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

Transcription

1 {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. longitudinal girder (beam) {shell, plate} 1 bay stiffened panel transverse frame (beam) stiffener (beam) beams (tubular members) joints (nodes) 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., 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. 1

2 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 ) and multiplanar K-joint (Dong et al., 2011, Long-term fatigue analysis of multi-planar tubular joints for jackettype offshore wind turbine in time domain, Engineering Structures, vol. 33, pp ). 2

3 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: s f = 2500 [mm], stiffener: bulb profile HP200x10 stiffener spacing: s s = 750 [mm], number of stiffeners: n s = 3 base plate thickness: t b = 16 [mm], number of frames: n f = 2 frame web height: h w = 800 [mm], stiffener area: A s = 2566 [mm 2 ] frame web thickness: t w = 12 [mm], stiffener area moment of inertia: I s = [mm 4 ] frame flange width: w f = 200 [mm], neutral axis position: n a = 119 [mm] frame flange thickness: t f = 15 [mm], stiffener height: h s = 200 [mm] The structure will be exposed to (local) pressure loading p = 10 [ton/m 2 ] = 0.1 [N/mm 2 ]. 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: σ y = 355 [N/mm 2 ], density: ρ s = 7800 [kg/m 3 ] 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: = ! " ' [mm 4 ] 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 ( = ) * = + * = ' [Nmm] 3

4 Nominal bending stress: / = = [MPa]; compressive Sandwich beam bending response: t f h c t f Nominal bending stress: /,7 ( The core height can be optimised (Wijker, 2008, Spacecraft Structures, ISBN: , Springer-Verlag, Berlin Heidelberg, Germany) w.r.t weight based on {strength, stiffness, face sheet dimpling}: 8 9 =: ; < * < =? 7 with C = {2, 4, 2/3} Note that the h c -t f relation is only used to have some guideline in the free design space. Substitution in the bending stress equation provides a t f formulation: 7 =@ ( : /,7 A 7 A 9 The larger C, the smaller the required t f will be. However, considering typical wall thickness constraints in the maritime industry from production, corrosion, and deformation perspective, t f 5 [mm]. In order to define the sandwich dimensions the {stiffened, sandwich} panel response level should be similar (σ b = σ b,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 t f criterion is satisfied: t f 6 [mm]. Consequently: h c 50 [mm]. Average sandwich core shear stress check: B 9 = + * 3 = 3 [MPa] This magnitude is acceptable, but should not be larger considering typical foam shear strength criteria. 4

5 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 m arc = 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 ) available for a butt joint: FAT183 and slope m FSW = 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 = cycles is adopted as resistance criterion: C= D : I F3 3 E G!D H Γ1+ K 8 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 C r ) of the stiffened panel response: C LMN = : "O9 Γ1+K IPQR LMN C "O9 : LMN Γ1+K "O9 ;: O F "O9 G!D? E G!D I S= H F "O9 For h c = 50 [mm], C r = Reducing the core height to respectively h c = 40 mm and h c = 30 [mm] in order to save some more weight, the sandwich panel response increases: C r40 = 1.33 and C r30 = The damage ratios (Fig. 5) show that for h c = 50 the fatigue damage reduces to ~10 [%] of its arcwelded equivalent; for h c = 40 up to ~30 [%]. For h c =30 [mm] no fatigue damage reduction is obtained damage ratio h c = 50 h c = 40 h c = 30 D FSW / D arc most likely max. stress range S in N = cycles [N/mm 2 ] Figure 5: fatigue damage ratio for different sandwich core heights. 5

6 Structural weight: Stiffened panel: T=U! V! 7 7 W + +V! 7 7 W! +! 7! V8 X X +Y 7 7 WZ A 10 4[ T 2150 [kg] Sandwich panel: T=VA A 7 7 W \V! 7 7 W! +! 7! 8 X 8 9 +Y 7 ] 10 4[ T^_ 2060 [kg] ~5 [%] reducton T`_ 1940 [kg] ~10 [%] reducton T a_ 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/m 2, meaning: Stiffened panel: Sandwich panel (~16 m 2 ): 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 t p 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 6

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: FS-welding: 30 /m weld length, meaning: 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 CO 2 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 midsized bulk carrier it can mean an annual saving of and for a large container ship for 1 [%] fuel consumption reduction (ABB.com); in average approximately 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 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, 5 th Int. Symp. on Life-Cycle Civil Engineering, Delft, The Netherlands) and extrapolated to industrial applications. SFA = (spectral) fatigue analysis (for design) HSM = hull structural monitoring (X year data collection) Figure 6: risk versus total cost of ownership (TOC). 7

8 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 h c = 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 h c = 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 n s ) = 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 CAPEX: stiffened panel sandwich panel - material and production 5 M estimate is conservatively rounded OPEX: - fuel saving - fatigue related maintenance 25 M -2 M 8 M note 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. 8

9 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 [%] 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. 9

10 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 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 /hr (technology is still immature) and the production rate at 10 kg/hr; i.e /kg, meaning the T-joint costs become in average 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: 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. 10

11 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 C r ) of the tubular joint response: C N0 C X1cd/9" = : X1cd/9" : N0 Γ1+K N0 ΓV1+K X1cd/9" W ;: O F X1cd/9" G!D I Rffg? E G!D I hijk/= H F X1cd/9" It is expected that using topology optimisation a response reduction of 10 [%] can be obtained: C r = 0.9. Plotting the damage ratios shows that for C r = 1 the fatigue damage reduces to ~25 [%] of its arcwelded equivalents; for C r = 0.9 up to ~15 [%] damage ratio C r = 1.0 C r = D WAAM / D weld,cast most likely max. stress range S in N = cycles [N/mm 2 ] Figure 9: fatigue damage ratio for different C r 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. 11

12 {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 CAPEX: as-welded joint wire+arc joint - 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 note Total: 0.6 M 0.5 M at least 15 [%] cost reduction obtained Table 2. {CAPEX, OPEX} for an asset value of 10 M. 12