Structural and Geotechnical Assessment of Onshore Wind Turbine Foundation for Service Life Extension: A Case Study

Abstract

This study presents a structural and geotechnical assessment of an onshore wind turbine foundation that has been in service for approximately 15 years. It aimed to evaluate its suitability for service life extension under the current operational conditions, within the broader context of decision-making in aging wind farms. The investigation integrated original design documentation, detailed field inspections, in situ and laboratory geotechnical testing, and advanced 3D numerical modeling incorporating soil–structure interaction effects. Verification procedures followed international standards and current guidelines for the design and reassessment of wind turbine foundations. Critical structural and geotechnical aspects, including internal forces and reinforcement demand, stiffness, bearing resistance, settlement, and global stability, are examined to verify performance under the current operational loading conditions. The results provide a sound technical basis for strategic decision-making regarding service life extension or decommissioning of wind turbines in established wind farms, and constitute an essential baseline for any future structural upgrading associated with repowering strategies.

1. Introduction

The global onshore wind energy sector is currently facing a critical transition as a significant portion of early wind farms approaches the end of their original 20-year design life [1,2]. Asset owners must decide whether to decommission the site, extend the service life of the existing infrastructure, or repower the farm with larger turbines [3]. Service life extension has emerged as a highly strategic and cost-effective approach, allowing continued energy generation without the immediate environmental and financial costs of dismantling or constructing new civil works [4]. However, extending the operational period requires a rigorous structural and geotechnical reassessment of the existing foundations to ensure they can safely withstand the accumulated and ongoing operational loads beyond their initial design assumptions [5].

These reassessments must adhere to modern, stringent international guidelines, including IEC 61400-6 [6], Eurocode 2 [7], Eurocode 7 [8], and DNV-ST-0126 [9]. Despite the industry’s need, integrated evaluations of foundations after more than a decade of operation remain scarce in the literature. Few studies combine field observations, in situ geotechnical data, back-analysis, and advanced 3D soil–structure interaction (SSI) modeling to normatively validate the continued use of aging foundations under their current operational loading conditions.

To address this gap, this study conducts a comprehensive structural and geotechnical assessment of existing shallow and piled foundations at a real wind farm in southern Brazil. Both foundation typologies feature a circular geometry—the most widely used solution for onshore wind turbines globally—and represent the most representative support systems in the industry [10]. The evaluation relies on technical documentation, visual inspections, geotechnical testing, and 3D numerical modeling using SAP2000 software (v.26, Computers & Structures, USA). Crucially, the loads considered in this analysis are exclusively those of the existing wind turbine, aiming to evaluate the foundation’s adequacy for service life extension under current operational conditions.

This study presents a replicable methodology for assessing the structural and geotechnical capacity of existing wind turbine foundations, verifying if they can continue operating safely after 20 years without the need for structural reinforcement. Furthermore, while the primary focus is on service life extension, this rigorous baseline assessment constitutes an essential first step for any future repowering strategies. The overall reassessment procedure, detailing the integration of field investigations, numerical modeling, and code-based verifications, is summarized in a flowchart presented in the Section 3.

2. Case Study

This study focuses on a wind farm located in Bom Jardim da Serra, Santa Catarina State (southern Brazil), at an elevation of approximately 1400 m (Figure 1). The project began operations in 2011 and comprises sixty-two 1.5 MW wind turbines, totaling 93 MW of installed capacity. After roughly 15 years, the farm is undergoing an assessment for continued operation, service life extension, and potential future repowering.

3. Materials and Methods

Two Vensys/Goldwind V77 turbines, IEC IIA class, each equipped with a 77 m rotor and a 100 m high tubular tower, operating between 3.0 m/s (cut-in) and 22 m/s (cut-out), were selected. The load data were obtained directly from the manufacturer. The wind farm has two foundation types: (i) shallow foundations with circular footings and perimeter tie rods, and (ii) deep foundations with excavated piles and a cap block. Each turbine was selected to represent a specific foundation solution, providing a basis for its joint characterization in service. The database includes original structural and geotechnical designs, formwork and reinforcement drawings, rotary drilling reports, photographic records, operation and maintenance logs, as well as a visual inspection conducted in September 2025 and compressive strength tests on concrete cores extracted in situ. These data support the structural and geotechnical analyses presented in subsequent sections, following procedures defined by international guidelines for the design and reassessment of wind turbine foundations.

3.1. Foundation Geometry

The foundations consist of shallow foundations with anchors and deep foundations with excavated piles, both widely used in onshore wind turbines [11]. These solutions employ the same reinforced concrete structural block (15.0 m in diameter), with a thickness ranging from 1.30 m at the edge to 2.00 m at the center, and concrete of $f_{ck}$ = 30/37 MPa. In the shallow foundation (Figure 2), the circular footing is equipped with 16 tie rods positioned at a radius of 7.0 m, designed to resist tensile forces resulting from overturning moments.

