# Performance assessment of a novel connector for timber frame facades on concrete slabs

Tran-Van Han, Alice Le Berder, Tuan-Anh Nguyen, Clemence

Lepourry-Nicollet, Hugues Somja

► To cite this version:

Tran-Van Han, Alice Le Berder, Tuan-Anh Nguyen, Clemence Lepourry-Nicollet, Hugues Somja. Performance assessment of a novel connector for timber frame facades on concrete slabs. Journal of Building Engineering, 2025, 114, pp.114397. 10.1016/j.jobe.2025.114397. hal-05421932

HAL Id: hal-05421932

https://univ-rennes.hal.science/hal-05421932v1

Submitted on 7 Jan 2026

HAL is a multi-disciplinary open access
archive for the deposit and dissemination of sci-
entific research documents, whether they are pub-
lished or not. The documents may come from
teaching and research institutions in France or
abroad, or from public or private research centers.

L'archive ouverte pluridisciplinaire HAL, est
destinée au dépôt et à la diffusion de documents
scientifiques de niveau recherche, publiés ou non,
émanant des établissements d'enseignement et de
recherche français ou étrangers, des laboratoires
publics ou privés.

<figure><img src="image_1.png" /></figure>

Distributed under a Creative Commons CC BY 4.0 - Attribution - International License
<figure><img src="image_2.png" /></figure>

ELSEVIER

Contents lists available at [ScienceDirect](https://www.sciencedirect.com)

Journal of Building Engineering

journal homepage: [www.elsevier.com/locate/jobe](https://www.elsevier.com/locate/jobe)

<figure><img src="image_3.png" /></figure>

Full length article

<figure><img src="image_4.png" /></figure>

# Performance assessment of a novel connector for timber frame facades on concrete slabs

Tran-Van Han<sup>a,c, ID,*</sup>, Alice Le Berder<sup>a,b</sup>, Tuan-Anh Nguyen<sup>a ID</sup>, Clemence Lepourry-Nicollet<sup>b</sup>, Hugues Somja<sup>a</sup>

## ARTICLE INFO

**Keywords:**

- Timber Frame Facades
- Concrete-façades connections
- Perfobond connector
- Pull-out anchorage resistance
- Finite element analysis
- Concrete Capacity method

## ABSTRACT

This study examines the structural performance of a U-profile fastening system designed to connect timber frame façades (TFFs) to concrete slabs in hybrid timber-concrete construction. The system consists of U-shaped steel profiles with flange cutouts, partially embedded in concrete to facilitate early-stage façade installation. Fifteen full-scale specimens, varying in embedment depth, cutout geometry, and concrete compressive strength, were tested under simulated wind loads representing suction conditions. Concrete cone breakout was identified as the predominant failure mode. Experimental results indicated that embedment depth and concrete strength were the main factors governing anchorage resistance. In contrast, variations in cutout geometry and localized reinforcement had minimal influence. Finite element simulations captured the overall structural behavior and failure mechanisms with acceptable accuracy, showing a maximum deviation of 7% in peak load relative to experimental results. An analytical model, based on the Concrete Capacity method and existing design standards, was developed to estimate anchorage capacity and showed consistent agreement across the tested configurations. The findings contribute to the structural design of timber-concrete façade connections incorporating embedded U-profile connectors.

# 1. Introduction

## 1.1. Overview of façade and fastening technologies

Non-structural components (NSCs) play a vital role in modern construction by improving building functionality, safety, and sustainability. Defined as elements not part of the primary load-bearing system [1], NSCs include architectural components, MEP systems, and furniture/fixtures [2]. Among these, lightweight façade systems have gained attention for their favorable balance between structural efficiency, functional performance, and environmental impact [3].

Façade systems serve both aesthetic and functional roles and are commonly classified as infill panels, claddings, or mixed systems [4]. Infill panels (e.g., brick walls) remain prevalent in Europe, while lightweight timber- or steel-framed panels are widely

<footnote><sup>a</sup> INSA Rennes, LGCGM Structural Engineering Research Group, 20 Avenue des Buttes de Coesmes, CS 70839, Cedex 7 F-35708, Rennes, France</footnote>

<footnote><sup>b</sup> INGENOVA, Civil Engineering Office, 5 Rue Louis Jacques Daguerre, Saint-Jacques-de-la-Lande 35136, France</footnote>

<footnote><sup>c</sup> Faculty of Civil Engineering, Ton Duc Thang University, Ho Chi Minh City, Viet Nam</footnote>

<footnote>* Corresponding author at: INSA Rennes, LGCGM Structural Engineering Research Group, 20 Avenue des Buttes de Coesmes, CS 70839, Cedex 7 F-35708, Rennes, France.</footnote>

<footnote>E-mail addresses: Van-Han.Tran@insa-rennes.fr, tranvanhan@tdtu.edu.vn (T.-V. Han).</footnote>

<footnote>https://doi.org/10.1016/j.jobe.2025.114397</footnote>

<footnote>Received 15 July 2025; Received in revised form 10 October 2025; Accepted 13 October 2025</footnote>

<footnote>Available online 17 October 2025</footnote>

<footnote>2352-7102/© 2025 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY license (http://creativecommons.org/licenses/by/4.0/).</footnote>
<figure><img src="image_5.png" /><figcaption>**Fig. 1.** Concept of the novel connector system.</figcaption></figure>

used elsewhere. Claddings-from curtain walls to precast panels-are typically attached with discrete connectors; mixed systems combine both concepts to improve insulation and architectural flexibility.

Connection systems are decisive for transferring façade actions to the structural frame. Depending on the façade type, connectors can be continuous (e.g., mortar beds or metal rails in infill systems) or discrete (e.g., angle brackets [3,5–10], anchors [11], or bolts [12–17]). Continuous solutions offer uniform support but limited adjustability; discrete solutions enable modularity and simpler installation. In practice, many discrete systems rely on angle brackets or post-installed anchors, whose performance is sensitive to workmanship, tolerances, and installation sequence. These limitations can lead to misalignment and reduced load resistance under combined mechanical and environmental actions.

Policy drivers for low-carbon construction-such as RE2020 in France [18], the EPBD in Europe [19], and Japan's energy performance act [20]-further motivate the use of timber façade systems. Timber frame façades (TFFs) offer lower embodied carbon and seismic mass than conventional concrete envelopes [3], along with favorable thermal and acoustic performance. However, their integration into hybrid timber-concrete buildings remains limited, largely due to the lack of efficient, robust connectors that ensure reliable load transfer, dimensional tolerance, and long-term durability.

Perfobond-type steel-concrete connectors have been extensively studied in composite structures, particularly for bridge applications. Since their introduction by Leonhardt et al. [21], numerous studies have explored their shear and tensile behavior [22–26]. These connectors achieve high strength through dowel interlocking and concrete confinement. However, existing research has focused primarily on beam–slab interfaces, not on façade anchorage systems where load paths, geometrical constraints, and installation conditions differ significantly.

## 1.2. Novel connector for timber façade systems

To overcome the limitations of conventional angle brackets and post-installed anchors in TFF applications, a novel cast-in-place connector was developed. Inspired by perfobond concepts, the proposed system utilizes U-shaped steel profiles with flange cutouts that allow fresh concrete to flow through and interlock, forming mechanical studs for shear transfer. Positioned along the edge of the concrete slab, the connector serves as a robust interface between the façade panels and structural frame, as illustrated in Fig. 1.

The proposed system consists of embedded U-profiles, vertical support brackets, serrated plates, and anti-tip devices. The connectors are fabricated by bending steel plates, with precisely cut holes forming dowel openings and serrations, ensuring high dimensional accuracy and consistency. On site, the vertical brackets and anti-tip devices are attached to the cast-in U-profiles using bolts and screws through oblong holes. The serrated plates interlock with the pre-formed serrations on the U-profile, allowing the brackets to slide vertically and rotate for alignment before tightening. Once fastened, the teeth engage to lock the position, providing robustness against minor placement deviations during casting. This modular assembly sequence simplifies installation, eliminates on-site drilling or welding, and accommodates geometric adaptability, making the system well suited for hybrid timber-concrete applications where standardization and speed of construction are critical.

