welded joints appear in relatively

complex geometries like bi-planar structural member intersections and the transition piece (Fig. 4).

To remove the welded joints to outside the member intersection area (i.e. to split stress

concentrations) a wire+arc additive manufactured node, welded to the structural members, allows

for {micro, meso}-scopic functional grading and microstructural tailoring and exploitation of the

provided topology freedom to control the far field stress level and (macroscopic) stress

concentration (i.e. {reference, peak} stress).

Figure 4: Structural member intersection (i.e. tubular joint) adopting wire+arc additive manufacturing

for a transition piece (Lee et al., 2016, Structural topology optimization of the transition piece for an

offshore wind turbine with jacket foundation, Renewable Energy, vol. 85, pp. 1214 – 1225) and multi-

planar K-joint (Dong et al., 2011, Long-term fatigue analysis of multi-planar tubular joints for jacket-

type offshore wind turbine in time domain, Engineering Structures, vol. 33, pp. 2002 – 2014).

dr. ir. J.H. den Besten Delft University of Technology

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1. Stiffened panel

To evaluate the {CAPEX, OPEX} for a sandwich panel relative to its conventional arc-welded

equivalent, an FPSO deck structure will be considered. The characteristic stiffened panel parameters

are defined as:

frame spacing: sf = 2500 [mm], stiffener: bulb profile HP200x10

stiffener spacing: ss = 750 [mm], number of stiffeners: ns = 3

base plate thickness: tb = 16 [mm], number of frames: nf = 2

frame web height: hw = 800 [mm], stiffener area: As = 2566 [mm2]

frame web thickness: tw = 12 [mm], stiffener area moment of inertia: Is = 1.02·107 [mm4]

frame flange width: wf = 200 [mm], neutral axis position: na = 119 [mm]

frame flange thickness: tf = 15 [mm], stiffener height: hs = 200 [mm]

The structure will be exposed to (local) pressure loading p = 10 [ton/m2] = 0.1 [N/mm2].

A global hull girder loading component is ignored for now as it applies to both the stiffened- and

sandwich panel. Considering the local response only is considered to be sufficient for the sake of

comparison. The local response predominantly defines the stress concentration; the global response

affects mainly the far field stress.

Material: S355

yield strength: σy = 355 [N/mm2], density: ρs = 7800 [kg/m3]

The response level of the conventional stiffened panel will be established first in order to define the

sandwich panel equivalent dimensions.

Plate-stiffener beam bending response:

From fatigue limit state point of view the governing hot spot is located on top of the stiffener at the

frame connection. The response is bending defined, meaning the neutral axis position is required:

�̅ = ��∙��∙� ��� ��������∙����� ≈ 30 [mm]

Area moment of inertia:

�� = ��∙����� + �� ∙ �� ∙ ��̅ − ��� �� + �� + � ∙ �� + !" − �̅�� ≈ 4.46 ∙ 10' [mm4]

Note that the plate is assumed to be fully effective; shear lag effects are ignored for now.

Bending moment: because of hierarchy and symmetry clamped boundary conditions are adopted

(� = − )∙�*�� = − +∙��∙�*�� = −3.90625 ∙ 10'[Nmm]

dr. ir. J.H. den Besten Delft University of Technology

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Nominal bending stress:

/� = 0�∙12� = 0�∙ ���3�45̅�2� ≈ −163 [MPa]; compressive

Sandwich beam bending response:

Nominal bending stress:

/�,7 ≈ (���89 ∙ �7

The core height can be optimised (Wijker, 2008, Spacecraft Structures, ISBN: 978-3-540-75552-4,

Springer-Verlag, Berlin Heidelberg, Germany) w.r.t weight based on {strength, stiffness, face sheet

dimpling}:

89 = : ∙ ;<*<=? ∙ �7 with C = {2, 4, 2/3}

Note that the hc-tf relation is only used to have some guideline in the free design space. Substitution

in the bending stress equation provides a tf formulation:

�7 = @ (���: ∙ /�,7 ∙ �A7A9�

The larger C, the smaller the required tf will be. However, considering typical wall thickness

constraints in the maritime industry from production, corrosion, and deformation perspective, tf ≥ 5

[mm]. In order to define the sandwich dimensions the {stiffened, sandwich} panel response level

should be similar (σb = σb,f) for a fair fatigue resistance comparison. The core configuration (e.g. foam,

corrugation) is not established yet, but its density is estimated as 10 [%] of the face sheet material.

