# Steel fiber-reinforced slag-based concretes for floats of FOWTs – The case of a slab as a preliminary approach: design, durability under chloride exposure and Life Cycle Assessment

Nicolas Reuge, Stéphanie Bonnet, Laurence Guihéneuf, Olivier Cornet, Anne Ventura

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Nicolas Reuge, Stéphanie Bonnet, Laurence Guihéneuf, Olivier Cornet, Anne Ventura. Steel fiber-reinforced slag-based concretes for floats of FOWTs – The case of a slab as a preliminary approach: design, durability under chloride exposure and Life Cycle Assessment. Case Studies in Construction Materials, 2025, 23, pp.e04824. 10.1016/j.cscm.2025.e04824. hal-05523300

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# Steel fiber-reinforced slag-based concretes for floats of FOWTs – The case of a slab as a preliminary approach: design, durability under chloride exposure and Life Cycle Assessment

Nicolas Reuge<sup>a, *</sup>, Stéphanie Bonnet<sup>a</sup>, Laurence Guihéneuf<sup>b</sup>, Olivier Cornet<sup>c</sup>,
Anne Ventura<sup>d</sup>

## ARTICLE INFO

### Keywords:

Steel fiber-reinforced concrete

Blast furnace slag

Durability

Chloride

Corrosion

Life cycle assessment

## ABSTRACT

Concrete Floating wind turbine is a solution of the future for Marine Renewable Energies (MRE). The B2FE project, i.e., steel fiber-reinforced concrete (SFRC) for wind turbine floats, aims at developing and characterizing a slag-based concrete (low-carbon concrete) reinforced with steel fibers (slag-based SFRC). Four concrete formulations composed of 50 % CEM I and 50 % blast furnace slag, including fiber contents of 0 (i.e., no fiber), 20, 40 and 60 kg.m⁻³, are studied. A simplified case study is investigated by considering a slab under the same exposure conditions as a FOWT. Considering given mechanical requirements (distributed load of 100 kPa, i.e. 10 t.m⁻²), the dimensioning of the slab is carried out with, on one hand, the slag-based RC solution (no fiber) and on the other hand, the three SFRC formulations (no rebar). Compared to steel bar reinforcements, it appears that steel fibers can provide similar mechanical properties with fewer concrete or steel which is a very interesting result. Then, knowing the physicochemical properties of the materials, simulations of chloride ions transport in the slab subjected to sea water tidal conditions (known as much "aggressive" as splashing) are performed to estimate the service life of these solutions: they appear to be around 30 years, which meet the expectations. Finally, a cradle-to-gate Life Cycle Analysis (LCA) demonstrates a clear advantage of fiber-reinforced formulations, by reducing carbon footprint and other indicators of environmental impacts. In terms of service life and carbon footprint, an innovative cross-analysis leads to an optimal steel fiber content of 40 kg.m⁻³.

## 1. Introduction

In marine environments, Reinforced Concretes (RC) are exposed to sea water and subjected to chloride ions penetration [1]. When chlorides reach the steel rebars in a sufficient quantity, corrosion of steel initiates [2], finally leading to severe cracking of the concrete matrix and radical loss of mechanical properties of the structure.

It is now well established that replacing a part of the cement by Blast-Furnace Slag (BFS) in the concrete formulation modifies the

<footnote><sup>a</sup> Nantes Université, Ecole Centrale Nantes, CNRS, GeM, UMR 6183, Saint-Nazaire F-44600, France</footnote>

<footnote><sup>b</sup> CAPACITÉS SAS, Nantes Université, 16 rue des Marchandises, Nantes F-44200, France</footnote>

<footnote><sup>c</sup> Charier TP, 2 bis rue des meuniers, Couëron F-44220, France</footnote>

<footnote><sup>d</sup> Univ Gustave Eiffel, MAST-GPEM, Campus de Nantes, Allée des Ponts et Chaussées, CS5004, Bouguenais cedex F-44344, France</footnote>

<footnote>* Corresponding author.</footnote>

<footnote>E-mail address: reuge@free.fr (N. Reuge).</footnote>
porous network of the material [3] in a very effective way to slow down chloride ingress. As a matter of fact, chloride diffusivity in slag-based concretes can be 5 times lower than in ordinary concretes with the same total open porosity [4]: this can be explained by the refinement of the porous network during the curing period due to the hydraulic latent hydration coupled with pozzolanic activity [5, 6]. Moreover, being a byproduct, BFS is a low-carbon binder.

Today, some new types of concretes such as Steel Fiber-Reinforced Concretes (SFRCs) are proposed to improve mechanical strength. It is established that the presence of steel fibers in the concrete allows significant gains in mechanical properties allowing to get rid of all or part of the rebars and to significantly reduce thicknesses of the walls [7]. Nevertheless, corrosion resistance of SFRCs is questionable: this is not a priori obvious for the common sense since steel fibers can be subjected to corrosion as well. However, from some results of the literature, it seems beneficial [8–13]. In Poorsaheli et al. [14], concrete beams (based on CEM II cement) featuring different contents of steel fiber were immersed in a tidal zone of the Persian Gulf. They showed that the onset corrosion of steel fibers was at about 26 months and 17 months for 0.33 % and 0.99 % fiber volume contents respectively. Actually, it is logical that the most exposed fibers (close to the surfaces of the beams) begin to corrode rapidly but this does not bring any indication on the actual durability of the structure. Kim et al. [8] studied slag-based SFRCs (fixed fiber volume fraction of 0.5 %) and various slag mass contents (30 % and 50 % of the binder). From their tests of long-term marine immersion, the fiber corrosion rate decreases as a function of the slag content. Moreover, assessments of service life using Life-365 model lead to 56 years and 161 years for slag content of 30 % and 50 % respectively. However, they did not take into account the modifications of the transport properties due to corrosion of fibers.

Furthermore, for fibers having very small diameters (usually ≤ 1 mm), the corrosion products would not involve as large cracks as corroded rebars. For reasons of durability or containment, crack opening is often limited to a threshold value to prevent the penetration of aggressive agents. In this case, the addition of fibers is an advantage: the post-peak behavior of the concrete is improved, providing residual tensile strength in the concrete, which is usually neglected in the case of non-fiber-reinforced concretes. This makes it possible to reduce the amount of rebar reinforcements [15]. This is very promising for SFRC structures in marine environment if the concretes are well formulated.

Excepted the aforementioned studies [8,14], which give only partial answers, there is clearly a lack in literature about durability of
steel fiber-reinforced concretes for coastal structures. Furthermore, steel fibers are studied here because, beyond their recognized
qualities, there is no European standard available yet for other types of fibers (such as glass or synthetic fibers).

Maximizing service life of concrete structures is a goal, but nowadays, reducing carbon footprint is of great concern too. From the review of Manso-Morato et al. [16], using steel fibers as reinforcements can avoid the energy-intensive works related to laying out steel reinforcement but increases the energy in terms of the manufacturing process of these fibers, affecting the LCA. They emphasize that the studied articles defined their own Functional Unit (FU) and system boundaries of the LCA, making comparisons difficult. Overall, they recommended the use of recycled steel fibers. However, from [17], the use of recycled steel fibers causes a reduction in the post-cracking behavior of the concretes in bending and consequently reduced their energy absorption capacity. In [18], the authors studied a slab and showed that using the combination of a low-carbon binder (limestone and calcined clay) and steel fiber as reinforcements could achieve carbonation resistance purposes and save 10 % of the carbon footprint of the structure compared to traditional RC. Adjustments of the concrete mix for lower binder content and the reuse of steel fibers could increase the relative decarbonization potential to 40 %.