## 1.3. Research objectives

This study investigates the anchorage performance of the proposed connector under wind-induced loading. The main objectives are to assess its mechanical behavior, identify failure modes, and establish practical design recommendations for use in hybrid façade systems.
<table><caption>**Table 1**
Test matrix.</caption><thead><tr><th>No.</th><th>Specimen</th><th>U-Profile</th><th>Rebar</th><th>RC block size (cm)</th><th>Load</th></tr></thead><tbody><tr><td>1</td><td>P</td><td>P</td><td>No</td><td>60 × 46 × 20</td><td>W-</td></tr><tr><td>2</td><td>PM</td><td>PM</td><td>No</td><td>60 × 46 × 20</td><td>W-</td></tr><tr><td>3</td><td>MC</td><td>MC</td><td>No</td><td>60 × 46 × 20</td><td>W-</td></tr><tr><td>4</td><td>MC HA</td><td>MC</td><td>Yes</td><td>60 × 46 × 20</td><td>W-</td></tr><tr><td>5</td><td>MCL-1</td><td rowspan="2">MCL</td><td>No</td><td>60 × 46 × 20</td><td>W-</td></tr><tr><td>6</td><td>MCL-2</td><td>No</td><td>60 × 46 × 20</td><td>W-</td></tr><tr><td>7</td><td>DCV-1</td><td rowspan="2">DCV</td><td>No</td><td>60 × 46 × 20</td><td>W-</td></tr><tr><td>8</td><td>DCV-2</td><td>No</td><td>60 × 46 × 20</td><td>W-</td></tr><tr><td>9</td><td>DCV HA</td><td>DCV</td><td>Yes</td><td>60 × 46 × 20</td><td>W-</td></tr><tr><td>10</td><td>DCV R-1</td><td rowspan="5">DCV R</td><td>No</td><td>60 × 46 × 20</td><td>W-</td></tr><tr><td>11</td><td>DCV R-2</td><td>No</td><td>60 × 66 × 20</td><td>W-</td></tr><tr><td>12</td><td>DCV R-3</td><td>No</td><td>60 × 66 × 20</td><td>W-</td></tr><tr><td>13</td><td>DCV R-4</td><td>No</td><td>60 × 66 × 20</td><td>W-</td></tr><tr><td>14</td><td>DCV R-5</td><td>No</td><td>60 × 66 × 20</td><td>W-</td></tr><tr><td>15</td><td>DCV R-6</td><td>No</td><td>No</td><td>60 × 66 × 20</td><td>W-</td></tr></tbody></table>

Noted: (W-) Wind suction load.

## 1.4. Methodology

A three-phase approach combining experimental investigation, numerical modeling, and analytical study was adopted. In the experimental phase, fifteen concrete slab specimens incorporating six connector configurations were tested under monotonic wind loading to failure. In the numerical phase, a validated 3D finite element (FE) model was developed in LS-DYNA [27]. A parametric study was then conducted to investigate the influence of key design parameters. In the third phase, based on an analytical study of similar anchorage types, design equations for practical implementation were developed and validated against both experimental and numerical results.

## 2. Experimental program

This study evaluates the anchorage performance of a novel connector system for TFFs embedded in structural slabs subjected to wind suction (W-) loads. The experimental program investigates the influence of connector geometry, reinforcement, slab dimensions, and material properties.

### 2.1. Description of the specimens

As shown in Fig. 2, all U-profile specimens have a uniform length of 189.5 mm, feature two cutouts, and are embedded 110 mm into the reinforced concrete slabs. The specimens were categorized into two main groups based on embedment characteristics and geometric modifications: (i) standard specimens, including P, PM, MC, MC HA, MCL, DCV, and DCV HA; and (ii) modified specimens with additional flanges, designated as DCV R. Specimens labeled “HA” (e.g., MC HA, DCV HA) included reinforcement bars within the U-profile cutouts, originally intended to enhance anchorage performance.

The baseline P specimen includes two circular cutouts (φ20 mm) located at distances of $h_1 = 20$ mm and $h_2 = 40$ mm from the top edge of the U-profile, as shown in Fig. 2. Variants were developed to assess the effects of cutout width (PM), shape (MC), increased embedment length (MCL), and additional reinforcement (HA series). The DCV configuration increases the cutout width from 25 mm to 30 mm, while the DCV R adds 40 mm-wide flanges to improve structural resistance.

Each specimen consists of two reinforced concrete blocks (A and B), with each block anchored to a steel base using four M16 bolts, representing structural slabs (see Fig. 3(a)). This symmetrical A-B configuration was adopted for all tests to ensure uniform load transfer and consistency across specimens. The blocks have dimensions of either 600 × 460 × 200 mm or 600 × 660 × 200 mm, enabling investigation of slab size effects on connector performance. The test matrix and loading conditions are summarized in Table 1. MCL and DCV specimens were tested in duplicate, while DCV R specimens were tested six times with varying block sizes. All tests were conducted under negative pressure (W-) conditions.

### 2.2. Fabrication

U-profiles were fabricated by Metal Industrie, and concrete casting was performed at Ouest Préfa. Prior to casting, profiles were greased to minimize friction. Concrete was poured horizontally with mechanical vibration to ensure compaction. Reinforcement cages, produced by Ouest Arma, followed consistent layouts adapted to block size, as illustrated in Fig. 4.

Detailed reinforcement drawings are provided in Fig. 22. Three reinforcement configurations were employed: (1) The standard layout, used for specimens P, PM, MC, MCL, DCV, and DCV R1 (Fig. 22a), consisting of orthogonal bars without any reinforcement
<figure><img src="image_6.png" /><figcaption>Fig. 2. U-Profile specimen configurations.</figcaption></figure>

crossing the U-profile cutout; (2) The cross-reinforced layout, adopted for MC HA and DCV HA specimens (Fig. 22b), which incorporated an additional HA6 bar passing through the U-profile cutout and connected to the main cage, aiming to enhance confinement and anchorage performance near the connector; and (3) The extended layout, applied to enlarged blocks (DCV R2–6, Fig. 22c), following the same reinforcement pattern as Type (1) but with wider spacing and an additional outer frame to accommodate the increased slab width of 660 mm.
<figure><img src="image_7.png" /><figcaption>Fig. 3. Test setup and measurement instrumentation.</figcaption></figure>

## 2.3. Material properties

Concrete strength was determined from compression tests on three 110 × 220 mm cylinders per specimen, performed on the test day in accordance with EN 12390-3 [28]. The measured average strengths ($f_{cm}$) ranged from 19.1 to 36.2 MPa, corresponding to curing ages between 4 and 54 days. This variation reflects the testing schedule adopted to accelerate the experimental program and to obtain a wider range of $f_c$ values from the same concrete mix. Results are summarized in Table 2.
<figure><img src="image_8.png" /><figcaption>Fig. 4. Construction process of specimens.</figcaption></figure>

The U-profiles were made of S235 steel, and their tensile properties were evaluated in accordance with ISO 6892-1 [29]. The measured yield strength ($f_y$) was 398.3 MPa, determined from the 0.2% proof stress ($R_{p0.2}$). Reinforcement bars (HA6, HA8, HA10) were characterized based on the mechanical property data reported in Kouakou et al. [30], with average yield strengths of 410 MPa (HA6), 411 MPa (HA8), and 368 MPa (HA10).

## 2.4. Experimental setup and instrumentation

The test setup (see Fig. 3(a)) employed a hydraulic jack to apply upward wind-suction loads through a central HEB 100 loading beam. The load was transferred to the U-profiles via hinged and T-hinged connections, while each specimen was anchored to a steel base using M16 bolts.

The loading protocol followed the general recommendations of Eurocode 4 [31]. The applied loading history, shown in Fig. 3(b), comprised two initial preconditioning cycles (0–5 kN), followed by 25 force-controlled cycles between 5% and approximately 40% of the estimated ultimate load, corresponding to the Serviceability Limit State (SLS) range. The target ultimate load was estimated based on preliminary analytical evaluations and reference formulations for perfobond-type connectors [22,26].

After completing the cyclic stage, specimens were subjected to monotonic displacement-controlled loading until failure at a constant rate of approximately 1.2 mm/min under quasi-static conditions.

Four linear variable differential transformers (LVDTs), labeled C1–C4, were positioned on the free length of the U-profile, 50 mm from the concrete block surface, to monitor displacement during testing.

## 3. Test results and discussion

This section presents the main experimental results. The results are analyzed in terms of load-deflection curves, failure mechanisms, and crack patterns.
<table><caption>**Table 2**
Material properties.</caption><thead><tr><th>No.</th><th>Components</th><th>Specimen</th><th>Age of concrete [Day]</th><th>f<sub>cm</sub> [MPa]</th><th>f<sub>y</sub> [MPa]</th><th>S<sub>d</sub></th></tr></thead><tbody><tr><td>1</td><td></td><td>P</td><td>5</td><td>27.0</td><td>-</td><td>0.7</td></tr><tr><td>2</td><td></td><td>PM</td><td>4</td><td>28.4</td><td>-</td><td>-</td></tr><tr><td>3</td><td></td><td>MC</td><td>5</td><td>24.9</td><td>-</td><td>-</td></tr><tr><td>4</td><td></td><td>MC HA</td><td>6</td><td>32.6</td><td>-</td><td>0.7</td></tr><tr><td>5</td><td></td><td>MCL-1</td><td>4</td><td>19.1</td><td>-</td><td>0.31</td></tr><tr><td>6</td><td></td><td>MCL-2</td><td>54</td><td>36.2</td><td>-</td><td>0.94</td></tr><tr><td>7</td><td rowspan="2">Concrete block</td><td>DCV-1</td><td>5</td><td>29.3</td><td>-</td><td>2.73</td></tr><tr><td>8</td><td>DCV-2</td><td>6</td><td>25.7</td><td>-</td><td>0.59</td></tr><tr><td>9</td><td rowspan="2">A and B</td><td>DCV HA</td><td>4</td><td>25.2</td><td>-</td><td>0.60</td></tr><tr><td>10</td><td>DCV R-1</td><td>4</td><td>24.6</td><td>-</td><td>0.41</td></tr><tr><td>11</td><td></td><td>DCV R-2</td><td>5</td><td>25.2</td><td>-</td><td>0.53</td></tr><tr><td>12</td><td></td><td>DCV R-3</td><td>5</td><td>25.2</td><td>-</td><td>0.53</td></tr><tr><td>13</td><td></td><td>DCV R-4</td><td>4</td><td>25.7</td><td>-</td><td>0.74</td></tr><tr><td>14</td><td></td><td>DCV R-5</td><td>4</td><td>25.7</td><td>-</td><td>0.74</td></tr><tr><td>15</td><td></td><td>DCV R-6</td><td>4</td><td>24.9</td><td>-</td><td>0.46</td></tr><tr><td>-</td><td>S235 U Profiles</td><td>All specimens</td><td>-</td><td></td><td>398.3</td><td>9.43</td></tr><tr><td>-</td><td>Reinforcement HA6</td><td>All specimens</td><td>-</td><td></td><td>410</td><td>-</td></tr><tr><td>-</td><td>Reinforcement HA8</td><td>All specimens</td><td>-</td><td></td><td>411</td><td>-</td></tr><tr><td>-</td><td>Reinforcement HA10</td><td>All specimens</td><td>-</td><td></td><td>368</td><td>-</td></tr></tbody></table>

Noted: $f_{cm}$: Average concrete strength; $f_y$: Yield strength; $S_d$: Standard deviation.