For the piled foundation (Figure 3), the block is rigidly connected to 36 excavated piles, each 12.00 m deep and 300 mm in diameter, amounting to 298 m³ of concrete and 30 t of steel in the block. The tower-foundation connection utilizes an anchor ring embedded at the top of the block, a typical solution in wind turbine structural systems and functionally equivalent to anchor bolt-based systems [4].

The foundation types share the same block geometry and receive identical overall turbine loads. Thus, structural modeling produces similar results, varying only in the support system (shallow or piled).

3.2. Geotechnical Characterization

The site’s geotechnical characterization, based on sixty-two standard penetration tests (SPT) and rotary drillings, revealed two distinct subsurface conditions that justify the coexistence of shallow and piled foundations. In the shallow foundation area, the subsoil features rocky outcrops or shallow, highly rigid basalt overlain by a thin residual soil layer. In the piled foundation area, the surface soil is followed by sandy silt and fractured rock, which is supported by a more competent rock at depth. These transitional geotechnical profiles between residual soil, fractured rock, and sound rock are widely described in classic geotechnical literature [12]. Consequently, shallow footings were adopted for the first condition, whereas 36 excavated piles, each 12.00 m deep and 300 mm in diameter, were utilized in the second condition to transfer loads to more competent strata.

3.3. Actions and Load Combinations

During their useful life, onshore wind turbines are subjected to cyclic overturning actions from wind loading, with variable direction over 360°. Given its random nature in time and space, wind is a highly variable action, suitably represented by statistical models [13]. Loads transmitted to the foundation (axial force, horizontal force, torsional moment, and bending moments) were provided directly by the manufacturer of the existing 1.5 MW Vensys/Goldwind V77 turbine operating at the studied wind farm. These loads represent the operational conditions for which the foundation is currently being assessed for service life extension. Accurately defining loads and foundation response is central to wind turbine foundation design, as articulated in leading references [14,15]. The load combinations adopted follow the requirements of IEC 61400-6 [6], which defines the specific design load cases for structural assessment. The characteristic values used, and the partial factors applied in each load type are shown in

Table 1. Factored design load cases for the existing 1.5 MW Vensys/Goldwind turbine.

Design Load Cases	Load Type	FS	Fz (kN)	Fxy (kN)	Mz (kN·m)	Mxy (kN·m)
110	QPC	1.0	2972	302	949	24,175
210	ACC	1.1	3201	773	1355	61,853
310	Rare	1.0	2918	629	681	59,010
311	Extreme	0.90–1.35	2626	849	919	82,620
312	Extreme	1.35	3939	849	919	82,620
410	Fat. Mean	1.0	2950	219	178	19,472
411	Fat. Min.	1.0	-	105	-	9001
412	Fat. Max.	1.0	-	333	-	29,943

FS: Factor of safety; Fz: axial force; Fxy: horizontal force; Mz: torsional moment; Mxy: bending moments; QPC: quasi-permanent combination; ACC: accidental action combination; Fat: fatigue.

.

3.4. Numerical Modeling

The wind turbine foundations were numerically modeled using SAP2000 software (v. 26, Computers & Structures, Berkeley, CA, USA). Shell elements with variable thickness were employed to capture the stiffness distribution within the foundation block. The choice of shell elements is justified by the foundation’s geometry, where the thickness is significantly smaller than its other main dimensions. This approach allows for an efficient representation of bending and membrane stresses, as widely discussed in finite element literature [16]. Shell structures can be considered a generalization of plates, incorporating axial forces (membrane effect) in the mid-surface plane, in addition to bending behaviors. The Reissner-Mindlin flat shell element formulation is utilized to model the behavior of plates and shells, assuming that the cross-section remains planar after deformation and that the stress and strain diagrams along the cross-section are linear.

Soil–structure interaction was simulated using the Winkler model, which represents the soil as a series of independent springs acting vertically at the foundation base. This approach, although simplified, is widely recognized for its applicability in foundation analyses, especially when soil stiffness is obtained directly from field tests, such as plate load tests [17]. For shallow foundations, the importance of calibrating the vertical reaction coefficient based on on-site plate load tests is fundamental. In the context of piled foundations, the Winkler model is also applied to simulate soil behavior along the pile shaft (lateral springs) and at the pile toe (toe spring), allowing for an effective representation of soil-pile interaction.

Although the Winkler model does not inherently capture soil continuity or shear transfer between springs, the extreme rigidity of the wind turbine foundation block forces a planar deformation of the support system. Consequently, the springs are compressed jointly, mitigating the theoretical limitations of independent springs. For the global assessment of wind turbine foundations, where the focus lies on rotational response and global settlements under cyclic loading, this model offers an adequate balance between accuracy and computational efficiency.