In [19], the authors performed the LCA of beams made of RC/SFRC. They considered concretes withstanding XC3 exposure class (i. e., moderate humidity environment) and fiber contents between 20 kg.m⁻³ and 40 kg.m⁻³ for SFRCs. In this study, corrosion is induced by cracks resulting from carbonation. The functional unit (FU) was defined to address given bending moments and a service life of 100 years. From their results, the average embodied carbon of SFRC beams were about 37 % lower than that of the conventional

<figure><img src="image_5.png" /><figcaption>Fig. 1. The Floatgen demonstrator.</figcaption></figure>
RC beams: this was mainly attributed to the material-saving properties of the steel fibers. Indeed, addition of steel fibers as reinforcements could reduce the average design depth by 17%.

To our knowledge, no research study was performed on SFRC structures exposed to sea environment by considering both durability and LCA for a functional unit based on experimental scenarios: this is what is proposed here.

It is important to note that most floats of FOWTs are entirely made of steel nowadays. It is clear that replacing a steel structure by a RC or SFRC would be a big step in the good direction in terms of costs and carbon emissions. This was the main purposes of BW Ideol™ in developing the Floatgen demonstrator which is shown in Fig. 1.

The aim of the project B2FE (French acronym, fiber-reinforced concrete for floats of floating wind turbines) is to study the solution of a slag-based SFRC that would be used for floats of Floating Offshore Wind Turbines (FOWTs) [20]. The aim of the project is actually twofold: (1) maximizing the service life of floats and (2) obtaining a favorable Life Cycle Assessment (LCA). As aforementioned, reducing the carbon footprint of concrete structures becomes of great concern. But as a matter of fact, LCA must be regarded against service life to makes sense [21]. A priori, it cannot be stated that these two points will be favorable using SFRCs, this is why it has to be investigated. Fig. 2 is a diagram presenting the studied approach.

The 1st point has been partly investigated in Reuge et al. [20]: measurements of mechanical strengths and durability indicators in the studied slag-based SFRCs have been carried out as a function of the steel fiber content in the concrete formulation (0, 20, 40 and 60 kg.m⁻³). Rapid Chloride Migration Tests (RCMTs) have revealed an increase of the chloride ions diffusivity as a function of the fiber content, very slightly between 0 and 40 kg.m⁻³ and more significantly between 40 kg.m⁻³ and 60 kg.m⁻³. Moreover, still in [20], an appropriate multiscale homogenization scheme has been developed, allowing to understand well the role of the fiber content on the evolutions of chloride diffusivity and electrical resistivity.

The multi-criteria approach developed here will allow to couple LCA and durability as the inputs are identical: the concrete mixtures. A study case will consider the specific case of a slab. It will bring useful information which could be transposed to floats of FOWTs.

The 1st point will be complemented performing simulations of chloride ingress in the studied slab with a face subjected to sea water exposure like floats of FOWTs. From on-site inspections, it is well established that exposure of concrete structure to tidal or splashing conditions is very harmful [22–24]. In conditions of immersion, chlorides are transported by water, they penetrate deeper and deeper into the porosity and can bind to the material. In conditions of emersion, bound chlorides progressively accumulate. Cycle after cycle, their quantity progressively increases to a critical concentration allowing the initiation of corrosion of the steel components [25]. Offshore splash conditions are severe but tidal cycles are known as being at least and maybe more "aggressive" [26]. Since splashing involves very steep conditions difficultly handled by numerical solvers, simulations of exposure to tidal conditions will be performed, they will provide some service life estimations of the studied slab in severe conditions of exposition.

The 2nd point, has to be investigated regarding the 1st point. According to the allocation rules set by the LCA standard NF EN 15804 [27], the use of secondary materials recovered from waste, such as cement blended with BSF, should decrease environmental impacts compared to concrete made with Ordinary Portland Cement (OPC). However, steel rebars and steel fibers production processes are different, and thus the environmental impacts need to be quantified and compared taking into account the quantities of materials used (concrete and steel) as a function of the rebar/fiber contents on the basis of the case study (i.e., the slab). Thus, by combining the service life estimations provided by the simulations with LCA, an original view of the studied solutions will appear.

First, a preliminary step will examine the dimensioning of the slab (made of slag-based RC or SFRC with fiber contents of 20, 40 or 60 kg.m⁻³) to meet given mechanical requirements, playing with its thickness. Then, durability performances of the slab will be

<figure><img src="image_6.png" /><figcaption>Fig. 2. Schematic diagram of the case study.</figcaption></figure>
studied. Mass transport properties of the studied materials will be completed from simulations adjusted on experiments of drying/imbibition. Subsequently, simulations of chloride ingress in the materials subjected to tidal conditions (in substitution of steep splashing conditions) will be performed considering the slag-based RC or the SFRC with the studied fiber contents. Finally, LCA will be performed for the slab made of slag-based RC or SFRC as a function of the fiber contents. All these results will allow to sort the proposed solutions through an innovative cross-analysis and to initiate a discussion transposing the case of a concrete slab to the case of a concrete wind turbine float.

## 2. Case study

This preliminary step examines the dimensioning of a slab (made of slag-based RC or SFRC with fiber contents of 20, 40 or 60 kg. m⁻³) to meet given mechanical requirements, playing with its thickness. This slab is intended to be fabricated near Saint-Nazaire (FR) in a marine environment and therefore exposed to severe conditions of chloride ingress. The expected service life of the structure is of 30 years.

### 2.1. Concrete formulations

Formulation of the reference concrete called F0 (no fiber) is given in Table 1. The binder is composed of 50 % OPC (reference: CEM I 52.5 N SR5 PM-CP2 HTS) and 50 % BFS provided by Ecocem™ (Dunkerque, France). Sand and gravels are provided by Groupe CB™ (quarry of the Boulonnais, France). The water to binder mass ratio W/B is of 0.48. Physical and chemical properties of OPC and BFS are given in Table 2. BFS is obtained by quenching the slag in water at the outlet of the Dunkerque blast furnaces using a new-generation granulator (INBA®), with a mean glass fraction above 90 %.

This concrete respect the requirements of XS3 (EN 206.1 standard) exposure conditions (tidal and splash conditions, sea spray) as required for floats of FOWTs.

Steel fibers are Dramix 5D from Bekaert™. A specimen is shown in Fig. 3: this is a double-hooked fiber of 0.9 mm diameter and of about 63 mm length. These hooked fibers are known to be particularly interesting to provide a good ductility to the concrete [28]. Physical properties of these fibers are given in Table 3.

As shown in Table 4, from the reference formulation F0 of the concrete (Table 1), three additional formulations are considered: F20, F40 and F60 with fiber contents of 20 kg.m⁻³, 40 kg.m⁻³ and 60 kg.m⁻³ respectively. These formulations have been validated by our industrial partner Charier™ and by the fiber provider Bekaert™. The fiber content of 60 kg.m⁻³ is the maximum that can be considered with regards to the workability of the fresh concrete and efficiency of the mixing process to avoid the formation of so-called "hedgehogs" or "urchins" of fibers [20].

Mechanical measurements (see Rc and Rt at day 28 in Table 4) have shown that F20 was not significantly more resistant than F0 whereas quite significant improvements have been obtained with F40 and F60: for Rc, + 9.3 % from F0/F20 to F40, + 6 % from F40 to F60 and for Rt, + 34 % from F0/F20 to F40, + 20.9 % from F40 to F60. Young modulus E of the concretes has been measured and revealed to be equal to about 41 ± 1 GPa whatever the fiber content. These mechanical characteristics will be used as input data for the mechanical dimensioning of the studied structure.

### 2.2. Geometry of the slab and mechanical requirements

A slab subjected to XS3 exposure conditions was considered here instead of a float. Clearly, the mechanical requirements of a slab and of a float are different, but as a first step, the coupled approach – mechanical requirements / LCA – is applied to a slab.