## 3.1. Failure mode and crack patterns

Two primary failure mechanisms were observed under suction wind loading (W–), identified based on crack propagation patterns and post-failure observations.

* **Mechanism 1 - Cone breakout:** This was the dominant failure mode across all configurations except MCL. Cracks initiated at the interface between the U-profile flanges and the surrounding concrete and propagated outward to form a conical fracture surface (Fig. 5). Load transfer occurred from the U-profile through the concrete dowels within the cutouts into the slab. Cross-sectional inspections of the MC specimen (Sections C–C and D–D in Fig. 6) confirmed that the cone apex was located near the lower edge of the U-profile, with fracture surfaces extending beyond the flange width, indicating efficient anchorage and well-distributed stresses. Section D–D further showed cracks initiating near the cutout edge and propagating diagonally to the free surface, but the U-profile exhibited no plastic deformation. Consistently, FE results demonstrated von Mises stresses at peak load below the yield strength of S235 steel ($f_y$ = 398 MPa, Fig. 15), confirming that steel yielding did not govern. Instead, the tensile resistance of the concrete was exhausted first, leading to cone breakout.

* **Mechanism 2 - Localized cracking:** This mode occurred only in one block of the MCL specimens. Cracks were confined between the U-profile flanges without developing into a full cone (Fig. 5d). Sectioned specimens revealed crack initiation at the cutout edges, where stress concentrations triggered premature crushing and detachment of the concrete dowels. Because the MCL configuration had a deeper embedment than MC, larger anchorage forces were mobilized through the same cutout geometry. This increased the demand on the concrete dowels and, in some cases, led to premature cracking at the cutout edges that suppressed full cone development. In contrast, MC specimens with shallower embedment and the same cutout geometry consistently developed stable cone breakout, albeit with lower peak loads but larger displacement at failure. DCV specimens, which shared the same embedment depth as MCL but had wider cutouts (30 mm vs. 25 mm), distributed stresses more uniformly and also developed full cone breakout. These comparisons confirm that localized cracking in MCL arises from the combined effect of deeper embedment and narrower cutout geometry, placing it in a transitional state between stable cone breakout (MC, DCV) and localized splitting.

## 3.2. Load–deflection response

The load–deflection behavior of the U-profile fastening system under suction wind loads (W–) was analyzed to investigate the influence of concrete compressive strength, embedment depth, cutout width, reinforcement, and concrete block width on anchorage resistance and stiffness. The deflection reported in the load–deflection curves represents the average of four LVDT measurements recorded during the tests.

### 3.2.1. Effect of concrete compressive strength on anchorage resistance:

Concrete strength had a clear impact on peak load capacity. Increasing $f_{cm}$ from 19.1 MPa to 36.2 MPa in specimens MCL-1 and MCL-2 resulted in a 26% increase in peak load (29.88 kN to 37.69 kN), while the displacement at peak load remained similar (Fig. 7), suggesting strength affects resistance more than deformability.
<figure><img src="image_9.png" /><figcaption>Fig. 5. Failure mode and cracks map of specimens under negative wind loads.</figcaption></figure>

<figure><img src="image_10.png" /><figcaption>Fig. 6. Typical concrete cone shapes of the specimens (MC specimen).</figcaption></figure>

### 3.2.2. Effect of embedment depth

Greater embedment depths improved both strength and stiffness. Comparing DCV-2 ($h_1$ = 66 mm) with MC ($h_1$ = 46 mm), peak load increased by 31% (29.37 kN to 38.43 kN), and displacement at peak reduced from 2.2 mm to 0.75 mm (Fig. 8(a)). Similarly, MC HA ($h_1$ = 46 mm) showed 43% higher resistance than P ($h_1$ = 40 mm), confirming the role of larger concrete cones in enhancing anchorage performance (Fig. 8(b)).
<figure><img src="image_11.png" /><figcaption>Fig. 7. Effect of concrete compressive strength on anchorage resistance.</figcaption></figure>

<figure><img src="image_12.png" /><figcaption>Fig. 8. Effect of embedment depth on anchorage resistance (noted: $h_1$: maximum embedment depth of the U-Profiles).</figcaption></figure>

### 3.2.3. Effect of the width of cutouts of the U-profiles

The impact of cutout width depended on embedment depth. In shallow embedments, P ($W_c$ = 20 mm) and PM ($W_c$ = 30 mm) showed similar peak loads (24.20 kN vs. 23.12 kN) and stiffness, indicating negligible influence (Fig. 9(a)). However, for deeper profiles, DCV specimens ($W_c$ = 30 mm) outperformed MCL ($W_c$ = 25 mm), despite the latter's higher $f_{cm}$, due to localized cracking in MCL that limited cone development (Fig. 9(b)).

### 3.2.4. Reinforcement within cutouts

Reinforcement passing through the cutouts had a negligible effect on the anchorage capacity. As shown by the comparison between DCV-HA and DCV-2, both specimens failed by concrete cone breakout with nearly identical crack patterns (Fig. 5f). Their peak loads were almost identical (39.02 kN vs. 38.13 kN) and exhibited similar displacements at peak (~0.75 mm), suggesting that the small variation observed was mainly due to material variability rather than the presence of reinforcement (Fig. 10(a)). Similarly, MC HA showed about 18% higher resistance than MC; however, this increase was associated with higher concrete strength ($f_{cm}$) rather than the added bar (Fig. 10(b)). The bars inserted through the cutouts remained elastic throughout the tests due to their small diameter, confirming that the load-displacement response and ductility were not affected. This observation aligns well with the findings of Sun et al. [26], who reported that rebars inserted in perfobond connector holes contribute negligibly to pull-out resistance, as the failure mechanism is dominated by concrete cone breakout rather than bar yielding or bond action. Overall, the anchorage resistance in the present study was primarily governed by the concrete cone breakout mechanism, rather than by the reinforcement passing through the cutout.
<figure><img src="image_13.png" /><figcaption>Fig. 9. Effect of the width ($W_c$) of cutouts of the U-Profiles on anchorage resistance.</figcaption></figure>

<figure><img src="image_14.png" /><figcaption>Fig. 10. Effect of reinforcement within cutouts on anchorage resistance.</figcaption></figure>

### 3.2.5. Effect of concrete block width

Concrete block width showed limited influence. DCV R-1 (460 mm) and DCV R-2 to R-6 (660 mm) had peak loads ranging from 40.47 to 45.14 kN (Fig. 11), a variation within experimental scatter, indicating block width did not significantly affect anchorage performance.

### 3.2.6. Effect of expanded flanges of the U-profiles

The addition of transverse flanges improved resistance. DCV R-1, with two 40 mm flanges, achieved a 15% higher peak load (44.14 kN) than DCV-2 (38.43 kN), despite slightly lower $f_{cm}$ (Fig. 12). DCV R-1 also had one-third the displacement at peak, attributed to increased contact area and better mechanical interlock.

Overall, embedment depth and concrete strength are the most influential parameters enhancing anchorage capacity and stiffness. Cutout geometry and flange extensions provide additional benefits, while reinforcement within cutouts and block width have negligible impact.

## 4. Finite element modeling

FE analysis is a widely adopted method in structural engineering for investigating complex parameters not easily accessible through experiments. It has been effectively applied to both steel [32,33] and concrete structures [34]. To better understand the mechanical behavior of the proposed U-profile connector, an FE study was performed. It is noted that numerical modeling focused on
<figure><img src="image_15.png" /><figcaption>Fig. 11. Effect of concrete block width on anchorage resistance (DCV R1: block width = 460 m; DCV R2-6: block width = 660 mm.</figcaption></figure>

<figure><img src="image_16.png" /><figcaption>Fig. 12. Effect of expanded flanges of the U-Profiles (DCV-2: without expanded flanges; DCV R1: with two 40 mm expanded flanges.</figcaption></figure>

representative configurations (MCL, DCV, DCV R) that captured key geometric and mechanical variations. The P and PM specimens, serving as baseline references, were excluded from FEA due to computational constraints and limited additional benefit to model validation.