For foam this is a typical value and for a corrugation a similar value can be achieved. Then only for C =

(2/3) the tf criterion is satisfied: tf ≈ 6 [mm]. Consequently: hc ≈ 50 [mm]. Average sandwich core

shear stress check:

B9 = +∙�*�∙3= ≈ 3 [MPa]

This magnitude is acceptable, but should not be larger considering typical foam shear strength

criteria.

hc

tf

tf

dr. ir. J.H. den Besten Delft University of Technology

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0 50 100 150 200 250 300 3500.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.50

most likely max. stress range S in N = 1.108 cycles [N/mm2]

DFS

W /

Dar

c

damage ratio

0 50 100 150 200 250 300 3500.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.50

most likely max. stress range S in N = 1.108 cycles [N/mm2]

DFS

W /

Dar

c

damage ratio

hc = 50

hc = 40

hc = 30

Fatigue resistance:

For the arc-welded reference case the DS T-joint is considered: FAT80 at best. The hot spot is a limit

case of type C. Alternatively an attachment (type A) could be considered, but it provides the same

FAT class (shortest attachment length because the attachment length is the frame web thickness).

Slope marc = 3.

For friction stir welding (FSW) the fatigue resistance is established using some recent test results

(Polezhayeva et al., 2015, Fatigue performance of friction stir welded marine grade steel, Int. J. of

Fatigue, vol. 81, pp. 162 – 170) available for a butt joint: FAT183 and slope mFSW = 5. The slope

indicates an increased contribution of initiation to the total fatigue damage. For an arc-welded butt

joint the fatigue strength is established as FAT112. Using the arc-welded joint FAT class ratio (80/112)

the fillet FSW class is estimated as FAT128.

Using a simplified (spectral) fatigue assessment the most likely max. stress range in N = 1.108 cycles is

adopted as resistance criterion:

C = D: ∙ E F3G! D�HI3 ∙ Γ �1 + K8�

Assuming the Weibull shape parameter h = 1 (i.e. exponential distribution), the FSW and arc-welded

fatigue damage ratio is obtained considering the FSW panel response as a multiplier (factor Cr) of the

stiffened panel response:

CLMNC"O9 = :"O9:LMN ∙ Γ 1 +KLMN�Γ 1 + K"O9� ∙ ;:O ∙ F"O9G! D� ?IPQR ∙ EG! D�F"O9 HI�S=

For hc = 50 [mm], Cr = 1.06. Reducing the core height to respectively hc = 40 mm and hc = 30 [mm] in

order to save some more weight, the sandwich panel response increases: Cr40 = 1.33 and Cr30 = 1.75.

The damage ratios (Fig. 5) show that for hc = 50 the fatigue damage reduces to ~10 [%] of its arc-

welded equivalent; for hc = 40 up to ~30 [%]. For hc =30 [mm] no fatigue damage reduction is

obtained.

Figure 5: fatigue damage ratio for different sandwich core heights.

dr. ir. J.H. den Besten Delft University of Technology

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Structural weight:

Stiffened panel:

T = U !� ∙ ��� ∙ V!7 ∙ �7W ∙ �+ + V!7 ∙ �7W ∙ !� ∙ � + !7 ∙ !� ∙ ��� ∙ V8X ∙ �X +Y7 ∙ �7WZ ∙ A� ∙ 104[

T ≈ 2150 [kg]

Sandwich panel:

T = VA9 ∙ 89 + 2 ∙ A7 ∙ �7W ∙ \V!7 ∙ �7W ∙ !� ∙ ��� + !7 ∙ !� ∙ ��� ∙ � 8X − 89� + Y7�] ∙ 104[

T^_ ≈ 2060 [kg] → ~5 [%] reducTon T̀ _ ≈ 1940 [kg] → ~10 [%] reducTon Ta_ ≈ 1820 [kg] → ~15 [%] reducTon

The weight reduction for metal core materials is limited; i.e. depends on core material density and

height.

Costs:

Material:

Foam will be considered as sandwich core material. However, steel foam prices are hardly available.