The slab must be able to support a distributed load of 100 kPa (i.e., 10 t.m⁻²). The reference scenario consists in concrete based on F0 formulation reinforced by steel rebars (RC F0). For this reference scenario, a slab thickness of 20 cm was imposed, and dimensioning consisted of determining the quantity of reinforcement.

For the scenarios based on fiber-reinforced concretes, i.e., SFRC F20, SFRC F40 and SFRC F60 (no rebar), dimensioning consisted of determining the slab thickness required to obtain mechanical properties identical to the reference scenario.

<table><caption>**Table 1**

Constituents of the reference concrete F0.</caption><thead><tr><td>Constituents</td><td>Mass content (kg.m<sup>-3</sup>)</td><td>Volume Fraction (%)</td></tr></thead><tbody><tr><td>Gravels, 12 mm – 24 mm</td><td>561.2</td><td>21.6</td></tr><tr><td>Gravels, 4 mm – 12 mm</td><td>432.0</td><td>16.7</td></tr><tr><td>Sand, ≤ 4 mm</td><td>867.8</td><td>33.1</td></tr><tr><td>Total Aggregates</td><td>1861.0</td><td>71.5</td></tr><tr><td>CEM I</td><td>173.5</td><td>5.6</td></tr><tr><td>BFS</td><td>173.5</td><td>6.1</td></tr><tr><td>Total Binder</td><td>347.0</td><td>11.8</td></tr><tr><td>Water</td><td>166.6</td><td>16.7</td></tr><tr><td>Total</td><td>2374.0</td><td>100</td></tr></tbody></table>
<table><caption>**Table 2**

Physical properties and chemical compounds of OPC and BFS.</caption><thead><tr><th></th><th>OPC</th><th>BFS</th></tr></thead><tbody><tr><td><strong>• Property</strong></td><td></td><td></td></tr><tr><td>Specific gravity (g/cm<sup>3</sup>)</td><td>3.17</td><td>2.9</td></tr><tr><td>Specific surface area (cm<sup>2</sup>/g)</td><td>3635</td><td>4450</td></tr><tr><td>Loss of ignition (950°C)</td><td>1.4</td><td>&lt; 1.5 %</td></tr><tr><td>Passing at 32 µm</td><td>–</td><td>≥ 0.5 %</td></tr><tr><td><strong>• % of components</strong></td><td></td><td></td></tr><tr><td>SiO<sub>2</sub></td><td>23.2</td><td>34.9</td></tr><tr><td>Al<sub>2</sub>O<sub>3</sub></td><td>3.1</td><td>10.9</td></tr><tr><td>Fe<sub>2</sub>O<sub>3</sub></td><td>2.24</td><td>0.5</td></tr><tr><td>CaO</td><td>66.4</td><td>42.3</td></tr><tr><td>MgO</td><td>0.76</td><td>6.9</td></tr><tr><td>SO<sub>3</sub></td><td>2.22</td><td>0.1</td></tr><tr><td>K<sub>2</sub>O</td><td>0.18</td><td>0.45</td></tr><tr><td>Na<sub>2</sub>O</td><td>0.11</td><td>0.35</td></tr><tr><td>TiO<sub>2</sub></td><td>0.15</td><td>0.7</td></tr></tbody></table>

<figure><img src="image_7.png" /><figcaption>**Fig. 3.** Double-hooked steel fiber, "Dramix 5D" from Bekaert™.</figcaption></figure>

<table><caption>**Table 3**

Physical properties of steel fibers.</caption><thead><tr><th>Property</th><th>Steel fiber</th></tr></thead><tbody><tr><td>Specific gravity (g/cm<sup>3</sup>)</td><td>7.8</td></tr><tr><td>Length / diameter (-)</td><td>65</td></tr><tr><td>Tensile strength (Mpa)</td><td>2300</td></tr><tr><td>Young Modulus (GPa)</td><td>200</td></tr></tbody></table>

<table><caption>**Table 4**

The four formulations with different steel fiber contents.</caption><thead><tr><th>Formulation</th><th>Fiber content (kg.m<sup>-3</sup>)</th><th>Fiber volume fraction (%)</th><th>Superplasticizer content (kg.m<sup>-3</sup>)</th><th>Open porosity at day 28 (%)</th><th>Rc – Rt at day 28 (MPa)</th></tr></thead><tbody><tr><td>F0</td><td>0</td><td>0</td><td>2.37</td><td>11.06</td><td>54.7 – 5</td></tr><tr><td>F20</td><td>20</td><td>0.256</td><td>2.44</td><td>11.66</td><td>54.8 – 5</td></tr><tr><td>F40</td><td>40</td><td>0.513</td><td>3.00</td><td>10.98</td><td>59.9 – 6.7</td></tr><tr><td>F60</td><td>60</td><td>0.769</td><td>3.95</td><td>10.66</td><td>63.5 – 8.1</td></tr></tbody></table>

## 2.3. Dimensioning

Here, a slab-on-grade is considered. The dimensioning was carried out using the expertise and numerical tools of Charier TP™ based on two standards: EN 1992-1-1 [29] for RC and fib Model Code 2010 [15] for SFRCs. The mechanical properties measured in Section 2.1 were used (see Table 4).

In order to be able to support a distributed load of 100 kPa, the slab must be able to withstand a maximum bending moment SLS (Serviceability Limit State) of 38 kN per linear meter in both horizontal directions.

For the reference scenario RC F0 (slab thickness of 20 cm), taking into account the aforementioned SLS moment criterion, calculations lead to the following quantities of reinforcing steel: 8.27 cm² per linear meter and 2.57 cm² per linear meter in both directions in the lower and upper part of the slab respectively. On this basis, the chosen reinforcements are the following:
- top layer: welded mesh ST25, 2.57 cm² per linear meter,

- bottom layer: HA14 e = 18 cm, 8.55 cm² per linear meter.

This leads to a steel content equal to 110 kg.m⁻³, and thus for the 20 cm thickness, to a surface quantity of steel of 22 kg.m⁻², as reported in Table 5.

For the SFRC F20, F40 and F60 scenarios, calculations were carried out in accordance with the theory of break lines using the "Slab on Ground" tool developed by Bekaert™ [30] and based on the fib Model Code 2010 standard [15]. Thus, typical values were used for the following parameters:

- characteristic coefficient: G = 1,

- coefficient of fibers orientation: K₀ = 1,

- long term effects of the concrete: αcc = 0.85,

- safety coefficient on fibers: γf = 1.

Moreover, finite elements calculations were performed using the software SCIA® (according a linear elastic approach).

Calculations lead to slab thicknesses of 23.5 cm, 18 cm and 15.5 cm for the respective SFRC F20, F40 and F60 scenarios, corresponding to respective surface quantities of steel of 4.7 kg.m⁻², 7.2 kg.m⁻² and 9.3 kg.m⁻². The results obtained for the four scenarios are reported in Table 5.

This dimensioning procedure clearly shows that all SFRC scenarios (F20, F40 and F60) reduce the amount of steel compared to RC F0, and that SFRC F40 and F60 also make it possible to use less concrete in significant proportions.

These significant differences in terms of steel quantities required for RC and SFRCs scenarios can be explained partly by the fact that calculations rely on different standards (EN 1992 vs. Model Code 2010) and mostly by the fact that tensile strengths of the steel of the rebars / the steel of the fibers are very significantly different: 500 Mpa / 2300 Mpa respectively.

### 3. Durability performances

In this part, durability of the slab under chloride exposure is investigated numerically. It involves knowledge of the transport properties of the studied concretes, the use of mass transport models and governing equations. First, mass transport properties of the studied materials are completed from simulations adjusted on experiments of drying/imbibition and then, simulations of chloride ingress under tidal conditions are performed.