## 4.1. Finite element modeling of the tested specimens

The simulations were carried out in LS-DYNA [27], which is well-suited for modeling nonlinear material behavior and contact interactions. To reduce computational cost, symmetry was utilized by modeling only half of the specimen. The model, built in LS-PrePost, consisted of six components: (1) concrete block, (2) concrete dowels, (3) U-profile, (4) threaded rods, (5) reinforcement bars, and (6) rigid loading plate (Fig. 13).

## 4.2. Finite element model development

### 4.2.1. Element type and mesh size

All major components (concrete, dowels, U-profile, bolts, and loading plate) were meshed using 8-node solid elements. Reinforcement bars were modeled with beam elements. An adaptive meshing approach was adopted: a fine 1 mm mesh was applied through the U-profile thickness, while 5 mm elements were used along its length and width to capture stress concentrations. A 5 mm mesh was also assigned to the loading plate, dowels, reinforcement, and the surrounding concrete. A coarser 10 mm mesh was used in less critical regions, such as the bolts and lower concrete volume, based on experimental observations.
<figure><img src="image_17.png" /><figcaption>Fig. 13. The FE model of the specimen DCV in LS-DYNA.</figcaption></figure>

<table><caption>**Table 3**
Input parameters for Winfrith concrete model.</caption><thead><tr><th>E<sub>c</sub> (MPa)</th><th>f<sub>c</sub> (MPa)</th><th>f<sub>t</sub> (MPa)</th><th>ν</th><th>ρ (g/mm<sup>3</sup>)</th><th>w<sub>c</sub> (mm)</th><th>d<sub>max</sub> (mm)</th></tr></thead><tbody><tr><td>32,000–35,000</td><td>19–36</td><td>2.1–3.3</td><td>0.17</td><td>0.0023</td><td>0.1</td><td>32</td></tr></tbody></table>

### 4.2.2. Contact and boundary conditions

Vertical displacement was applied to the U-profile via a rigid plate using the \*PRESCRIBED\_MOTION\_RIGID command. The bottom nodes of the threaded rods were fixed to ensure global stability.

Contact between the rigid plate and U-profile, as well as between dowels and the U-profile, was modeled using \*CON- TACT\_AUTOMATIC\_SURFACE\_TO\_SURFACE. Bolts were bonded to the concrete block via node merging. Reinforcement bars were embedded in concrete using the \*CONSTRAINED\_LAGRANGE\_IN\_SOLID method [35], assuming perfect bond and neglecting bond-slip for simplification.

### 4.2.3. Concrete material model

The Winfrith Concrete Model (MAT_084/085) was used to simulate nonlinear concrete behavior, including crack initiation and propagation, due to its proven effectiveness in prior studies [34,36]. Key inputs include compressive strength $f_c$, modulus $E_c$, Poisson's ratio, tensile strength $f_t$, crack width $w_c$, and maximum aggregate size $d_{max}$. Parameters were derived from tests or estimated using CEB-FIP formulas [37].

The tensile strength and elastic modulus were computed as:

$$f_t = 0.3(f_{cm})^{2/3}, \quad E_c = 22 \left( \frac{f_{cm}}{10} \right)^{0.3}.$$
<table><caption>**Table 4**
Steel properties for MAT_003.</caption><thead><tr><th>Material</th><th><i>E</i> (MPa)</th><th><i>f</i><sub>y</sub> (MPa)</th><th><i>E</i><sub>tan</sub> (MPa)</th><th><i>ν</i></th><th><i>ρ</i> (kg/m<sup>3</sup>)</th><th><i>β</i></th></tr></thead><tbody><tr><td>Rebar</td><td>205,000</td><td>400</td><td>41,700</td><td>0.30</td><td>7850</td><td>1</td></tr><tr><td>U-profile</td><td>205,000</td><td>398.3</td><td>20,600</td><td>0.30</td><td>7850</td><td>1</td></tr></tbody></table>

<table><caption>**Table 5**
Summary of the numerical and experimental results.</caption><thead><tr><th rowspan="2">No.</th><th rowspan="2">Specimen</th><th colspan="3">Peak load (kN)</th><th colspan="3">Displacement (mm)</th></tr><tr><th>Experiment</th><th>FEA</th><th>Error (%)</th><th>Experiment</th><th>FEA</th><th>Error (%)</th></tr></thead><tbody><tr><td>1</td><td>MC</td><td>29.37</td><td>29.9</td><td>1.80</td><td>2.16</td><td>2.82</td><td>30.56</td></tr><tr><td>2</td><td>MCL-1</td><td>29.88</td><td>31.87</td><td>6.66</td><td>0.60</td><td>0.48</td><td>20.00</td></tr><tr><td>3</td><td>DCV-1</td><td>41.48</td><td>41.33</td><td>0.40</td><td>0.68</td><td>0.72</td><td>5.88</td></tr><tr><td>4</td><td>DCV-2</td><td>38.43</td><td>39.33</td><td>2.34</td><td>0.78</td><td>0.77</td><td>1.28</td></tr><tr><td>5</td><td>DCV R-1</td><td>44.14</td><td>45.12</td><td>2.22</td><td>0.37</td><td>0.43</td><td>16.22</td></tr><tr><td>6</td><td>DCV R-3</td><td>45.14</td><td>46.01</td><td>1.93</td><td>0.35</td><td>0.43</td><td>22.86</td></tr></tbody></table>

<table><caption>**Table 6**
Comparison of initial stiffness obtained from FE simulation and experimental tests.</caption><thead><tr><th rowspan="2">No.</th><th rowspan="2">Specimen</th><th colspan="2">Initial stiffness (kN/mm)</th><th rowspan="2">Error (%)</th></tr><tr><th>Experiment results</th><th>FE results</th></tr></thead><tbody><tr><td>1</td><td>MC</td><td>41.45</td><td>36.18</td><td>12.71</td></tr><tr><td>2</td><td>MCL-1</td><td>160.41</td><td>166.39</td><td>3.73</td></tr><tr><td>3</td><td>DCV-1</td><td>154.62</td><td>126.75</td><td>18.02</td></tr><tr><td>4</td><td>DCV-2</td><td>172.69</td><td>121.75</td><td>29.50</td></tr><tr><td>5</td><td>DCV R-1</td><td>291.70</td><td>223.92</td><td>23.24</td></tr><tr><td>6</td><td>DCV R-2</td><td>523.37</td><td>460.32</td><td>12.05</td></tr></tbody></table>

Fracture energy under uniaxial tension $G_{Ft}$ and crack width $w_c$ were estimated based on aggregate size and compressive strength:

$$G_{Ft} = G_{F0} \left( \frac{f_{cm}}{f_{cm0}} \right)^{0.7}, \quad \text{for } f_{cm} \le 50 \text{ MPa,} \qquad (1)$$

$$G_{Ft} = G_{F0} \ln \left( 1 + \frac{f_{cm}}{f_{cm0}} \right), \quad \text{for } f_{cm} > 50 \text{ MPa.} \qquad (2)$$

where $f_{cm0} = 10$ MPa, and the base fracture energy ($G_{F0}$) depends on the maximum aggregate size ($d_{max}$):

$$G_{F0} = 0.021 + 5.357 \times 10^{-4} d_{max}. \qquad (3)$$

The critical crack width ($w_c$) was determined as:

$$w_c = \frac{2G_f}{f_f}. \qquad (4)$$

Compressive behavior followed an elastic-plastic law with Ottosen's yield criterion [38]. Strain-rate effects were ignored (RATE = 1). The input data for the Winfrith concrete used in this study are presented in Table 3, where $v$ is the Poisson's ratio and $\rho$ is the density of concrete [39].

### 4.2.4. Steel material model

Steel components such as the U-profile and reinforcement were modeled using the bilinear Plastic-Kinematic model MAT_003 [40] (see Fig. 13). The material parameters are listed in Table 4, where $E$ is the elastic modulus, $f_y$ is the yield strength, $E_{tan}$ is the tangent modulus, $v$ is the Poisson's ratio, $\rho$ is the material density, and $\beta$ is the hardening parameter, which defines the model as either isotropic hardening ($\beta = 1$) or kinematic hardening ($\beta = 0$). The loading plate was modeled as a rigid body using MAT_020.

### 4.3. Finite element results and validation

The numerical results were validated against experimental data in terms of load-deflection responses and crack patterns.

#### 4.3.1. Load-deflection response

Fig. 14 compares experimental and FE load-deflection curves. As shown in Table 5, the FE models predict peak loads with high accuracy (error < 7%), and displacements with reasonable agreement (max. error ~30%). Notably, the DCV-1 and DCV R-1 specimens exhibited the closest match.

Initial stiffness values were determined using a linear regression method, applied to the initial portion of the load-displacement curve up to 10% of the peak load, in order to capture the elastic response of the specimens [41]. As summarized in Table 6,
<figure><img src="image_18.png" /><figcaption>(a) MC specimen</figcaption></figure>

<figure><img src="image_19.png" /><figcaption>(b) MCL-1 specimen</figcaption></figure>

<figure><img src="image_20.png" /><figcaption>(c) DCV-1 specimens</figcaption></figure>

<figure><img src="image_21.png" /><figcaption>(d) DCV-2 specimens</figcaption></figure>

<figure><img src="image_22.png" /><figcaption>(e) DCV R-1 specimens</figcaption></figure>

<figure><img src="image_23.png" /><figcaption>(f) DCV R-2 specimens</figcaption></figure>

<caption>**Fig. 14.** Comparison between experimental and numerical results.</caption>

The FE-predicted initial stiffness shows acceptable consistency with the experimental measurements, with a maximum deviation of 29.5%.