The interface between the block and the upper structure was represented by link elements. The steel tower was modeled as a rigid body using constraint body features applied at the interface point. Mesh discretization was progressively refined until further refinement produced no significant change in global stresses or displacements, validating adequate convergence. The discretization level directly affects numerical accuracy; finer meshes improve solution fidelity while increasing computational cost [18]. The discretized model, a 3D extrusion of the shallow foundation, and of the piled foundation are presented in Figure 4.

3.5. Regulatory Criteria and Verifications

The structural and geotechnical verification of foundations followed the IEC 61400-6 [6], establishing the structural supports’ design for onshore wind turbines. The Eurocode 2 [7] criteria were applied for reinforced concrete, while soil bearing capacity and geotechnical safety checks were conducted according to the Eurocode 7 [8]. Complementary guidelines for provisions for shallow and deep foundations were drawn from DNV-ST-0126 [9], widely used in wind farm projects. These normative requirements were consistently applied to both foundation types since their structural geometry is similar and they are subjected to equivalent turbine loads. This approach aligns with established recommendations in foundation engineering literature, such as the fundamentals of soil behavior [4], and the structural principles applied to towers and their support systems [10].

3.5.1. Geotechnical Analyses

(a)

Soil bearing capacity: The ultimate bearing capacity was determined by a plate load test at the foundation level, conducted under slow loading with loading and unloading cycles, as per standard geotechnical protocols [19,20]. The test directly provided the ultimate capacity and the reaction modulus required in numerical modeling (Equation (1)):

$$qu={\displaystyle \frac{Pult}{A}}$$

(1)

where $qu$ denotes the soil’s ultimate capacity, $Pult$ is the final load applied to the plate, and $A$ is the plate area. Soil–structure interaction was modeled using Winkler’s concept, where the soil serves as a series of springs of stiffness $Ks$, idealized as linearly elastic for small deformations [21]. The test also allowed determination of the soil reaction modulus (Equation (2)), relating applied stress to observed displacement:

$$Ks={\displaystyle \frac{q}{\delta }}$$

(2)

where $Ks$ is the soil reaction modulus, $q$ is the unit area load, and δ is the corresponding foundation displacement under stress.

(b)

Stress distribution: For shallow, gravity-type foundations, contact pressure beneath the foundation is assumed to be linearly distributed, following the classical principles of material resistance. The general expression for contact stress is presented in Equation (3):

$$\sigma max={\displaystyle \frac{Nk+G(soil+footing)}{Afooting}}+{\displaystyle \frac{M}{W}}$$

(3)

where $\sigma max$ is the maximum stress at the soil-foundation interface, it is derived from the combined normal load $Nk$, the cover soil weight added to the footing weight, $G$, and the footing area, $Afooting$. Moreover, the base moment $M$ and the section modulus $W$ defined by the foundation geometry, govern the stress distribution used in stability assessment.

(c)

Overturning: Overturning evaluation (Equation (4)) compares the stabilizing moments from vertical loads with destabilizing moments from horizontal and bending actions. The safety factor for overturning should exceed 1.5 in the serviceability limit state (SLS) and 1.0 in the ultimate limit state (ULS) [11].

$${SF}_t={\displaystyle \frac{Mr}{Ms}} > 1.5 \left(\mathrm{SLS}\right), 1.0 \left(\mathrm{ULS}\right)$$

(4)

where, $Mr$ is the sum of vertical forces times the foundation radius, and $Ms$ is the moment load applied by the turbine manufacturer. Their ratio defines the overturning safety factor ${SF}_t$.

(d)

Sliding: To verify sliding, horizontal actions should not exceed the friction resistance at the soil-foundation interface. Following Milititsky [11], a minimum safety factor of 1.5 is used. For drained conditions, this verification can be expressed by Equation (5):

$${SF}_d={\displaystyle \frac{F_r}{F_x}}>1.5$$

(5)

where the friction coefficient μ is tan (2/3 Φ), the soil-base interaction. $Dt$ is the combined weight of base and superimposed soil, and $F_z$ is the applied axial force (Equation (6)). The reduction coefficient $\delta$ is used according to regulatory criteria, and these parameters yield the sliding safety factor ${SF}_d$, evaluating the foundation’s horizontal stability.

$$F_r=\mu \times (\delta \times Dt+F_z)$$

(6)

(e)

Decompression with “no gapping”: For decompression assessment, eccentricity is evaluated under SLS and ULS conditions. Eccentricity must be <$K_1$ in SLS, ensuring a fully compressed area, and <$K_2$ in ULS, securing at least 50% compression [22]. These limits are described in Equations (7) and (8).

$$K_1=0.25\times R$$

(7)

$$K_2=0.25\times R\times \pi$$

(8)

The inertia core respective $K_1$ and $K_2$ boundaries are illustrated in Figure 5.