#### 3.1. Governing equations

In presence of gas (air and water vapor) and liquid water, mass balance equations in 1D axial direction are given in [31]. They involve Darcy flows ruled by liquid and gas pressures, Pᵢ and Pⱼ respectively, intrinsic permeability of the porous network to liquid and gas, Kᵢ and Kⱼ respectively, and the corresponding relative permeabilities kᵣ,ᵢ and kᵣ,ⱼ. These governing equations were used here to simulate samples drying and imbibition.

Under tidal conditions, ambient RH always remains higher than 50 % (RH varying between 80 % and 100 %), thus gas transport can be ignored in the description of the hygric transport [31]. By adding transport of chlorides, Reuge et al. [32] have shown that this simplification remains valid. Thus, mass balance equations in molar form can be written as follows:

$$ \frac{\partial}{\partial t}(\varphi S_i) - \frac{\partial}{\partial x}\left(\frac{K_i k_{r,l}}{\eta_l} \frac{\partial P_l}{\partial x}\right) = 0 \qquad (1) $$

$$ \frac{\partial}{\partial t}(\varphi S_i c_{Cl,f}) + C_{Cl,b} - \frac{\partial}{\partial x}\left(D_{eff} \frac{\partial c_{Cl,f}}{\partial x}\right) - \frac{\partial}{\partial x}\left(c_{Cl,f} \frac{K_i k_{r,l}}{\eta_l} \frac{\partial P_l}{\partial x}\right) = 0 \qquad (2) $$

where the degree of saturation Sᵢ is the volumetric ratio of liquid water in the open porosity φ, cCl,f is the free Cl ions concentration (mol.m⁻³ of solution) and CCl,b is the bound Cl concentration (mol.m⁻³ of concrete). Deff is the effective diffusivity of free chloride ions in the liquid phase of an unsaturated porous medium.

Eqs. (1) and (2) describe the transport of liquid water and chlorides respectively. Some approaches considering the presence of other ions have been developed previously [33,34] but the interaction of chlorides with other ions seems to be of minor importance compared to the other transport mechanisms.

<table><caption>**Table 5**
Concrete and steel contents calculated by dimensioning the slab.</caption><thead><tr><th>Mix design scenario</th><th>Fiber content (kg.m<sup>-3</sup>)</th><th>Rebar content (kg.m<sup>-3</sup>)</th><th>Slab thickness (cm)</th><th>Steel content (kg.m<sup>-2</sup> of slab)</th><th>Concr. content (kg.m<sup>-2</sup> of slab)</th></tr></thead><tbody><tr><td>RC F0</td><td>0</td><td>110</td><td>20</td><td>22</td><td>490</td></tr><tr><td>SFRC F20</td><td>20</td><td>0</td><td>23.5</td><td>4.7</td><td>576</td></tr><tr><td>SFRC F40</td><td>40</td><td>0</td><td>18</td><td>7.2</td><td>441</td></tr><tr><td>SFRC F60</td><td>60</td><td>0</td><td>15.5</td><td>9.3</td><td>380</td></tr></tbody></table>
## 3.2. Mass transport models

Only mass transport models for the liquid phase and the chloride ions will be given here. Regarding the gas phase, see [31].

In the materials, $S_l$ as a function of local relative humidity $H_r$ is commonly described by the Van Genuchen (VG) model [35] which has been used by many authors to describe the moisture transport in concretes [31, 36–38]. It is expressed as follows:

$$S_l(Hr) = \left[ 1 + \left( -R T M_w^{-1} a^{-1} \cdot \ln(Hr) \right)^{1/(1-m)} \right]^{-m} \quad (3)$$

Parameters $a$ (Pa) and $m$ (-) are the VG coefficients, adjustable on measurements. Note that they can have significantly different values in adsorption and desorption conditions because adsorption and desorption isotherms differ.

The relation between liquid and gas total pressures is given by:

$$P_l = P_g - P_c(S_l) \quad (4)$$

where $P_c$ is capillary pressure which can be well estimated by the following equation involving the VG coefficients from Eq. (3) and Kelvin's law [35]:

$$P_c(S_l) = a(S_l^{-1/m} - 1)^{1-m} \quad (5)$$

Still according to Van Genuchten [35], liquid relative permeability can be estimated by the following semi-empirical model:

$$k_{r,ads}(S_l) = S_l^{1/2} \left[ 1 - \left( 1 - S_l^{1/m} \right)^m \right]^2 \quad (6)$$

This expression of $k_{r,l}$ is only valid in adsorption conditions, there is no analytical relationship available in desorption conditions (this will be discussed later).

The effective diffusivity of free chloride ions $D_{eff}$ in the liquid phase of an unsaturated porous medium can be modeled as follows [39]:

$$D_{eff}(S_l)/D_{eff,0} = S_l^{\lambda} \quad (7)$$

where $D_{eff,0}$ is the Cl ions diffusivity in the saturated porous medium (which can be determined by migration in steady state [40]) and $\lambda$ is an adjustable exponent. From adjustments of simulations [39] on experiments [41], a value of 6 has been obtained for $\lambda$. However, it appears that the global shape of the data of Francy and François [41] can be better described by the following model:

$$\begin{cases} D_{eff}(S_l)/D_{eff,0} = 2.25 \cdot 10^{-3} \exp(6.6S_l) & \text{for } S_l < 0.924 \\ D_{eff}(S_l)/D_{eff,0} = 1 & \text{for } S_l \ge 0.924 \end{cases} \quad (8)$$

Thus, model (8) will be used here.

From Gelhen et al. [42], chlorides diffusion in concrete is not constant and decreases with time. This would be due to the structural evolution of the porous network at long time scales. Based on some on-site measurements, they propose a model which can be transposed (from apparent diffusion coefficient to effective diffusion coefficient) as follows:

$$D_{eff,0}(t) = D_{eff,0}(t_{cure}) \left( \frac{t_{cure}}{t + t_{cure}} \right)^{\alpha} \quad (9)$$

where $t_{cure}$ (days) is the duration of the cure period, $D_{eff,0}$ is deduced from measurements at $t_{cure}$, $t$ (days) is the time of service (i.e., $t = 0$ at $t_{cure}$) in days and $\alpha$ is the so-called ageing factor having a value between 0.2 and 0.6 [42].

Presently, all the "engineering" models used by the designers take the ageing factor into account (see fib Model Code described in Gelhen et al. [42] and Model Life-365 [43]).

Finally, $c_{Cl,f}$ and $C_{Cl,b}$ are linked by binding isotherms, which can be described by the following relationship:

$$C_{Cl,b}(c_{Cl,f}, S_l) = \mu(c_{Cl,f})^{\gamma} \quad (10)$$

where $\mu$ and $\gamma$ are the Freundlich parameters.

<table><caption>**Table 6**
Mass transport properties of the reference concrete (F0).</caption><thead><tr><th>VG param.<br>a<sub>ads</sub> - a<sub>des</sub> (MPa)</th><th>VG param.<br>m (-)</th><th>Freundlich param.<br>μ (SI)</th><th>Freundlich param.<br>γ (-)</th></tr></thead><tbody><tr><td>9.97<sup>*</sup> - 24.93</td><td>0.46<sup>*</sup></td><td>11.59<sup>*</sup></td><td>0.536<sup>*</sup></td></tr></tbody></table>

<footnote>* Deduced from measurements of Ben Fraj et al. [4]</footnote>
### 3.3. Mass transport properties

Ben Fraj et al. [4] have studied a slag-based concrete (BSFC60) whose formulation is identical to the F0 formulation considered here. By adjustments on their measurements, the VG parameters *a* and *m* in adsorption conditions and the Freundlich parameters can be deduced, as reported in Table 6.