Since for aluminium foam pricing is available, the material costs will be established for an aluminium

{stiffened, sandwich} first. It is assumed that the results can be extrapolated to steel. Costs of

aluminium {plate, profile} material: €10/kg. For an aluminium sandwich panel with foam core it is

estimated at €1000/m2, meaning:

Stiffened panel: € 21.500

Sandwich panel (~16 m2): € 16.000

The ratio of the stiffened- and sandwich panel costs as obtained in aluminium is assumed to be the

same for the steel counterparts. For the sake of simplicity it will be assumed that material costs for

the {stiffened, sandwich} panel will be the same.

Production (equipment):

FSW tools for steel are currently expensive (€ 3000/tool) and the service life L ~ 100 [m] at the

moment (comparison: for aluminium € 300/tool and L ~ 1000 [m]), but it is sufficient for

demonstrator development proving increased fatigue resistance and reduced life cycle costs. Friction

stir welds in steel are limited by tool technology to tp ≤ 20 [mm] but this is practically no limitation for

sandwich panels. The weld length for a sandwich panel in comparison to a stiffened panel is

dr. ir. J.H. den Besten Delft University of Technology

7

considered to be reduced to 80 [%], assuming that for a sandwich butt joint 2 groove welds are

required and for a sandwich T-joint 2 fillet welds as well as 2 groove welds at the back (Fig. 3).

Friction stir welding costs are already comparable to arc-welding costs for aluminium. Assuming that

arc-welding costs are the same for steel and aluminium the following estimate is obtained for steel

marine structures assuming it contains 100 [km] weld seam length (25 stiffeners in double {bottom,

shells}; i.e. 6 walls, welded at 2 sides over 200 [m] length, multiplied by 1.5 for frame welding ~ 100

[km]):

Arc-welding: € 0.30 /m weld length, meaning: € 30.000

FS-welding: € 30 /m weld length, meaning: € 3.000.000

Fuel:

A 1 [%] reduction in structural (i.e. light) weight reduces fuel consumption in the range 0.1 to 0.3 [%]

(Ship Energy Efficiency Measures: status and guidance, American Bureau of Shipping). The impact on

CO2 emissions is likewise.

Using the Admiralty coefficient for ships 0.7 [%] fuel reduction can be obtained for 1 [%] reduction in

total weight (Energy savings by light-weighting-II, 2004, International Aluminium Institute). For a mid-

sized bulk carrier it can mean an annual saving of € 45.000 and for a large container ship € 270.000

for 1 [%] fuel consumption reduction (ABB.com); in average approximately € 160.000 annually for a

marine structure.

Comparing the fuel consumption reduction for {light, total} weight, an average fuel consumption

reduction of 0.2 [%] for 1 [%] reduction in structural (light) weight corresponds to 0.2/0.7 = 0.3 [%]

reduction in total weight. It means that for an estimated light weight reduction of 10 [%] obtained

using sandwich panels a 3 [%] total weight reduction can be achieved.

For a total weight reduction of 3 [%] the fuel consumption can be reduced by 2.1 [%], meaning an

average annual cost saving of €215.000.

Fatigue induced maintenance:

Figures w.r.t. maintenance costs in relation to Total Cost of Ownership (TOC) are very limited and

some results available for Navy ships are used for reference (Fig. 6; Stambaugh K., Kaminski M.L.,

2016, Ship structure fatigue and life cycle risk management approaches, 5th Int. Symp. on Life-Cycle

Civil Engineering, Delft, The Netherlands) and extrapolated to industrial applications.

Figure 6: risk versus total cost of ownership (TOC).

SFA = (spectral) fatigue analysis (for design)

HSM = hull structural monitoring (X year data collection)

dr. ir. J.H. den Besten Delft University of Technology

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Asset value estimate:

Frigate: 500 M$, approximately 4 [M$/m] assuming an average length of 120 [m].

Bulk carrier: 250 M$, approximately 1 [M$/m] assuming an average length of 250 [m].

FLNG carrier (Prelude): 5000 M$, approximately 10 [M$/m] assuming an average length of 500 [m].

Approximation: 1€ = 1$.

A fatigue damage reduction to 10 [%] of its original value, as obtained for a sandwich panel with hc =

50, means a factor 10 reduction in probability of failure. Hence the risk reduces by the same factor

10. Assuming that an SFA is applied for design and that the step to HSM-15 is equivalent to the

friction stir welding induced fatigue resistance improvement (simply because it is in between HSM-5

and HSM-30), for an asset value of 500 M€ (navy ship) a risk reduction factor of 10 comes along with

a TOC decrease from 25M€ to 5M€, introducing the ratio 100:5:1 (asset value: maintenance cost

without fatigue resistance improvement : maintenance costs with fatigue resistance improvement).