(f)

Rotational stiffness: To assure proper tower-foundation-soil system performance, wind turbine manufacturers require a minimum rotational stiffness, controlling the system’s natural frequency. This performance is checked at SLS with the characteristic action combination. The IEC 61400-6 [6] specifies dynamic and static soil behavior stiffness limits, as outlined in Equation (9):

$$K_{R,dyn}={\displaystyle \frac{8\times G_0\times R^3}{3\times (1-\upsilon )}}$$

(9)

where dynamic rotational stiffness ($K_{R,dyn}$) corresponds to the foundation’s ability to resist overturning moments. $G_0$ (Equation (10)) is the initial distortion modulus of the supporting soil. $R$ is the foundation radius, $\upsilon$ is the Poisson’s ratio of the soil material, and $E$ is the modulus of elasticity that determines ground deformation under rotational loads.

$$G_0={\displaystyle \frac{E}{2\times (1+\upsilon )}}$$

(10)

Rotational stiffness can be estimated more accurately by finite element models that include explicit soil representation or distributed springs under the foundation, incorporating the footing flexibility and deriving stiffness from the ratio of applied moment $M$ to the resulting rotation $\theta$ (Equation (11)).

$$K_{\theta }={\displaystyle \frac{M}{\theta }}$$

(11)

(g)

Load distribution on piles and geotechnical sizing: For piled foundations, the pile group is typically idealized as a rigid set of $n$ piles. distributed at a constant radius from the block’s center. When subjected to vertical load $N$ and overturning moment $M_R$, and assuming plane sections remain plane, the axial load on the i-th pile at position $z_i$ is given by Equation (12):

$$N_n={\displaystyle \frac{N}{n}}\mp {\displaystyle \frac{M_R\times z_i}{{\sum }_j\times x^2}}$$

(12)

where $N_n$ is the axial load on pile i, $N$ is the total vertical load applied, $n$ is the number of piles, $M_R$ is the applied moment, $z_i$ is the pile’s moment arm along the moment axis, and ${\sum }_j\ast x^2$ is the pile group’s second moment of area, quantifying its moment resistance. The axial pile capacity comprising both tip and lateral resistance along the embedded pile length is $R_{c,k}$, Then, the safety factor must be applied. Detailed expressions for tip and shaft resistances are presented in Equation (13).

$$R_{c,k}{=\left[\right(R}_{b,k}=q_{b,k}\times A_b)+{(R}_{s,k}=\Sigma (q_{s,k}\times A_i)]$$

(13)

where $R_{b,k}$ is the characteristic pile base resistance, $q_{b,k}$ the unit resistance mobilized at the pile’s base, and $A_b$ is the base area. Similarly, $R_{s,k}$ is shaft resistance, $\Sigma$ is the sum of the overall soil layer contribution, $q_{s,k}$ is the layer unit resistance, and $A_i$ is the pile area per layer. The structural resistance of the pile is detailed in Equation (14), where $R_{d }$ is the design pile resistance and $f_{cd}$ is the design concrete compressive strength.

$$R_d=A_b\times f_{cd}$$

(14)

3.5.2. Structural Analyses

(a)

Flexural and shear reinforcement: Flexural and shear reinforcement were sized per Eurocode 2 [7], based on classical reinforced concrete relations (Equation (15)).

$$A_s={\displaystyle \frac{M_{Ed}}{z\times f_{yd}}}$$

(15)

where, $A_s$ is the required tensile reinforcement area, $M_{Ed}$ is the design bending moment, $z$ is the internal lever arm, $f_{yd}$ is the steel yield capacity. The formulation accounts for compressed concrete, tensile reinforcement, and shear resistance contributions (Equation (16)). Thus, both minimum and required steel areas for ULS and SLS are determined by the requests obtained in the numerical modeling.

$$V_{Rd,c}=\left[C_{Rd,c}\times k\times {\left(100\times \rho l\times f_{ck}\right)}^{{\displaystyle \frac{1}{3}}}+k_1\times {\sigma }_{cp}\right]\times b_w\times d$$

(16)

The stress check considers the concrete strength $V_{Rd,c}$, the coefficient $C_{Rd,c}$, the scaling factor $k$, the longitudinal reinforcement ratio $\rho l$, and the compressive strength $f_{ck}$. Parameter $k_1$ incorporates average compressive stress ${\sigma }_{cp}$ from the axial load, $b_w$ is the effective width, and $d$ is the effective height. According to Eurocode 2 [7], the minimum longitudinal reinforcement for flexural elements is determined by Equation (17):

$$A_{s,min}=0.26\times {\displaystyle \frac{f_{ctm}}{f_{yk}}}\times b_w\times d$$

(17)

where $f_{ctm}$ is the mean concrete’s tensile strength and $f_{yk}$ is the characteristic steel strength.