Adjusted isotherms can be compared to measurements in Fig. 4a (water sorption) and 4b (binding of Cl ions). Because of the relatively small volume fractions of steel fibers in the SFRC studied here (≤ 0.77 %), the values of these parameters will be kept unchanged whatever the fiber content. The isotherm of desorption is a priori not known, but in agreement with some studies [39,44], *m* has been kept equal to 0.46 whereas *a* has been multiplied by 2.5 to get *a*<sub>des</sub> (see Table 6 and Fig. 4a).

Thanks to RCMTs and multiscale homogenization modeling, Reuge et al. [20] have determined *D*<sub>eff,0</sub> as a function of the fiber content. Results are presented in Fig. 4c, with the results of measurements and, for the calculations, hypothetical interphase of corroded steel around the fibers of thickness Δ*r*<sub>int</sub>. Corrosion was induced by the testing method itself (i.e., application of a voltage difference between the two faces of the samples). This interphase is considered as an open porosity. For the basic formulation without fiber, the measured value of *D*<sub>eff,0</sub> is of 1.21·10<sup>-12</sup> m<sup>2</sup>.s<sup>-1</sup>, which is typical for a slag-based concrete. From the tests and the calculations, from F0 to F40 formulations, *D*<sub>eff</sub> increases according an interphase thickness Δ*r*<sub>int</sub> of 22 µm, which means that the corroded fraction τ of the fibers is of about 10 %. Fig. 4c also shows the results for a hypothetical corroded fraction τ equal to 0 % (Δ*r*<sub>int</sub> = 0, *D*<sub>eff</sub> remains almost constant and equal to *D*<sub>eff,0</sub>) and τ equal to 50 % (Δ*r*<sub>int</sub> = 128 µm, for F40, *D*<sub>eff</sub> is equal to 2.31·10<sup>-12</sup> m<sup>2</sup>.s<sup>-1</sup>). These data will be used for the simulations of Cl ingress under tidal conditions.

At this point, regarding liquid transport properties, some elements are still missing:

- the liquid intrinsic permeabilities *K*<sub>l</sub>

- a law for the liquid relative permeability of desorption *k*<sub>r,l,des</sub>

They will be determined in the next section.

<figure><img src="image_8.png" /><figcaption>**Fig. 4.** Physicochemical properties of the concrete – (a) Isotherms of adsorption/desorption, measurements, VG model and adjustment parameters – (b) Binding isotherm, measurements, Freundlich model and adjustment parameters – (c) Effective diffusivity of Cl ions as a function of the fiber content and the degree of corrosion, measurements, calculations [20].</figcaption></figure>
## 3.4. Drying and imbibition

### 3.4.1. Experimental data

After 28 days of moist cure, cylindrical specimens of 22 cm height and 11 cm diameter were cut to obtain samples of 5 cm height and 11 cm diameter. Then, the radial sides of the samples were sealed using resin. Therefore, water exchange with outdoor environment was only possible in the axial direction. For each of the four formulations, two samples were prepared.

Due to the moist cure, samples porosity was saturated with water. The first step consisted in drying: the sample were placed in an oven at a temperature of 60 °C and relative humidity of about 5 %. The samples were weighted regularly to follow water losses. Fig. 5a shows the evolutions of the temporal evolutions of water losses (the ratio of the volume of water loss on sample volume). Each curve is an average of the measurements performed on two samples. The stabilizations of the masses of the samples were obtained after 4–5 months of drying: this indicated a quasi-complete evaporation of water initially present in the samples. From Fig. 5a, drying of F0 and F20 samples were faster than drying of F40 and F60 samples. The final ratios (volume of water loss on sample volume) indicates that porosity of the samples is in the range of 10.6–11.6 %, which leads to an average of 11.1 %.

At this point, as 2nd step, the samples were subjected to gas permeability measurements. In order to do so, a gas permeameter has been used. Despite the inherent uncertainties of this kind of measurements, the gas intrinsic permeability $K_g$ was estimated to about $10^{-16}$ m² as reported in Table 6, with no clear effect on the presence or absence of fibers.

The 3rd step consisted of plunging the dried samples in freshwater at ambient temperature (about 25 °C). The samples were weighted regularly to follow the water gains. Fig. 5b shows the evolutions of the temporal evolutions of the water gains (the ratio of the volume of water gain on sample volume). The stabilizations of the masses of the samples were obtained after 3–7 days of imbibition. From Fig. 4b, imbibition of F0 and F20 samples were clearly faster than imbibition of F40 and F60 samples.

Therefore, these results clearly show that the presence of fibers tends to slow down the water mass transport.

### 3.4.2. Simulations

Mass transport Eq. (1D) [31] and constitutive relationships have been implemented in a finite element solver platform called FlexPDE®. Its solver is based on a modified Newton-Raphson iteration procedure. Timestep and mesh density are adaptative.

Simulations of imbibition were performed first because $k_{r,l,ads}(S_l)$ is known (Eq. (6)). A priori, the only missing parameter is $K_l$. For every formulation, $K_l$ was adjusted in order to obtain the best possible agreement of the results on the measurements.

For these simulations, an initial $Hr$ of 5 % was considered inside the domain ($0 < x \le 5$ cm). At $x = 0$, a value of $Hr$ of 100 % was applied. At $x = 5$ cm, no flux boundary conditions were applied.

In the sake of clarity, results are reported in Fig. 6a only for F0 and F40. For each formulation, Table 7 reports the adjusted values of $K_l$ and the coefficients of determination ($r^2$) obtained.

It is noteworthy that, from F0-F20 to F40-F60, the deduced value of $K_l$ decreases by a factor of about 2. It is consistent with many studies which shows that $K_l$ tends to decrease with the fiber content (see the dedicated paragraph in [20]).

Then, simulations of drying were performed. For these simulations, an initial $Hr$ of 100 % (which means a porosity fully saturated with liquid water) was considered inside the domain ($0 < x \le 5$ cm). At $x = 0$, a value of $Hr$ of 5 % was applied, it corresponds to the estimated relative humidity in the oven at 60°C. They were performed using the values of $K_l$ determined above for every formulation.

The laws of $k_{r,l,des}(S_l)$ being unknown, an analytical expression derived from previous studies [45,46] was used to describe $k_{rl,des}$:

$$k_{r,l,des}(S_l) = \frac{1}{3}(2 - 0.7/S_l^{0.315})\exp\left[\beta_0\left(\frac{S_l - 1}{1 + 2.34S_l^{0.05} - S_l^{\beta_z}}\right)\right] \quad (11)$$

<figure><img src="image_9.png" /><figcaption>Fig. 5. Temporal results of drying/imbibition for the four formulations – (a) Relative volume of evaporated water during drying – (b) Relative volume of adsorbed water during imbibition.</figcaption></figure>
<figure><img src="image_10.png" /><figcaption>**Fig. 6.** Comparisons of measurements/calculations for (a) drying, (b) imbibition for F0 and F40 formulations, adjustment of $K_I$ and (c) models of $k_{rl}$ depending on the adjustments parameters $\beta_s$ and $\beta_0$.</figcaption></figure>

<table><caption>**Table 7**

Results of simulations – values of $K_I$ (adjustments) and correlations.</caption><thead><tr><td>Formulation</td><td>$K_I$ (10<sup>-20</sup> m<sup>2</sup>)</td><td>Imbibition – cor. (r<sup>2</sup>)</td></tr></thead><tbody><tr><td>F0</td><td>8</td><td>0.9962</td></tr><tr><td>F20</td><td>7.5</td><td>0.9946</td></tr><tr><td>F40</td><td>4</td><td>0.9789</td></tr><tr><td>F60</td><td>4.25</td><td>0.9760</td></tr></tbody></table>

where $\beta_0$ and $\beta_s$ are adjustable parameters. Broadly, playing with $\beta_s$ allows to adjust the slope of the curve and playing with $\beta_0$ allows to adjust the magnitude of $k_{r,l,des}(S_I)$. The value of $\beta_s$ was fixed in such a way that $k_{r,l,des}(S_I)$ is parallel to $k_{r,l,ads}(S_I)$ for $S_I$ in the range of 0.3–0.9, as suggested by the curves available in [39]: it leads to a value of 2.7. Thus, the adjustments of simulations of drying on