The numerical models effectively captured both peak loads and the general shape of response curves. Minor discrepancies in stiffness likely stem from modeling assumptions, such as rigid contact and perfect bond.
<figure><img src="image_24.png" /><figcaption>Fig. 15. FE results: von Mises equivalent stress distribution in U-profiles at peak load.</figcaption></figure>

<caption>**Table 7**</caption>

Pull-out anchorage resistance models for perfobond connectors.

<table><thead><tr><th>Researcher</th><th>Calculation formula</th></tr></thead><tbody><tr><td>Zheng et al. [22]</td><td>N<sub>r</sub> = γ<sub>w</sub> [0.95 (d<sub>h</sub><sup>2</sup> - d<sub>r</sub><sup>2</sup>) f<sub>c</sub> + 0.45d<sub>r</sub><sup>2</sup> f<sub>ry</sub> + 0.18d<sub>h</sub>t<sub>p</sub>f<sub>sy</sub>]</td></tr><tr><td>Sun et al. [26]</td><td>N<sub>r</sub> = (3u + d<sub>h</sub>)<sup>2</sup> · 3(u + <sup>d<sub>h</sub></sup>&frasl;<sub>3</sub>)<sup>-0.56</sup> √f<sub>c</sub></td></tr></tbody></table>

Note: $N_r$ is the pull-out capacity (N); $d_h$, $d_r$: hole and rebar diameters (mm); $t_p$: steel plate thickness (mm); $f_c$, $f_{ry}$, $f_{sy}$: material strengths (MPa); $u$: embedment depth (mm).

### 4.3.2. Damage and cracks prediction

Von Mises stress contours of the U-profiles at peak load (Fig. 15) remained below the yield strength of S235 steel ($f_y$ = 398 MPa). This further confirms that anchorage failure was governed by concrete cone breakout rather than steel yielding.

Crack patterns from FE simulations were compared with experiments for selected specimens (Fig. 16). The FE model reproduced the failure modes with reasonable accuracy:

* MC specimen: Failed to capture full cone failure, showing localized crushing instead.

* MCL-1 specimen: Accurately reflected local crushing and asymmetric cracks.

* DCV-2, DCV R-1 specimens: Captured cone-shaped propagation and interface cracks well.

* DCV R-2 specimen: Showed asymmetry but underestimated crack spread angle.

Differences may be attributed to simplified concrete material models, especially in the post-peak softening range. Nonetheless, the simulations offered valuable insight into internal stress evolution, cone formation, and anchorage failure progression, confirming their validity for parametric studies.

## 5. Parametric analysis and discussion

The validated LS-DYNA model was used to conduct a parametric study on the effects of concrete strength and U-profile embedment depth. These parameters were varied independently while maintaining constant material and geometric configurations elsewhere.

### 5.1. Effect of the compressive strength of the concrete blocks

A parametric study was carried out to investigate how varying the concrete compressive strength affects the performance of MC, DCV, and DCV R-1 specimens. All other material properties and geometric parameters were kept constant.

As shown in Figs. 17(a, c, e), increasing the concrete strength from 16 MPa to 40 MPa significantly improved both peak load and stiffness across all configurations.
<figure><img src="image_25.png" /><figcaption>Fig. 16. Comparison of the crack failure of specimens in the test and numerical model.</figcaption></figure>

Figs. 17(b, d, f) show a strong linear relationship between the peak load and the square root of compressive strength ($\sqrt{f_{cm}}$), with $R^2$ values ranging from 0.986 to 1.000. This trend aligns with previous studies on concrete cone failure in anchorage systems, such as perfobond rib connectors [26], and design guidelines from CEB [42], fib [43], and ACI 318-14 [44], further validating the FE model. The rate of load increase also depends on U-profile geometry. The DCV specimen ($h_1$ = 66 mm) showed a higher slope (8.2) than the MC specimen ($h_1$ = 46 mm, slope 7.1). The DCV R-1, which incorporates an expanded plate, exhibited the highest slope (12.4), indicating enhanced interaction between the profile and surrounding concrete. These results highlight the combined role of concrete strength and connector geometry in improving tensile resistance.
<figure><img src="image_26.png" /><figcaption>(a) Load-deflection response - MC</figcaption></figure>

<figure><img src="image_27.png" /><figcaption>(c) Load-deflection response - DCV</figcaption></figure>

<figure><img src="image_28.png" /><figcaption>(e) Load-deflection response - DCVR1</figcaption></figure>

<figure><img src="image_29.png" /><figcaption>(b) Concrete strength vs peak load - MC</figcaption></figure>

<figure><img src="image_30.png" /><figcaption>(d) Concrete strength vs peak load - DCV</figcaption></figure>

<figure><img src="image_31.png" /><figcaption>(f) Concrete strength vs peak load - DCVR1</figcaption></figure>

<caption>Fig. 17. Effect of the compressive strength.</caption>

## 5.2. Effect of the embedment depth of the U-profiles

A parametric study was conducted to assess the impact of embedment depth on the performance of the standard DCV specimen and the modified DCV R-1 configuration. The concrete compressive strength $f_{cm}$ was kept constant, while the effective embedment depth $h_{ef}$ was varied by shifting the anchor cutout position relative to the concrete surface [45].
<figure><img src="image_32.png" /><figcaption>(a) Concrete strength vs peak load-DCV</figcaption></figure>

<figure><img src="image_33.png" /><figcaption>(c) Load-deflection response-DCV R1</figcaption></figure>

<figure><img src="image_34.png" /><figcaption>(b) Concrete strength vs peak load-DCV</figcaption></figure>

<figure><img src="image_35.png" /><figcaption>(d) Concrete strength vs peak load-DCV R1</figcaption></figure>

<caption>Fig. 18. Effect of the embedment depth of the U-Profiles.</caption>

The effective embedment depth was defined as the average between the minimum and maximum values (see Fig. 6), as shown in Eq. (5):

$$h_{ef} = \frac{h_{\min} + h_{\max}}{2} \qquad (5)$$

As seen in Figs. 18(a, b), increasing embedment depth from 53 mm to 103 mm significantly improved the load capacity of the DCV system (from 35.83 kN to 60.02 kN), with a nearly linear trend ($R^2 = 0.996$). Stiffness also increased proportionally.

For the DCV R-1 configuration (Figs. 18(c, d)), the load capacity rose from 29.6 kN to 58.6 kN over the same embedment range, showing a strong linear correlation ($R^2 = 0.947$), though stiffness remained largely unchanged. This behavior suggests that while deeper embedment improves anchorage, the ultimate capacity may be governed by the strength of the external U-profile portion, particularly if its stiffness is not increased correspondingly.

### 5.3. Global variation analysis

To complement the individual parametric observations, a global sensitivity analysis was conducted to quantify the relative influence of key parameters on the anchorage resistance. The analysis employed both the experimental and FE datasets (Table 8), excluding specimens with flange extensions (DCV-R) to ensure a consistent cone breakout mechanism. Three independent variables were considered: the square root of concrete compressive strength ($\sqrt{f_c}$), the equivalent embedment depth ($h_{eq}$), and the cutout width ($W_c$).

A standardized linear regression model was applied to assess the contribution of each variable to the variation in peak load. Results indicate that $h_{eq}$ and $\sqrt{f_c}$ together account for approximately 99% of the total variance in anchorage resistance, confirming
<figure><img src="image_36.png" /><figcaption>Fig. 19. Global sensitivity ranking of key parameters on anchorage resistance based on standardized regression.</figcaption></figure>

that these two parameters predominantly govern the anchorage behavior. The effect of cutout width was minor (< 5%), consistent with the experimental observations discussed in Section 3.2.3. The normalized influence of each parameter is illustrated in Fig. 19.

These findings quantitatively demonstrate that the anchorage capacity is primarily controlled by the concrete strength and effective embedment depth, while geometric details such as the cutout width play a secondary role. This sensitivity-based evaluation provides a rational foundation for the simplified analytical expression proposed in Section 6, where $h_{eq}$ and $\sqrt{f_c}$ are adopted as the governing parameters for estimating the anchorage resistance.

# 6. Analytical study

## 6.1. Concrete cone breakout resistance of a single anchor

The concrete cone breakout strength is a classical failure mode for anchors subjected to direct tension and provides a theoretical basis for evaluating pull-out behavior. One of the earliest models, introduced in ACI 349 [46], assumes a 45° failure cone with uniform tensile stress and leads to a breakout capacity proportional to $h_{ef}^2$. However, this model tends to overestimate capacity for large embedment depths due to the size effect [47].