(b)

Crack control: Crack control was checked using standard reinforced concrete relationships to estimate the characteristic crack opening as a function of reinforcement stress and the section’s deformability (Equation (18)). Verification was performed for service combinations, ensuring that the calculated opening remained within the limit of w ≤ 0.3 mm, in line with SLS performance requirements.

$$w_k=S_{r,max}\times \left({ϵ}_{sm}-{ϵ}_{cm}\right)\le 0.30 \mathrm{m}\mathrm{m}$$

(18)

where, $w_k$ is the service-state characteristic crack width at SLS, $S_{r,max}$ is the maximum crack spacing from the steel-concrete bond model, ${ϵ}_{sm}$ is the mean post-cracking tensile reinforcement’s deformation, ${ϵ}_{cm}$ is mean concrete strain between cracks, affected by shrinkage, creep, and bond-based stress transfer.

(c)

Steel fatigue control: Reinforcement was checked for fatigue per Eurocode 2 [7], with stress variation from turbine loading cycles kept <70 MPa permissible limit (Equation (19)), achieving adequate fatigue performance over the structure’s service life.

$${∆}_{\sigma s}\le {∆}_{\sigma Rsk}=70 \mathrm{M}\mathrm{P}\mathrm{a}$$

(19)

Here, ${∆}_{\sigma s}$ is the steel stress variation from cyclic actions, ${∆}_{\sigma Rsk}$ is characteristic fatigue strength. For conventional reinforcement, ${∆}_{\sigma Rsk}$ = 70 MPa is defined for maximum stress variation in any cycle.

(d)

Fatigue control in concrete: Fatigue control followed Eurocode 2 [7], evaluating the relationship between maximum and minimum compressive stresses under repeated actions. Concrete fatigue resistance was calculated for $f_{ck}$ ≤ 50 MPa (Equation (20)), ensuring compressive stress cycles remain within the standard fatigue resistance limits.

$${\displaystyle \frac{{\sigma }_{c,max}}{f_{cd,fat}}}\le 0.5+0.45\times {\displaystyle \frac{{\sigma }_{c,min}}{f_{cd,fat}}}\le 0.9 for f_{ck}\le 50 \mathrm{M}\mathrm{P}\mathrm{a}$$

(20)

Here, ${\sigma }_{c,max}$ is the maximum compressive stress in the concrete under the fatigue load cycle, while ${\sigma }_{c,min}$ is the minimum. $f_{cd,fat}$ is the design compressive strength for fatigue, obtained by applying a partial safety factor and fatigue reduction to $f_{ck}$. The upper limit of 0.9 applies to concretes with $f_{ck}$ ≤ 50 MPa under current standards.

These verifications support a comprehensive assessment of the foundation’s capacity subjected to acting forces, confirming its functionality for the remaining life cycle, and informing future decisions to extend the service life or repower the turbines.

To facilitate the application of the proposed approach to other onshore wind farms, a comprehensive overview of the methodology is presented. Figure 6 summarizes the overall reassessment procedure adopted in this study. This workflow integrates data collection, detailed field and laboratory investigations, geotechnical and structural characterization, advanced 3D soil–structure interaction modeling, and code-based verifications. The process culminates in a structured decision framework designed to support informed choices regarding service life extension, repowering, or decommissioning of existing wind turbine foundations.

4. Results

4.1. Soil Bearing Capacity and Stress Distribution in the Soil

The soil bearing capacity was determined via plate load testing, which yielded an ultimate stress of 800 kN/m². Applying a safety factor of 2, the allowable stress was set at 400 kN/m², with a measured settlement of 2.55 mm. For ULS, stresses calculated according to Ortiz [23] and the classical material resistance formulation produced maximum and minimum values of 382.04 kN/m² and −129.14 kN/m², respectively. For SLS, maximum and minimum stresses were 169.00 kN/m² and 18.08 kN/m², respectively. Numerical modeling resulted in a maximum soil stress of 350 kN/m² and a minimum of −105 kN/m² for ULS, as shown in the stress map (Figure 7). For SLS, maximum and minimum stresses were 220 kN/m² and 77 kN/m², respectively (Figure 7). All stresses obtained for the SLS remained compressive, confirming the absence of gapping in the quasi-permanent combination. All calculated stresses were below the allowable 400 kN/m², confirming the adequacy of the foundation for the considered loads.

4.2. Decompression, Eccentricity, Overturning, and Sliding

The decompression limits were verified for SLS and ULS, based on criteria defined for wind turbine foundations in accordance with IEC 61400-6 [6]. For SLS, K₁ = 1.49 m was obtained, lower than the permissible limit of 1.87 m. Thus, the base was fully compressed with no gapping. In ULS, K₂ was 3.77 m, also below the limit of 5.89 m, establishing that at least half of the contact area remained compressed under design loading. These values fully satisfy the eccentricity and compression criteria for the analyzed foundation. Furthermore, the global stability of the shallow foundation was evaluated. The calculated overturning safety factors were 4.95 for the Serviceability Limit State (SLS) and 2.05 for the Ultimate Limit State (ULS). Similarly, the sliding safety factors were determined to be 22.98 (SLS) and 8.38 (ULS), demonstrating adequate stability under the considered load combinations.