<table><caption>**Table 8**

Results of simulations – values of $\beta_0$ and $\beta_s$ (adjustments) and correlations.</caption><thead><tr><td>Formulation</td><td>$K_I$ (10<sup>-20</sup> m<sup>2</sup>)</td><td>$\beta_s$ (-)</td><td>$\beta_0$ (-)</td><td>Drying – cor. (r<sup>2</sup>)</td></tr></thead><tbody><tr><td>F0</td><td>8</td><td>2.7</td><td>48.35</td><td>0.9998</td></tr><tr><td>F20</td><td>7.5</td><td>2.7</td><td>48.35</td><td>0.9988</td></tr><tr><td>F40</td><td>4</td><td>2.7</td><td>48.35</td><td>0.9979</td></tr><tr><td>F60</td><td>4.25</td><td>2.7</td><td>48.35</td><td>0.9988</td></tr></tbody></table>
experiments were carried out playing on the values of β₀. Actually, in these conditions of drying, simulations revealed to be quite insensitive to βₛ. On the other hand, sorption parameters (a and m) play a more important role than in imbibition.

In the sake of clarity, results are reported in Fig. 6b only for F0 and F40. Table 8 reports the adjusted values of α₀ and αₛ and the coefficients of determination (r²) of the simulations. kᵣ,ᵢ,ads(Sᵢ) and kᵣ,ᵢ,des(Sᵢ) are shown in Fig. 6c.

### 3.5. Simulations of chloride ingress under tidal conditions

Mass transport Eq. (1D) (Eq. (3-6)) and constitutive relationships (Eqs. (8-10)) have been implemented in FlexPDE® [32,45].

The exposed face of the materials was considered alternately immersed/emerged for cycle durations of 6 h/6 h respectively which led to a periodicity of 12 h. In the time range [0, 6 h], a progressive temporal evolution of ambient relative humidity (RH) is given by the following purely theoretical function:

$$RH(t) = Hr(x = 0, t) = RH_{\max} - (RH_{\max} - RH_{\min}) \frac{\tanh[b_1(t - c_1)/t_{month} + a_1] - \tanh(a_1)}{\tanh[b_1(6 - c_1)/t_{month} + a_1] - \tanh(a_1)} \quad (12)$$

with: a₁ = 18.33, b₁ = -5686.67, c₁ = 0.647 h, tmonth = 729.98 h (number of hours in a month of 30.416 days), RHmin was fixed at 80 % (classical RH of ambient air in maritime environments) and RHmax at 99.9 % (conditions of immersion). The symmetrical function is used for temporal evolution of RH in the time range [6 h, 12 h] and this scheme is repeated for the next periods as shown in Fig. 7.

As Sᵢ is used as an unknown, the temporal evolution at the material surface is given by:

$$S_{i}(x = 0, t) = S_{i}[RH(t)] \quad (13)$$

where RH(t) is given by the just aforementioned scheme and Sᵢ(RH) is given by the VG model (Eq. (3)).

In immersed conditions, CClᵢf(x = 0) was fixed at a typical sea water chloride concentration cCl,sea of 598 mol.m⁻³. In emerged conditions, no flux boundary condition was imposed for CClᵢf at x = 0. In the core of the concrete (x = L = 10 cm), no flux boundary conditions were imposed: this corresponds to conditions found in the core of a structural element.

Initial liquid saturation degree inside the material Sᵢ₀ was fixed as high as 0.9 because of the strong humidity of the marine environment. Initial CClᵢf was set to zero.

Inside the domain, total chloride content CClᵢtot is defined as follows:

$$C_{Cl,tot} = \varphi S_{i}c_{Cl,f} + C_{Cl,b} \quad (14)$$

Simulations were stopped at time tcrit for which CClᵢtot reaches the critical concentration defined as follows at 5 cm depth:

$$C_{Cl,tot}(t = t_{crit}, x = 5\text{cm}) = M_{Cl}^{-1}B \cdot C_{crit} \quad (15)$$

where MCl is Cl molar mass (kg.mol⁻¹), B the binder mass content (kg.m⁻³) and Ccrit the critical chloride content (expressed in weight %/binder). According to [25,42], a typical value of 0.4 % was considered for Ccrit. Note that 5 cm corresponds to the cover thickness above the steel rebars in marine environments according to the standard EN 1992 [29] and that this depth is still relevant for fiber-reinforced formulations for two reasons:

- the possibility of comparing Cl ingress in slag-based RC and SFRC at the same position,

- the design of floats of FOWTs made of SFRC (which is under study) might incorporate steel fibers associated with steel rebars (i.e., hybrid solutions for optimal specifications).

For these 1D calculations, computational cost was of about 1 h for 5 years of physical time. As aforementioned, timestep and mesh density are adaptative, which means that they were dynamically adjusted to ensure a high accuracy. In practice, number of nodes

<figure><img src="image_11.png" /><figcaption>Fig. 7. Temporal evolution of RH in tidal conditions, Eq. (12).</figcaption></figure>
remained around 100. Simulation tests performed with a fixed number of nodes of 200 led to the same results. Timesteps evolved in a range of magnitude to 1 s to 1 h. The results of this simulation tool had been compared to many on-site measurements in [45]: overall, the agreements were good.

For numerical reasons (i.e., numerical discontinuities), simulations cannot be carried out using $k_{rl,ads}$ and $S_{l,ads}$ in immersion phases and $k_{rl,des}$ and $S_{l,des}$ in drying phases (physical reality actually relies on complex hysteretic phenomena). For the sorption isotherm, a median value of $a$ (between $a_{des}$ and $a_{ads}$) of 17.45 resulting in a median isotherm was considered. Regarding the liquid relative permeability, the VG law of $k_{rl,ads}$ (Eq. (7)) was considered except for one case. For the ageing factor, values of $\alpha$ of 0 (no ageing) and 0.4 (median value between 0.2 and 0.6 [42]) were considered.

Table 9 shows the values of the parameters considered for the simulations and the results in terms of $t_{crit}$. As a comparison basis, conditions of permanent immersion were considered for two simulations (I0–1 and I0–2). Regarding corroded fraction $\tau$ of the fibers, extreme situations of Fig. 4c were considered with values of 0 % for most simulations or 50 % (resulting a greater $D_{eff}$) for simulation T40–3. Finally, simulation T40–4 was carried out considering logarithmic average between $k_{rl,ads}$ and $k_{rl,des}$.

Considering $\alpha$ equal to 0 (no ageing effect) and F0 formulation, I0–1 (Immersion) and T0–1 (tidal conditions) simulations leads to $t_{crit}$ of 10.3 years and 9.51 years respectively: compared to immersion, applying tidal conditions results in a decrease of $t_{crit}$ of 7.7 %. Regarding F40 formulation, T40–1 simulation leads to $t_{crit}$ of 8.5 years.

Considering $\alpha$ equal to 0.4 and F0 formulation, I0–2 (Immersion) and T0–2 (tidal conditions) simulations leads to $t_{crit}$ of 43.2 years and 33 years respectively: compared to immersion, applying tidal conditions results in a decrease of $t_{crit}$ of 23.6 %. Regarding F40 formulation, T40–2 simulation leads to $t_{crit}$ of 38.4 years: compared to F0, the increase of $t_{crit}$ is due to the lower $K_l$ (by a factor 2). This effect was not observed between T0–1 and T40–1 simulations, therefore it takes time to occur and reflects the complexity of the phenomena involved.