To address this, the Concrete Capacity (CC) method [48] introduces a 35° failure cone and accounts for the size effect by reducing the exponent on $h_{ef}$ to 1.5. This method is widely adopted in standards such as ACI 318-14 [44] and EN 1992-4 [45]. The generalized form is:

$$N_{u,m} = \alpha \cdot \sqrt{f_c} \cdot h_{ef}^{\beta} \qquad (6)$$

where $\alpha$ and $\beta$ are empirical coefficients depending on $h_{ef}$. EN 1992-4 further introduces correction factors for geometry and loading conditions:

$$N_r = k_1 \sqrt{f_c} h_{ef}^{1.5} \frac{A_{c,N}}{A_{c,N}^0} \psi_{s,N} \psi_{re,N} \psi_{ec,N} \psi_{M,N} \qquad (7)$$

These models form the basis for evaluating the pull-out resistance of embedded components.

## 6.2. Pull-out anchorage resistance of perfobond connectors

While perfobond connectors have been extensively studied under shear [21,24,49], limited attention has been paid to their performance under tensile (pull-out) forces [22,26]. Inspired by the CC method, several models have been proposed to estimate their anchorage capacity under axial tension, as summarized in Table 7. These approaches assume centrally loaded connectors positioned away from concrete edges, neglecting edge effects and cone overlap.

## 6.3. Proposed analytical model

An analytical model is developed to estimate the pull-out resistance of standard U-profile fixing systems under suction wind loading, based on the CC method. The formulation draws from the anchorage behavior of single bolts and perfobond connectors.

Unlike existing models (Table 7), the proposed model accounts for eccentric loading effects specific to U-profiles. The force distribution is shown in Fig. 20.
<figure><img src="image_37.png" /><figcaption>Fig. 20. Force distribution on U-profiles under suction wind load.</figcaption></figure>

The moment equilibrium is established between the anchorage resistance of the U-profile (F) and the anchorage resistance of its cutouts ($F_p$).

The resistance of the cutouts is proposed as:

$$F_p = 2\alpha h_{eq} \sqrt{f_c} \qquad (8)$$

The anchorage resistance of the U-profile is given by:

$$F = 2 \frac{\alpha h_{eq} \sqrt{f_c} l_1}{l_2} \quad (N) \qquad (9)$$

where $\alpha$ is a calibration coefficient; $h_{eq}$ is the equivalent embedment depth (mm); $f_c$ is the concrete compressive strength (MPa); and $l_1$, $l_2$ are geometric moment arms (mm). A linear dependence on $h_{eq}$ is adopted to represent the partial-cone breakout observed in the U-profile tests (Section 3) and FEA simulations (Section 5), in contrast to the nonlinear $h_{ef}^{1.5}$ dependence typically used in CCD anchor models.

The equivalent embedment depth is defined as (see Fig. 20):

$$h_{eq} = \frac{h_{min} + h_{max}}{2} \quad (mm) \qquad (10)$$

## 6.4. Model validation

The parameter $\alpha = 134$ was calibrated using test data from standard U-profile specimens (excluding DCV-R). Validation against experimental results showed prediction errors between -17.5% and +16.7% (Fig. 21(a)), with a mean ratio of 0.96 and $R^2 = 0.682$.

For the FE simulations (Table 8), the model yielded errors from -16.7% to +10.6%, with a mean ratio of -0.05 and $R^2 = 0.942$ (Fig. 21(b)). Overall, the model offers a reasonable and slightly conservative prediction across the evaluated cases.

It should be noted that the proposed formulation is empirical, relying on a single calibrated coefficient ($\alpha = 134$) obtained by regression of the tested specimens and further confirmed by FEA simulations. While the model provides reasonable agreement with the database, scatter remains significant ($R^2 = 0.682$), reflecting variability of cone cracking and dowel-concrete detachment. The applicability of the present equations is limited to standard U-profiles without transverse flanges, as DCV-R specimens with enlarged flanges exhibited additional confinement mechanisms not captured by the model. Moreover, edge effects were not considered in the test program, since the connectors were centrally placed and block width was found negligible; however, in real slabs, slab-edge proximity may truncate the breakout cone and reduce capacity, which is beyond the scope of the current model. Finally, unlike CCD-based approaches for headed anchors, which follow $N \propto h_{ef}^{1.5}$ due to full-cone breakout and fracture mechanics size effect [50–52], the U-profile connectors tested here developed partial-cone breakout. Consequently, the anchorage resistance increased nearly linearly with the equivalent embedment depth ($h_{eq}$), as confirmed by both experimental and FE results.
<figure><img src="image_38.png" /><figcaption>(a) Experimental results</figcaption></figure>

<figure><img src="image_39.png" /><figcaption>(b) FE results</figcaption></figure>

<caption>Fig. 21. Comparison of predicted anchorage resistance with test and FE results.</caption>

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

Comparison of anchorage resistance predicted by the analytical model with experimental and FE results.</caption><thead><tr><th rowspan="2">No.</th><th rowspan="2">Specimen</th><th><i>l</i><sub>1</sub></th><th><i>l</i><sub>2</sub></th><th><i>h</i><sub>min</sub></th><th><i>h</i><sub>max</sub></th><th><i>f</i><sub>c</sub></th><th colspan="3">Anchorage resistance (kN)</th><th colspan="2">Error (%)</th></tr><tr><th>[mm]</th><th>[mm]</th><th>[mm]</th><th>[mm]</th><th>[Mpa]</th><th>Experimental</th><th>FEA</th><th>Analytical model</th><th>Experimental</th><th>FEA</th></tr></thead><tbody><tr><td>1</td><td>MC</td><td>68</td><td>157</td><td>40</td><td>46</td><td>16</td><td>-</td><td>21,14</td><td>19,97</td><td>-</td><td>5,56</td></tr><tr><td>2</td><td>MC</td><td>68</td><td>157</td><td>40</td><td>46</td><td>24,9</td><td>29,37</td><td>29,9</td><td>24,91</td><td>15,20</td><td>16,70</td></tr><tr><td>3</td><td>MC</td><td>68</td><td>157</td><td>40</td><td>46</td><td>40</td><td>-</td><td>37,74</td><td>31,57</td><td>-</td><td>16,35</td></tr><tr><td>4</td><td>MCL</td><td>68</td><td>157</td><td>60</td><td>66</td><td>19,1</td><td>29,88</td><td>31,87</td><td>31,96</td><td>6,96</td><td>0,28</td></tr><tr><td>5</td><td>DCV</td><td>68</td><td>157</td><td>60</td><td>66</td><td>10</td><td>-</td><td>23,45</td><td>23,13</td><td>-</td><td>1,39</td></tr><tr><td>6</td><td>DCV</td><td>68</td><td>157</td><td>60</td><td>66</td><td>16</td><td>-</td><td>30,57</td><td>29,25</td><td>-</td><td>4,31</td></tr><tr><td>7</td><td>DCV</td><td>68</td><td>157</td><td>60</td><td>66</td><td>25,7</td><td>38,43</td><td>39,33</td><td>37,07</td><td>3,53</td><td>5,74</td></tr><tr><td>8</td><td>DCV</td><td>68</td><td>157</td><td>60</td><td>66</td><td>29,3</td><td>41,48</td><td>41,33</td><td>39,58</td><td>4,57</td><td>4,22</td></tr><tr><td>9</td><td>DCV</td><td>68</td><td>157</td><td>60</td><td>66</td><td>40</td><td>-</td><td>47,97</td><td>46,25</td><td>-</td><td>3,58</td></tr><tr><td>10</td><td>DCV</td><td>68</td><td>177</td><td>60</td><td>66</td><td>29,3</td><td>-</td><td>36,4</td><td>35,11</td><td>-</td><td>3,54</td></tr><tr><td>11</td><td>DCV</td><td>68</td><td>137</td><td>60</td><td>66</td><td>29,3</td><td>-</td><td>47,2</td><td>45,36</td><td>-</td><td>3,89</td></tr><tr><td>12</td><td>DCV</td><td>68</td><td>157</td><td>40</td><td>46</td><td>29,3</td><td>-</td><td>28</td><td>27,02</td><td>-</td><td>3,51</td></tr><tr><td>13</td><td>DCV</td><td>68</td><td>157</td><td>100</td><td>106</td><td>29,3</td><td>-</td><td>60</td><td>64,72</td><td>-</td><td>7,86</td></tr><tr><td>14</td><td>DCV</td><td>68</td><td>157</td><td>40</td><td>46</td><td>16</td><td>-</td><td>21,8</td><td>19,97</td><td>-</td><td>8,42</td></tr><tr><td>15</td><td>DCV</td><td>68</td><td>157</td><td>40</td><td>46</td><td>50</td><td>-</td><td>38,2</td><td>35,29</td><td>-</td><td>7,61</td></tr><tr><td>16</td><td>DCV</td><td>68</td><td>157</td><td>70</td><td>76</td><td>29,3</td><td>-</td><td>46,6</td><td>45,87</td><td>-</td><td>1,57</td></tr><tr><td>17</td><td>DCV</td><td>68</td><td>157</td><td>50</td><td>56</td><td>29,3</td><td>-</td><td>35,8</td><td>33,30</td><td>-</td><td>6,98</td></tr><tr><td>18</td><td>DCV</td><td>68</td><td>157</td><td>90</td><td>96</td><td>29,3</td><td>-</td><td>56,6</td><td>58,43</td><td>-</td><td>3,24</td></tr><tr><td>19</td><td>DCV</td><td>68</td><td>157</td><td>90</td><td>96</td><td>15</td><td>-</td><td>37,8</td><td>41,81</td><td>-</td><td>10,61</td></tr><tr><td>20</td><td>DCV</td><td>68</td><td>157</td><td>50</td><td>56</td><td>40</td><td>-</td><td>47,2</td><td>38,91</td><td>-</td><td>17,57</td></tr><tr><td>21</td><td>DCV</td><td>68</td><td>157</td><td>90</td><td>96</td><td>35</td><td>-</td><td>69,6</td><td>63,86</td><td>-</td><td>8,24</td></tr><tr><td>22</td><td>DCV</td><td>68</td><td>157</td><td>50</td><td>56</td><td>20</td><td>-</td><td>30,6</td><td>27,51</td><td>-</td><td>10,09</td></tr><tr><td>23</td><td>DCV</td><td>68</td><td>157</td><td>90</td><td>96</td><td>40</td><td>-</td><td>77,2</td><td>68,27</td><td>-</td><td>11,56</td></tr><tr><td>24</td><td>P</td><td>70</td><td>157</td><td>34</td><td>40</td><td>27</td><td>24,2</td><td>-</td><td>22,97</td><td>5,07</td><td>-</td></tr><tr><td>25</td><td>PM</td><td>65</td><td>157</td><td>34</td><td>40</td><td>28,4</td><td>23,12</td><td>-</td><td>21,88</td><td>5,37</td><td>-</td></tr><tr><td>26</td><td>MC HA</td><td>68</td><td>157</td><td>40</td><td>46</td><td>32,63</td><td>34,56</td><td>-</td><td>28,51</td><td>17,50</td><td>-</td></tr><tr><td>27</td><td>MCL</td><td>68</td><td>157</td><td>60</td><td>66</td><td>36,2</td><td>37,69</td><td>-</td><td>44,00</td><td>16,74</td><td>-</td></tr><tr><td>28</td><td>DCV HA</td><td>68</td><td>157</td><td>60</td><td>66</td><td>25,2</td><td>39,02</td><td>-</td><td>36,71</td><td>5,92</td><td>-</td></tr></tbody></table>