4.3. Rotational Stiffness of the Foundation

The rotational stiffness of the shallow foundation, obtained from SAP2000 modeling, was 6.08 × 10¹⁰ N·m/rad, exceeding the turbine manufacturer’s minimum requirement of 4.5 × 10¹⁰ N·m/rad. Additionally, the calculated differential settlement for the shallow foundation was (10.45 mm), which remains safely below the allowable limit of (14.11 mm). The analysis used a soil elasticity modulus of 141 MPa and a shear modulus of 54 MPa, with a stiffness reduction factor of 0.7, yielding G/G₀ = 38 MPa. For the piled foundation, rotational stiffness was calculated from the ratio of applied moment to angular rotation obtained by numerical modeling (Figure 8). The rotation corresponded to the ratio of maximum differential deflection (8.23 mm) to the foundation width (15.0 m), expressed in radians. Dividing the critical moment of 84,688 kN·m by the corresponding rotation produced a rotational stiffness of 1.54 × 10¹¹ N·m/rad, greater than the manufacturer’s requirement.

The results confirm that the foundations present adequate stiffness for the turbine’s operation under representative loading conditions. Moreover, the deflection differences remained within the 3 mm/m criterion commonly referenced when no limit is specified [24]. The force distribution in the piles was evaluated for SLS and ULS combinations, following IEC 61400-6 requirements [6]. For DLC 110 (SLS), the most stressed pile had a compressive force of 644 kN, while the least had 264 kN, both remaining entirely compressed, meeting the no tension criterion (“no pile tension”) for the quasi-permanent combination. For DLC 312 (ULS), the maximum compressive force was 1296 kN, with a tensile force of 48 kN in the most stressed pile. The geotechnical design, based on the characterized stratigraphic profile, resulted in an ultimate resistance capacity of 1661.67 kN per pile, used as a reference value for verifying the stresses obtained. Furthermore, the structural design resistance of the pile was determined to be 1413.17 kN, confirming its capacity to withstand the maximum applied axial forces.

4.4. Summary of Internal Stresses in the Foundation

The internal stresses were determined along ten representative circular sections throughout the foundation block, as illustrated in Figure 9.

The radial bending moment, circumferential bending moment, radial shear, and circumferential shear maps were obtained via numerical modeling (Figure 10).

The figures corresponding to the shallow foundation demonstrate the bending and shear gradients resulting from the load combinations. For the piled foundation, the stress maps display very similar behavior and magnitudes, as both solutions share the same block geometry and are subjected to identical global loads. The resulting distributions yield precise identification of high-stress regions, serving as the graphical basis for analyzing the foundation’s performance.

The complete set of numerical values obtained for internal stresses in the shallow foundation under service load (DLC 110—SLS) and ultimate load (DLC 312—ULS) combinations, summarizing the bending moments and shear stresses used for subsequent checks are presented in

Table 2. Bending moments and shear forces by section (shallow foundation).

Section	M11	M22	V13	V23	M11	M22	V13	V23
	Serviceability Limit State				Ultimate Limit State
	kN·m	kN·m	kN	kN	kN·m	kN·m	kN	kN
1	0	0	0	−59	28	−31	−9	−210
2	−166	−151	−128	−104	−385	−424	−365	−347
3	−440	−298	−158	−119	−1288	−958	−624	−414
4	−1090	−414	−230	−63	−3715	−1413	−1074	−243
5	−682	−146	−129	−22	−2472	−529	−580	−96
6	727	171	272	35	2395	563	753	107
7	1084	416	507	90	3693	1409	1436	278
8	224	212	237	162	966	806	724	423
9	33	57	63	97	129	295	157	339
10	11	71	26	63	11	147	59	212

M11: Radial bending moment; M22: circumferential bending moment; V13: radial shear; V23: circumferential shear.

. The results adhere to the standard sign convention: negative values indicate tension and positive values indicate compression in structural elements, facilitating direct interpretation and comparison across the different limit states evaluated.

The piled foundation, comprising a rigid block connected to the pile group, was evaluated under combinations DLC 110 (SLS) and DLC 312 (ULS), leading to corresponding internal forces obtained via numerical modeling. The complete values are shown in

Table 3. Bending moments and shear forces per section (piled foundation).