Considering $\tau= 50 \%$, T40–3 simulation leads to $t_{crit}$ of only 13 years: this is due to the much higher $D_{eff,0}$.

Finally, using the logarithmic average between $k_{rl,ads}$ and $k_{rl,des}$, T40–4 simulation leads to $t_{crit}$ of 32.6 years which is lower than the $t_{crit}$ obtained from T40–2.

Overall, results of the simulations are greatly impacted by the ageing factor and by the corrosion degree of fibers $\tau$ (which cannot a priori be known). Thus, these investigations show the great interest in adequately estimating the value of $\alpha$ and the corroded fraction $\tau$ of fibers, this latter parameter actually increasing with time and decreasing with depth in the material.

# 4. Life cycle assessment

## 4.1. Study context and framing

The results provided by mechanical dimensioning of the slab (Table 5) provide amounts of concrete and steel for each studied scenario considering mechanical requirements (i.e. able to support a distributed load of 100 kPa) as the Functional Unit (FU) for a 1 m² surface area. Thus, the FU volume is 1 m² multiplied by the thickness of the slab. Thus, this gives different values of reference flows for each scenario expressed in m³ of concrete slab over: 0.2 m for RC F0, 0.235 m for SFRC 20, 0.18 m for SFRC 40 and 0.155 m for SFRC 60.

The parametric analysis was carried out by using the dedicated numerical tool BETie®. This tool, developed by SNBPE and PwC Ecobilan, allows to produce EPD (Environmental Product Declaration) sheets in conformity to the NF EN 15804 standard [27] and to perform LCA of structures made of ready-mixed concretes produced by French concrete mixer plants with cement and aggregates produced in France taking into account the French electric mix.

## 4.2. The different steps

The parametric analysis takes into account the whole-life cycle of a project: module (A) with production (A1) and transport of materials (A2), and to construction site (A4), utilization (B1), end of life treatment (C3) or landfill (C4) as well as possible waste recovery (D). Fig. 8 [27] shows all the different steps of the LCA.

<table><caption>**Table 9**
Simulations of Cl ingress in concretes exposed to sea water – parameters and results.</caption><thead><tr><th>Simulation</th><th>Formulation</th><th>Hygric conditions</th><th>D<sub>eff,0</sub><br>(10<sup>-12</sup> m<sup>2</sup>.s<sup>-1</sup>)</th><th>α (-)</th><th>K<sub>l</sub><br>(10<sup>-20</sup> m<sup>2</sup>)</th><th>t<sub>crit</sub> (year)</th></tr></thead><tbody><tr><td>I0–1</td><td>F0</td><td>Immersion</td><td>1.214</td><td>0</td><td>8</td><td>10.3</td></tr><tr><td>T0–1</td><td>F0</td><td>Tidal</td><td>1.214</td><td>0</td><td>8</td><td>9.51</td></tr><tr><td>T40–1</td><td>F40</td><td>Tidal</td><td>1.205</td><td>0</td><td>4</td><td>8.5</td></tr><tr><td>I0–2</td><td>F0</td><td>Immersion</td><td>1.214</td><td>0.4</td><td>8</td><td><b>43.2</b></td></tr><tr><td>T0–2</td><td>F0</td><td>Tidal</td><td>1.214</td><td>0.4</td><td>8</td><td><b>33.0</b></td></tr><tr><td>T40–2</td><td>F40</td><td>Tidal</td><td>1.205</td><td>0.4</td><td>4</td><td><b>38.4</b></td></tr><tr><td>T40–3</td><td>F40</td><td>Tidal</td><td>2.314 (τ=50 %)</td><td>0.4</td><td>4</td><td>13.0</td></tr><tr><td>T40–4<sup>*</sup></td><td>F40</td><td>Tidal</td><td>1.205</td><td>0.4</td><td>4</td><td>32.6</td></tr></tbody></table>

<footnote>* Using logarithmic average between $k_{rl,ads}$ and $k_{rl,des}$</footnote>
<figure><img src="image_12.png" /><figcaption>Fig. 8. Building Assessment Modules for LCA according to EN 15804:2012 [27].</figcaption></figure>

Regarding the production (module A), for the construction step (A1), generic data provided by BETie® for steel fibers, super-plasticizer and granulometric classes of sand and gravels were used. The transport of concrete (A4) was considered by road (mixer truck) from Saint-Nazaire to the construction area in marine environment (distance: 7 km). The installation step (A4-A5) takes into account the energy consumed on site and the production and the transport of auxiliary materials (including the rebars).

During the use phase (module B), the slab does not require any maintenance during its service life of 30 years. Carbonation of concretes was taken into account according to NF EN 16757 [47] and results as a negative CO₂ emission during both the utilization (B1) and landfill (C4) phases.

Regarding the end of life (module C), 75 % of the concrete slab was considered as recycled (C3) and 25 % as landfilled (C4). Note that end of life scenarios of BETie® are conform to the concrete PCR (Product Category Rules). For the RC F0 scenario, steel rebars were considered as recyclable, leading to calculation of benefits beyond the life cycle (module D), whereas for SFRC F20, F40 and F60 scenarios, steel fibers were not considered as recyclable. Note that these choices are questionable because: (1) solutions to extract steel fibers from the concrete exist such as the FCB RODAX® 4D unit (selective crushing and grinding as specified by Fives™), (2) with or without fibers, it may not be worth recycling a block full of chlorides.

That said, potential reuse of the materials (module D) is considered as beyond the building life cycle and, therefore, was not taken into account in calculation of indicators of environmental impact.

## 4.3. Results

Fig. 9 shows Global Warming Potential indicator obtained for the different studied scenarios (expressed in kg CO₂ equivalent/FU). It appears that the impacts are mostly due to the production steps. Production of SFRCs is significantly more impactful than production of RC due to the manufacturing process of fibers which is more energy consuming than the one of rebars. Nevertheless, for the RC F0 scenario, the impact of the installation step is really significant, which is not observed for the SFRC F20, F40 and F60 scenarios.

<figure><img src="image_13.png" /><figcaption>Fig. 9. Global warming potential indicator for the studied scenarios.</figcaption></figure>
Actually, regarding rebars, shaping and welding of the meshes are very energy consuming. Overall, global carbon footprint of RC F0 and SFRC F20 are similar. A significant decrease around 15 %-20 % is however obtained for SFRC F40 and F60 scenarios, these two scenarios providing similar results. Note that negative values during landfill (C4) are due to concrete carbonation. Scenario RC F0 has more important benefits beyond the life cycle (module D) due to steel recovery, whereas it is not considered for SFRC scenarios.

Fig. 10 shows relative impact for each indicator and for the four scenarios. Recycling and reuse were only considered for the extraction of rebars at the end of life for RC F0 scenario, which explains null values for those indicators for SFRC scenarios. For other indicators, significant improvements are obtained for SFRC F40 and F60 scenarios, up to 20 % for some indicators, compared to RC F0 and SFRC F20. Regarding the environmental indicators, the most favorable results are obtained with scenario SFRC F40 for: global warming, photochemical ozone creation, exhaustion of abiotic resources (foss.), and air pollution. The most favorable results are obtained with scenario SFRC F60 for: depletion of the ozone layer, soil and water acidification, eutrophication, exhaustion of abiotic resources (el.), and water pollution.

Broadly, these results are in agreement with the literature [18,19] on the fact the replacing steel reinforcement by steel fibers results in more favorable LCA. However, as stated by [16], the two aforementioned studies have defined their own FU making comparisons difficult. Moreover, they focused on durability with regard to the carbonation process and not chloride ingress.