The TOC saving is 20M€. For an asset value of 250 M€ the TOC saving estimate is 10 M€.

For a damage reduction to 30 [%] of its original value, as obtained for a sandwich panel with hc = 40,

the risk reduction is a factor ~3. The TOC decrease is from 25M€ to 9M€ for an asset value of 500 M€.

The ratio becomes: 60:3:1. The cost saving estimate is 16M€. For an asset value of 250 M€ the TOC

saving is estimated at 8 M€.

The reduction in number of governing hot spots from (2·ns) = 6 to 2 for each stiffened panel bay in

the considered configuration, reducing the probability of failure and risk even further, is not taken

into account.

{CAPEX, OPEX} evaluation:

Considering CAPEX only, costs will increase. The friction stir welding technology is still immature, but

over time production costs are expected to decrease. Cost savings will come from the OPEX part in

terms of fuel and fatigue related maintenance.

expenses stiffened

panel

sandwich

panel

note

CAPEX:

- material and production

5 M€

estimate is conservatively rounded

OPEX:

- fuel saving

- fatigue related maintenance

25 M€

-2 M€

8 M€

half the estimated value for 20 years

Total: 25 M€ 11 M€ at least 30 [%] cost reduction obtained

Table 1. {CAPEX, OPEX} for an asset value of 250 M€.

dr. ir. J.H. den Besten Delft University of Technology

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2. Tubular joint

To evaluate the {CAPEX, OPEX} for a wire+arc additive manufactured node relative to its conventional

arc-welded equivalent, a simple tubular T-joint (Fig. 7) will be considered.

Figure 7: tubular T-joint.

Several sources provide figures to increase insight in the potential cost savings for offshore wind

turbines. First, the asset value is estimated (Offshore wind project cost outlook, 2014 edition): ~3.5

M€/MW including installation; excluding installation ~1.8 M€/MW. For a 5 MW turbine, costs

including installation ~17.5 M€; excluding installation ~9 M€. Support structure and foundation costs

are 10…20 [%] of the total. The total costs are typically divided into 80 [%] CAPEX and 20 [%] OPEX.

The potential support structure cost reduction is estimated at ~5 [%] of total wind turbine

investment costs (Offshore wind project cost outlook, 2014 edition; Offshore wind power priorities

for R&D and innovation, Scottish Enterprise).

A maintenance cost reduction estimate: 5…10 [%] of total wind turbine investment costs (offshore

wind power priorities for R&D and innovation, Scottish Enterprise), but is expected to be a result of

smart maintenance programs.

dr. ir. J.H. den Besten Delft University of Technology

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Costs:

For the simple tubular T-joint (Fig. 7), the structural weight is estimated at 670 [kg].

Material and production:

• Arc-welded tubular joint: estimating for robotic welding the all-in material and production

costs at 50 €/hr and the production rate at 10 kg/hr; i.e. 5 €/kg, the costs are estimated at

€3.350. Note that joint complexity is not explicitly considered.

• Cast node: a weight reduction in the range 5 … 10 [%] in comparison to its arc-welded

equivalent can be achieved. A ~10 [%] cost reduction as well as ~10 [%] cost increase is

possible, depending on joint complexity; in average the same (Webster et al., 1981, Cast

steel nodes – their manufacture and advantages to offshore structures, J. of Petroleum

Technology, vol. 33).

Figure 8: {welded, cast} tubular single-planar K-joint (Wang, 2013, balance fatigue design of cast steel

nodes in tubular steel structures, Scientific World Journal, vol. 2013).

A cast node benefits from having no welds in the fatigue sensitive member intersection

regions (i.e. straight forward welding at the tubular member connections only; Fig. 8) and

optimised topology (i.e. reduced stress concentrations). For production, using a classical

foundry the material reference fatigue resistance is not necessarily improved because of the

casting process related {voids, pores, inclusions}. Vacuum steel production benefits from

excellent mechanical properties, fracture resistance and homogeneity.