Section	M11	M22	V13	V23	M11	M22	V13	V23
	Serviceability Limit State				Ultimate Limit State
	kN·m	kN·m	kN	kN	kN·m	kN·m	kN	kN
1	0	−8	−9	−66	0	−63	−33	−218
2	−49	−62	−75	−102	−209	−297	−300	−348
3	−251	−173	−136	−120	−1018	−756	−590	−420
4	−819	−315	−221	−64	−3229	−1264	−1070	−246
5	−514	−117	−123	−21	−2281	−539	−576	−91
6	868	215	282	37	2586	571	787	123
7	1346	507	538	95	4046	1531	1492	294
8	369	323	277	134	1211	952	631	438
9	74	136	127	107	222	393	266	357
10	0	36	15	66	0	95	40	219

M11: Radial bending moment; M22: circumferential bending moment; V13: radial shear; V23: circumferential shear.

, enabling direct comparison with the shallow foundation, and constitute the basis for structural assessment.

The existing steel area in the shallow foundation and the recalculated values for each section analyzed are compared in Figure 11. Although a few sections exhibit reinforcement areas slightly lower than the calculated values, the total steel distributed throughout the foundation exceeds the required amount from the analyses, thereby maintaining compliance with the adopted structural criteria.

The comparison between the existing steel area in the piled foundation and the recalculated values from this study is presented in Figure 12. Minor variations arise from moderate differences in the bending moments for this configuration. However, the total existing reinforcement area exceeds the calculated requirement. Similar to the shallow foundation, the overall steel distribution in the piled foundation is consistent with the evaluated results.

4.5. Crack Evaluation

Serviceability analyses indicated maximum crack widths of 0.15 mm for the shallow foundation and 0.21 mm for the piled foundation, both within the adopted limit of w_k ≤ 0.30 mm. Cracks are concentrated at the underside of the block, along the edge opposite the applied load, where the highest positive bending moments occur in the quasi-permanent combinations. The block thickness and continuous ring reinforcement ensured adequate crack control in the systems.

4.6. Fatigue Evaluation

The calculated stress ranges remained within regulatory limits. The shallow foundation presented ${∆}_{\sigma s}$ = 63.75 MPa, below the limit ${∆}_{\sigma Rsk}$ = 70 MPa, while the piled foundation showed ${∆}_{\sigma s}$ = 56.21 MPa at the critical section, also within the permissible range. For concrete fatigue assessment, the maximum-to-minimum stress ratio was 0.21 for the configurations, below the prescribed limit of 0.50, confirming that cyclic loading remains within the allowable safety domain for concrete.

5. Discussion

This section critically summarizes the geotechnical and structural results obtained for two foundation solutions, comparing them with the regulatory limits established by IEC 61400-6 [6], Eurocode 2 [7], Eurocode 7 [8], and DNV-ST-0126 [9]. The analysis is presented comparatively for the shallow and piled foundations, assessing performance under SLS and ULS conditions and evaluating the compatibility of the structural responses with the requirements applicable to the existing 1.5 MW onshore turbines. The discussion focuses on resistance capacity, stiffness, stability, reinforcement demand, and fatigue assessment, with emphasis on the suitability of existing foundations for service life extension.

For the shallow foundation, the mobilized bearing capacity under ULS resulted in a contact stress of 382.04 kN/m², below the permissible value of 400 kN/m², satisfying geotechnical safety requirements. Eccentricities were within normative limits, with K₁ = 1.49 m (limit 1.87 m) and K₂ = 3.77 m (limit 5.89 m), confirming full compression under SLS and at least 50% compressed area under ULS. The overturning safety factors were 4.95 (SLS) and 2.05 (ULS), exceeding the minimum requirement of 2.0, while sliding safety factors were substantially higher (22.98 for SLS and 8.38 for ULS).

The calculated rotational stiffness (6.08 × 10¹⁰ N·m/rad) surpassed the manufacturer’s minimum requirement of 4.50 × 10¹⁰ N·m/rad. Differential settlement reached 10.45 mm, below the allowable limit of 14.11 mm. From a structural standpoint, the required longitudinal reinforcement (303.21 cm²) was fully satisfied by the provided area (330.53 cm²), and the installed transverse reinforcement (40.12 cm²) exceeded the minimum requirement (34.43 cm²). Serviceability and fatigue assessments were also satisfied: crack width (0.15 mm ≤ 0.30 mm), steel fatigue (63.75 MPa < 70 MPa), and concrete fatigue (0.21 < 0.50). Overall, the performance of the shallow foundation is consistent with experimental evidence indicating that such systems often exhibit ultimate capacities exceeding operational demand [25], and aligns with the mechanisms described by Das [12] and Milititsky [11].