This LCA study has been performed with the tool BETie® which doesn't provide sufficient scope for integrating many innovative aspects. Indeed, this tool is commonly used by companies in France, and provides results in line with current standards. However, to integrate more innovative proposals (here fibers, but it could be recycled aggregates from precast rejects, even plant fibers, biochar...) requires a different, more modular and transparent tool. The current concrete standards specify mechanical performances in the description of the product, and use m³ (or kg) and as a functional unit. Theoretically, in that case study, because concretes with different performances are compared, performances define the functional unit, and should be used to calculate the reference flow (i.e. amount of product necessary to fulfill the FU). This calculation was not possible because there is no existing model. So, this study only explored these performances based on experimental studies and compared LCAs of each experimental scenario. This is indeed a first step, that should lead in a longer perspective to model the link between concrete formulation and its performances, that should then serve to calculate the reference flow from the functional unit.

## 5. Cross-analysis: first recommendations

The multi-criteria approach developed in the framework of the project finally led to quite interesting results.

From Fig. 4c, accelerated migration testing has shown that the lowest Cl ions diffusivity is obtained for the studied concrete without fiber and that its increase is slight up to a fiber content of 40 kg.m⁻³. At a fiber content of 60 kg.m⁻³, its increase is more significant.

Simulations of chloride ingress under tidal conditions presented here show that, for the scenario SFRC F40, $C_{crit}$ at 5 cm depth is reached at $t_{crit}$ of 13 years of exposure for the less favorable case considered here ($\tau = 50 \%$) and at $t_{crit}$ of 38.4 years of exposure for the most favorable case ($\tau = 0 \%$). In reality, corrosion rate $\tau$ slowly increases with time and $t_{crit}$ would be somewhere between these two extremes (knowing that $t_{crit}$ linearly increases with $\tau$). This has to be compared with the RC F0 scenario which leads to $t_{crit}$ of 33 years. Excepted at very high degree of fiber corrosion extending deep inside the structure, values of $t_{crit}$ are quite respectable and whatever the scenario considered (RC F0 and SFRC F40), simulations tend to show that the service life of the structure would be suitable.

Furthermore, LCA performed here shows that SFRC F40 and SFRC F60 scenarios leads to significantly better results than RC F0 scenario in terms of environmental indicators.

Fig. 11 shows service life based on simulations (tidal conditions) taking into account originally measured values of $D_{0,eff}$ (see Fig. 4c) and carbon footprint as a function of the fiber content. For instance, for F40 formulation, RCMTs led to a value of $\tau$ of about 10 %. As explained in Reuge et al. [20], it is noteworthy that the potential difference applied during these tests (29 V) inevitably resulted in a decrease in the electrochemical potential of the steel and therefore in a strong acceleration of the corrosion process [48]. Moreover, the acceleration of the corrosion process is likely stronger than the acceleration of the Cl diffusion process. Thus, we consider that the simulation of Cl ingress with a value of $\tau$ of 10 % probably leads to an underestimation of $t_{crit}$.

Based on this Figure, increasing fiber content further provokes a loss of the durability functionality, but also a decrease of GWP indicator. However, the GWP is not improved above 40 kg.m⁻³ fiber content. Considering all the above elements, the SFRC F40 scenario thus seems to be the most promising for the case of a slab with the lowest carbon footprint (70 kg CO₂ eq/FU) and a service life of almost 25 years (which is in the error margin from 13 years to 38 years provided by the extreme cases studied in Section 3.4).

However, the transposition to the case of a float is not straightforward because mechanical requirements are different. Floats of FOWTs are subjected to bending moments, shear forces and sometimes torsion. Wave and swell have important effects: as waves pass, the distribution of buoyant force changes depending on the shape of the wave, which can cause hogging and sagging moments and deformation of the float. Moreover, the tower and nacelle apply a concentrated vertical load at the top of the float creating vertical bending moments in the structure and significant dynamic forces when the wind blows (due to oscillations) [49,50]. In practice, magnitudes of these various forces depend on the size of the structure. Thus, a new mechanical dimensioning and a new LCA will have to be envisaged for the case of a float. A hybrid structure (SFRC with a few rebars) might be a viable option.

## 6. Conclusions

The solution of a slag-based SFRC that would be used for floats of FOWT was studied considering the simplified case of a slab.

Mechanical dimensioning of the slab has been performed as a function of the fiber content and chloride ingress in the concrete subjected to tidal conditions has been simulated also as a function of the fiber content. These cases were compared with the solution of
<figure><img src="image_14.png" /><figcaption>Fig. 10. Relative impact of environmental indicators for the studied scenarios.</figcaption></figure>

<figure><img src="image_15.png" /><figcaption>Fig. 11. Service life (simulations, under tidal conditions) and carbon footprint as a function of fiber content for the studied case of a slab.</figcaption></figure>

RC slab. This case study leads to the following points:

- this mechanical dimensioning clearly shows that all SFRC F20, F40 and F60 scenarios reduce the amount of steel compared to the reference scenario RC F0,

- and that SFRC F40 and F60 also make it possible to use less concrete in significant proportions,

- the service life is sensitive to the corrosion degree of the fibers which was considered uniform in the concrete depth and constant with time. The case of a corrosion degree increasing with time and decreasing with depth in the material is not included here but will have to be studied in the future,

- the service life of a structure made of F40 formulation (fiber content: 40 kg.m⁻³) and subjected to cyclic sea water exposition can reasonably be estimated of around 30 years.

Based on these scenarios, a LCA has been performed by using the Functional Unit fulfilling the mechanical requirements. It leads to the following results for a French context:

- Global Warming Potential indicator is significantly impacted by the production step: production of SFRCs is significantly more impactful than production of RC due to the manufacturing process of fibers which is more energy consuming than the manufacturing process of rebars,

- Nevertheless, for the RC F0 scenario, the Global Warming Potential indicator of the installation step is really significant due to the
steel rebars implementation, which is not the case for the SFRC F20, F40 and F60 scenarios,

- Regarding the other environmental impact categories, the most favorable results are obtained with scenario SFRC F40 or F60 depending on the indicator.

- This LCA, based on experimental scenarios, was a first step that should lead in a longer perspective to model the link between concrete formulation and its performances, that should then serve to calculate the reference flow from functional unit based on concrete performances.

Finally, considering all the above elements, a cross-analysis (multi-criteria approach) has been conducted with the following conclusions:

- SFRC F40 seems to be the most promising for the case of a slab,

- the transposition from the case of a slab to the case of a float is not straightforward because mechanical requirements are different. Regarding the floats, a hybrid structure (SFRC with a few rebars) might be envisaged.

## CRediT authorship contribution statement

**Olivier Cornet:** Software, Methodology, Investigation, Formal analysis. **Anne Ventura:** Writing - review & editing, Methodology, Formal analysis. **REUGE Nicolas:** Writing - original draft, Visualization, Validation, Software, Methodology, Investigation, Formal analysis, Conceptualization. **Stéphanie Bonnet:** Writing - original draft, Validation, Supervision, Project administration, Methodology, Funding acquisition, Formal analysis, Conceptualization. **Laurence Guihéneuf:** Writing - original draft, Software, Methodology, Investigation, Formal analysis.

## Declaration of Competing Interest

The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.

## Acknowledgements

This study has been carried out within the framework of the WEAMEC program through the B2FE project. WEAMEC and its local partners, Région Pays de la Loire (France), FEDER, CARENE and Nantes Métropole, are gratefully acknowledged for their support. The mass transport simulation tool used in this study have been developed in the frame of the DEMCOM project (grant ANR-20-CE22-0008-03) funded by the French National Research Agency.

## Data availability

Data will be made available on request.

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