## 7. Conclusions

This study experimentally and numerically evaluated the anchorage performance of a novel U-profile fastening system designed to connect timber frame façades (TFFs) to concrete structures. Fifteen full-scale specimens, incorporating variations in U-profile geometry, embedment depth, and concrete strength, were tested under suction wind loading scenarios. The key findings, which are limited to the configurations examined in this study, are summarized as follows:

* Under suction wind loading, failure was primarily governed by the surrounding concrete. Cracks initiated near the embedded U-profiles and propagated outward in a typical concrete cone failure pattern.

* Anchorage resistance was governed by concrete cone breakout rather than yielding of the U-profile. Performance was primarily controlled by the capacity of the surrounding concrete cone, with embedment depth, cutout geometry, flange configuration of the U-profile, and concrete compressive strength identified as the key influencing parameters.
<figure><img src="image_40.png" /><figcaption>(a) Specimens P, PM, MC, MCL, DCV, and DCV R1</figcaption></figure>

<figure><img src="image_41.png" /><figcaption>(b) Specimens MC HA and DCV HA</figcaption></figure>

<figure><img src="image_42.png" /><figcaption>(c) DCV R2-6</figcaption></figure>

<caption>**Fig. 22.** Reinforcement configurations of concrete blocks.</caption>

* Variations in cutout width, including configurations with expanded flange cutouts, had minimal impact on anchorage resistance. Similarly, the inclusion of rebar within the cutout provided limited benefit, as neither modification significantly increased the effective concrete volume contributing to cone formation.
* Finite element simulations using LS-DYNA accurately reproduced the load-displacement responses and failure modes of the specimens. Peak load predictions deviated by no more than 7.0% from experimental results, and the models effectively captured both the initial stiffness and crack evolution.

* An analytical model was developed with reference to the Concrete Capacity method and relevant standards (EN 1992-4, ACI 318). The formulation incorporates embedment depth and concrete compressive strength and provides a practical tool for estimating the pull-out resistance of standard U-profile connectors. The model showed reasonable agreement with experimental and FE results, but remains empirical and is limited to standard U-profiles, neglecting edge effects and excluding configurations with flanges (DCV-R).

These results underscore the importance of embedment depth and concrete compressive strength in determining anchorage performance, while demonstrating that modifications to cutout geometry and internal reinforcement offer limited structural advantages. The analytical formulation, supported by both experimental and numerical validation, provide a reliable framework for the structural design of U-profile connectors.

A key limitation of this study is that local failure mechanisms - such as concrete crushing at the steel - concrete interface or localized yielding of the U-profile - were not explicitly investigated. These mechanisms may become critical under different geometric or loading conditions and should be addressed in future work. Additionally, the long-term performance of the fastening system under fatigue and cyclic wind loading warrants further investigation to ensure durability in service applications.

### CRediT authorship contribution statement

**Tran-Van Han:** Writing – original draft, Validation, Software, Methodology, Investigation, Data curation, Formal analysis, Visualization. **Alice Le Berder:** Writing – review & editing, Methodology, Formal analysis, Data curation, Conceptualization, Investigation, Validation, Visualization. **Tuan-Anh Nguyen:** Writing – review & editing, Validation, Software, Investigation, Formal analysis, Data curation, Methodology, Conceptualization. **Clemence Lepourry-Nicollet:** Writing – review & editing, Conceptualization, Investigation, Methodology, Project administration. **Hugues Somja:** Writing – review & editing, Supervision, Resources, Project administration, Methodology, Investigation, Formal analysis, Conceptualization, Funding acquisition.

### 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.

### Acknowledgments

The authors gratefully acknowledge financial support from the French National Research Agency (ANR, Paris) with grant number ANR-21-CHIN-0003 through the ANR FREEINBTP project.

### Appendix

See Fig. 22 and Table 8.

### Data availability

Data will be made available on request.

### References

[1] H.V. Tran, S.C. Kim, J. Shin, K. Lee, Shaking table testing of an innovative wireway vibration attenuation system, J. Constr. Steel Res. 197 (2022) 107477, http://dx.doi.org/10.1016/j.jcsr.2022.107477.

[2] E. Fema, Reducing the Risks of Nonstructural Earthquake Damage A Practical Guide, Federal Emergency Management Agency, Washington, DC, 2011.

[3] G. Di Nunzio, O. Al-Mansouri, G. Muciaccia, A. Gabs, In-plane behaviour of lightweight façade systems installed in masonry and concrete under simulated seismic loading, J. Build. Eng. 63 (2023) 105442, http://dx.doi.org/10.1016/j.jobe.2022.105442.

[4] A. Baird, A. Palermo, S. Pampanin, P. Riccio, A. Tasligedik, Focusing on reducing the earthquake damage to facade systems, Bull. N. Z. Soc. Earthq. Eng. 44 (2) (2011) 108-120, http://dx.doi.org/10.5459/bnzsee.44.2.108-120.

[5] X. Chong, Y.-F. Guo, J.-Q. Huang, Q. Jiang, M. Zhao, J.-G. Dai, Seismic performance of a steel frame structure with dissipative cladding connections, J. Constr. Steel Res. 197 (2022) 107508, http://dx.doi.org/10.1016/j.jcsr.2022.107508.

[6] S. Tulebekova, K.A. Malo, A. Rønnquist, Dynamic identification and model calibration of connection stiffness in multi-storey cross-laminated timber buildings, J. Build. Eng. 72 (2023) 106607, http://dx.doi.org/10.1016/j.jobe.2023.106607.

[7] J. Yang, S. Liang, X. Zhu, L. Dang, W. Wang, B. Sun, Research on point-supported rotational semi-rigid energy-dissipation connection for precast-cladding-panel at connection level, J. Constr. Steel Res. 214 (2024) 108455, http://dx.doi.org/10.1016/j.jcsr.2024.108455.

[8] J.-W. Baek, S.-M. Kang, H.-G. Park, Direct shear test and cyclic loading test for seismic capacity of dimension stone panel cladding with dowel pin connection, J. Build. Eng. 38 (2021) 102205, http://dx.doi.org/10.1016/j.jobe.2021.102205.
[9] J.-W. Baek, H.-J. Hwang, S.-M. Kang, H.-G. Park, Shaking table test for dimension stone cladding with dowel pin connection, J. Build. Eng. 70 (2023) 106390, http://dx.doi.org/10.1016/j.jobe.2023.106390.

[10] A. Aloisio, F. Boggian, R. Tomasi, Design of a novel seismic retrofitting system for RC structures based on asymmetric friction connections and CLT panels, Eng. Struct. 254 (2022) 113807, http://dx.doi.org/10.1016/j.engstruct.2021.113807.

[11] B. Huang, J. Shi, Z. Yang, H. Dong, Hysteric behavior and seismic analysis of granite cladding with dowel-pin connection, Eng. Struct. 322 (2025) 119060, http://dx.doi.org/10.1016/j.engstruct.2024.119060.

[12] R. Eligehausen, R. Mallée, J.F. Silva, Anchorage in Concrete Construction, vol. 10, John Wiley & Sons, 2006.

[13] L. Xie, H. Sha, X. Chong, L. Hou, Y. Guo, Performance-based seismic design of RC frame structure with energy dissipative cladding panel connection system using equivalent energy design procedure, J. Build. Eng. 59 (2022) 105076, http://dx.doi.org/10.1016/j.jobe.2022.105076.