For the piled foundation, composed of a rigid block identical to that of the shallow foundation and supported by 36 piles, each 12.00 m long and 300 mm in diameter, the mobilized geotechnical resistance was 1296 kN per pile, below the ultimate capacity considered in design (1661.67 kN per pile). No tensile stresses were observed under the quasi-permanent load combination. The global rotational stiffness, 1.54 × 10¹¹ N·m/rad, substantially exceeded the minimum requirement of 4.50 × 10¹⁰ N·m/rad, while the calculated differential settlement (8.23 mm) remained below the allowable limit of 14.11 mm.

From a structural perspective, the maximum axial force in the piles (1413.17 kN) was compatible with the design resistance. The existing longitudinal reinforcement (489.37 cm²) and transverse reinforcement (29.38 cm²) differ from the original design values (555.87 and 90.79 cm², respectively), yet satisfy the demands imposed by the evaluated load combinations. Serviceability and fatigue assessments were also met: crack width (0.21 mm ≤ 0.30 mm), steel fatigue (56.21 MPa < 70 MPa), and concrete fatigue index (0.21 < 0.50). Although the piles are not of the PHC type investigated by Charles et al. [24], the observed behavior is consistent with findings that well-confined deep foundations exhibit stable performance even under the accumulated cyclic loading associated with extended operational periods.

Comparatively, the results reflect the distinct resistance mechanisms in each solution: surface stress distribution in the shallow foundation and load transfer through shaft friction and tip resistance in the piled foundation. These mechanisms explain the greater rotational stiffness and reduced settlement observed in the piled foundation. Both behaviors are consistent with modern geotechnical theory [11,12].

These results corroborate recent literature on the service life extension of wind turbine foundations. Studies such as that by Jeong et al. [4] demonstrate that aged foundations, when not subjected to significant material degradation, maintain high structural safety margins over time. Furthermore, the adequate rotational stiffness and low differential settlements observed in our models align with the parameters highlighted by Toña et al. [3] as decisive for the continued use of these structures. Adding to this, recent experimental evidence presented by Lago et al. [25] proves that wind turbine foundation systems frequently possess ultimate capacities that far exceed standard operational demands. This convergence of findings—material integrity over time, adequate global stiffness, and high ultimate capacity—reinforces the premise that it is technically feasible and safe to maintain the continued operation of these structures beyond their original 20-year design life, without the need for strengthening interventions.

Overall, the findings demonstrate that both foundation systems maintain adequate geotechnical safety, stiffness, and structural performance for continued operation under the existing 1.5 MW turbine loads. This outcome aligns with the guidelines discussed by Toña et al. [3], which indicate that continued operation is viable when critical parameters are within regulatory limits and supports the feasibility of service life extension. The methodology presented herein offers a robust and replicable framework for engineers and asset managers to assess similar existing foundations, thereby contributing to the broader understanding of infrastructure longevity in wind energy. It is also in line with the multifactorial framework proposed by Gil García et al. [26], emphasizing technical optimization and efficient utilization of existing infrastructure. The compatibility with current civil infrastructure constraints, such as fixed anchorage systems, limited site access, and existing foundation geometry, further reinforces the technical and economic advantages of maintaining existing foundations for extended operational periods. Thus, the results demonstrate that the continued use of the existing foundation systems is technically feasible and structurally compatible with current operational demands, laying the groundwork for future repowering considerations.

6. Conclusions

Geotechnical and structural analyses demonstrate that both evaluated foundation solutions (shallow and piled, with 36 piles rigidly connected to the block) comply with IEC 61400-6, Eurocode 2, Eurocode 7, and DNV-ST-0126 for the continued operation of the existing 1.5 MW onshore turbines. In both cases, safety margins, rotational stiffness, and differential settlements remained within allowable limits, confirming the technical viability of the existing foundations for extended service life without the need for additional reinforcement.

However, this study has some limitations. The geotechnical characterization relied on only two soil profiles, limiting the assessment of spatial variability and potentially affecting stiffness and bearing capacity estimates. Fatigue assessment did not include a full S-N curve approach with Palmgren-Miner damage accumulation, as recommended for long-term assessments. Additionally, the structural model adopted a simplified 3D reinforcement idealization, which may influence local predictions of punching shear, cracking, and stiffness.

From a practical standpoint, the findings support the feasibility of extending the service life of existing foundations and serve as a reliable baseline for future repowering projects, thereby reducing decommissioning costs, environmental impacts, and the need for new civil works. For engineering practice, the study highlights the value of integrating inspection data, back-analysis, and soil–structure interaction modeling to guide maintenance or life-extension decisions. From a policy perspective, it demonstrates that legacy infrastructure can achieve adequate safety levels when reassessed under updated regulatory criteria.

Future research should expand the database of aged foundations assessed for life extension, incorporate probabilistic treatment of geotechnical and structural uncertainties, evaluate hybrid strengthening alternatives, and implement long-term structural health monitoring to validate performance over time. Furthermore, the framework established in this study provides a critical baseline for future investigations into the feasibility of repowering these sites with larger capacity turbines.