• Wire+arc additive manufactured node: costs are estimated at 200…300 €/hr (technology is

still immature) and the production rate at 10 kg/hr; i.e. 20..30 €/kg, meaning the T-joint costs

become in average €16.750. The state-of-the-art wire+arc manufacturing induced surface

roughness requires special attention, either in terms of WAAM process modifications or

post-deposition machining, ~10 [%] more costs are considered: €19.000.

A wire+arc additive manufactured node benefits from optimal topology (reduced stress

concentrations in the fatigue sensitive regions), straight forward welding at the tubular

member connections and the possibility to apply functional grading and microstructurally

tailoring to improve the fatigue resistance.

dr. ir. J.H. den Besten Delft University of Technology

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0 50 100 150 200 250 300 3500.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.50

most likely max. stress range S in N = 1.108 cycles [N/mm2]

DW

AA

M /

Dw

eld,

cast

damage ratio

0 50 100 150 200 250 300 3500.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.50

most likely max. stress range S in N = 1.108 cycles [N/mm2]

DW

AA

M /

Dw

eld,

cast

damage ratio

Cr = 1.0

Cr = 0.9

Fatigue induced maintenance:

Adopting the DNV-GL regulation (DNVGL-RP-C203) the fatigue resistance curve for tubular joints (i.e.

welded node) in air or in seawater with cathodic protection is prescribed as: FAT90 (T-class), slope m

= 3. For cast nodes FAT100 (C-class) is assigned. Note that because of the involved structural hot spot

stress concept the geometry and loading induced stress concentrations are not incorporated. For

cast nodes these stress concentrations are typically smaller than for as-welded joints meaning the

difference in fatigue strength will be even larger in favour of cast nodes. Introducing functional

grading and microstructural tailoring using wire+arc additive manufacturing a fatigue strength

increase of ~30 [%] is expected to be possible: FAT115, w.r.t. its arc-welded equivalent. The curve

slope is estimated at m ~ 4 (the largest value adopted for tubular joints / nodes in the DNV-GL

regulation) assuming the crack initiation resistance has increased as a result of functional grading.

Using a simplified fatigue assessment (same assumptions as for the {stiffened, sandwich} panel) the

wire+arc additive manufactured and welded node fatigue damage ratio is obtained considering the

wire+arc node response as a multiplier (factor Cr) of the tubular joint response:

CN��0CX1cd/9"�� = :X1cd/9"��:N��0 ∙ Γ 1 + KN��0�ΓV1 + KX1cd/9"��W ∙ ;:O ∙ FX1cd/9"��G! D� ?IRffg ∙ E G! D�FX1cd/9"��HIhijk/=��

It is expected that using topology optimisation a response reduction of 10 [%] can be obtained: Cr =

0.9. Plotting the damage ratios shows that for Cr = 1 the fatigue damage reduces to ~25 [%] of its arc-

welded equivalents; for Cr = 0.9 up to ~15 [%].

Figure 9: fatigue damage ratio for different Cr values.

Let’s assume the asset value is 10 M€, 2 [%] of the 500 M€ asset the numbers have been provided for

(Fig. 6); the decrease in TOC will scale accordingly. A damage reduction to 15 [%] means a factor 6.5

reduction in probability of failure and hence a risk reduction of the same factor 6.5. The TOC

reduction will be from 0.5 M€ to 0.15 M€. In case once in the turbine life time the jacket requires a

fatigue damage repair, the costs for a jack-up are estimated at 0.2 M€. This number is of the same

order of magnitude as the extrapolated estimate(0.5 M€).

A damage reduction to 25 [%] means a factor 4 reduction in probability of failure and risk; the TOC

decrease is from 0.5 M€ to 0.2 M€.

dr. ir. J.H. den Besten Delft University of Technology

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{CAPEX, OPEX} evaluation:

Considering CAPEX only, costs will increase. The wire+arc additive manufacturing technology is still

immature, but over time production costs are expected to decrease. Cost savings will come from the

OPEX part in terms of fatigue related maintenance.

expenses as-welded

joint

wire+arc

joint

note

CAPEX:

- material and production

0.1 M€

0.35 M€

20 nodes for the jacket are considered

OPEX:

- fatigue related maintenance

0.5 M€

0.15 M€

Total: 0.6 M€ 0.5 M€ at least 15 [%] cost reduction obtained

Table 2. {CAPEX, OPEX} for an asset value of 10 M€.