[14] H. Sha, X. Chong, L. Xie, L. He, Q. Jiang, J. Gao, Study on seismic performance of RC frame with UHPC cladding panels and two-level friction damping connections, J. Build. Eng. 95 (2024) 110118, http://dx.doi.org/10.1016/j.jobe.2024.110118.

[15] W. Sui, L. Zhu, H. Yang, Z. Wang, Y. Ma, Cyclic behaviour of friction energy dissipation connection systems with precast concrete cladding panel: Experimental and numerical studies, J. Build. Eng. 98 (2024) 111301, http://dx.doi.org/10.1016/j.jobe.2024.111301.

[16] L. Ostetto, J. Fonseca, R. Sousa, P. Fernandes, H. Rodrigues, Experimental campaign of horizontal cladding wall panels connection in industrial precast reinforced concrete buildings, Eng. Struct. 322 (2025) 119106, http://dx.doi.org/10.1016/j.engstruct.2024.119106.

[17] G. Di Nunzio, N. Pinoteau, G. Muciaccia, T. Guillet, R. Mège, Derivation of an innovative testing protocol for plastic anchors for fixing light façade claddings under seismic action, Eng. Struct. 248 (2021) 113191.

[18] Ministry of Ecological Transition and Territorial Cohesion of France, Environmental regulation 2020, 2022, Published by the Government of France.

[19] Directive 2010/31/EU on the energy performance of buildings (EPBD), 2010, https://energy.ec.europa.eu/topics/energy-efficiency/energy-efficient-buildings/energy-performance-buildings-directive_en. (Accessed 23 May 2025).

[20] Act on the improvement of energy consumption performance of buildings, 2015, https://www.mlit.go.jp/en/toshi/building-energy.html. (Accessed 23 May 2025).

[21] F. Leonhardt, W. Andrä, H.-P. Andrä, W. Harre, Neues, vorteilhaftes verbundmittel für stahlverbund-tragwerke mit hoher dauerfestigkeit, Beton-Und Stahlbetonbau 82 (12) (1987) 325-331, http://dx.doi.org/10.1002/best.198700500.

[22] S. Zheng, Y. Liu, Y. Liu, C. Zhao, Experimental and parametric study on the pull-out resistance of a notched perfobond shear connector, Appl. Sci. 9 (4) (2019) 764, http://dx.doi.org/10.3390/app9040764.

[23] T. Hosaka, K. Mitsuki, H. Hiragi, Y. Ushijima, Y. Tachibana, H. Watanabe, An experimental study on shear characteristics of perfobond strip and its rational strength equations, J. Struct. Eng. JSECE 46 (2000) 1593-1604.

[24] E. Oguejiofor, M. Hosain, Numerical analysis of push-out specimens with perfobond rib connectors, Comput. Struct. 62 (4) (1997) 617-624, http://dx.doi.org/10.1016/S0045-7949(96)00270-2.

[25] J.-H. Ahn, C.-G. Lee, J.-H. Won, S.-H. Kim, Shear resistance of the perfobond-rib shear connector depending on concrete strength and rib arrangement, J. Constr. Steel Res. 66 (10) (2010) 1295-1307, http://dx.doi.org/10.1016/j.csr.2010.04.008.

[26] L. Sun, Y. Liu, F. Shi, H. Wang, Pull-out performance of perfobond rib connectors in steel high-strength concrete composite bridge pylons, Constr. Build. Mater. 366 (2023) 130205, http://dx.doi.org/10.1016/j.conbuildmat.2022.130205.

[27] L.S.T. Corporation, LS-DYNA keyword user's manual, version R10.0, LTC, Livermore, CA, 2017.

[28] B. En, 12390-3:(2009). Testing hardened concrete. Compressive strength of test specimens, 2019, British Standard Institution, London 16.

[29] International Organization for Standardization, ISO 6892-1:2016 - metallic materials — Tensile testing — Part 1: Method of test at room temperature, 2016.

[30] C.H. Kouakou, I. Bakayoko, M.A. Serifou, E. Emeruwa, Mechanical Characteristics Test of Concrete Steel Bars Available in Côte d'Ivoire, J. Mater. Sci. Chem. Eng. 8 (9) (2020) 1-13.

[31] CEN-European Committee for Standardization, Eurocode 4: Design of Composite Steel and Concrete Structures - Part 1-1: General Rules and Rules for Buildings, CEN-European Committee for Standardization, Brussels, Belgium, 2005.

[32] T.-V. Han, M. Jeong, Y. Kim, D. Lee, N.H. Cuong, K. Lee, Experimental and numerical evaluation of the structural performance of novel S-CN connections in modular construction, in: Structures, vol. 71, Elsevier, 2025, 107930, http://dx.doi.org/10.1016/j.istruc.2024.107930.

[33] T.-V. Han, J. MoonSook, Y. Kim, D. Lee, N.-V. Luat, K. Lee, Numerical study on compressive behavior of a 3D-printed sand mold-cast S-CN connector, J. Constr. Steel Res. 213 (2024) 108336, http://dx.doi.org/10.1016/j.cjsr.2023.108336.

[34] N.H. Cuong, H. An, H.V. Tran, S. An, K. Lee, Structural test and analyses of UHPC horizontal connection with shear keys for slab-external PC wall, Constr. Build. Mater. 419 (2024) 135410, http://dx.doi.org/10.1016/j.conbuildmat.2024.135410.

[35] M. Abdallah, P. Heng, F. Palas, H. Somja, Mechanical behavior of a new thermal break system for exterior wall-slab joints under monotonic loading, Eng. Struct. 303 (2024) 117563, http://dx.doi.org/10.1016/j.engstruct.2024.117563.

[36] S. Azad, S.R. Mirhaderi, S. Epackachi, Numerical investigation of steel and composite beam-to-encased composite column connection via a through-plate, in: Structures, vol. 31, Elsevier, 2021, pp. 14-28, http://dx.doi.org/10.1016/j.istruc.2021.01.040.

[37] Du Beton, Comité Euro-International, CEB-FIP model code 1990, Des. Code (1991) 54-58.

[38] N.S. Ottosen, A failure criterion for concrete, J. Eng. Mech. Div. 103 (4) (1977) 527-535, http://dx.doi.org/10.1061/JMCEA3.000224.

[39] V.H. Tran, Winfrith model input generator, 2025, http://dx.doi.org/10.5281/zenodo.14883590.

[40] N.-V. Luat, J. Shin, J.-H. Park, K. Lee, Experimental and numerical investigation on seismic performance of retrofitted RC columns with web direct/indirect bonding external H-section, J. Build. Eng. 44 (2021) 103404, http://dx.doi.org/10.1016/j.jobe.2021.103404.

[41] A.M. Billah, M.S. Alam, Seismic performance of concrete columns reinforced with hybrid shape memory alloy (SMA) and fiber reinforced polymer (FRP) bars, Constr. Build. Mater. 28 (1) (2012) 730-742, http://dx.doi.org/10.1016/j.conbuildmat.2011.10.020.

[42] Comité Euro-International du Béton (CEB), Design of Fastenings in Concrete, Thomas Telford, Lausanne, Switzerland, 1997, p. 92, CEB Design Guide.

[43] Fédération Internationale du Béton (fib), Design of Anchorages in Concrete: Guide to Good Practice, vol. 58, International Federation for Structural Concrete, Lausanne, 2011, p. 280.

[44] A.C. 318, Building Code Requirements for Structural Concrete and Commentary (ACI 318-14), American Concrete Institute, Farmington Hills, MI, 2014, p. 524.

[45] EN 1992-4, Eurocode 2: Design of concrete structures-part 4: Design of fastenings for use in concrete, 2018.

[46] A.C. Institute, ACI 349-90: Code Requirements for Nuclear Safety-Related Concrete Structures, American Concrete Institute, Detroit, MI, 1990, ACI Committee 349.

[47] Z.P. Bažant, Size effect in blunt fracture: concrete, rock, metal, J. Eng. Mech. 110 (4) (1984) 518-535, http://dx.doi.org/10.1061/(ASCE)0733-9399(1984)110:4(518).

[48] W. Fuchs, R. Eligehausen, J.E. Breen, Concrete capacity design (CCD) approach for fastening to concrete, Struct. J. 92 (1) (1995) 73-94.

[49] Y. Liu, Experimental and analytical investigation of Naning steel-concrete composite bridge arch, vol. 72, Bridge Engineering Department, Tongji University, China, 2006.

[50] R. Eligehausen, P. Bouška, V. Červenka, R. Pukl, Size effect of the concrete cone failure load of anchor bolts, in: Z.P. Bažant (Ed.), Fracture Mechanics of Concrete Structures, FRAMCOS 1, Elsevier Applied Science, Breckenridge, 1992, pp. 517-525.

[51] R. Eligehausen, R. Mallee, G. Rehm, Befestigungstechnik, Furche, J. (1994). Zum Trag- und Verschiebungsverhalten von Kopfbolzen bei zentralischem Zug Dissertation, University of Stuttgart, Ernst & Sohn, Berlin, Germany, 1997, See also.

[52] J. Ožbolt, R. Eligehausen, H.W. Reinhardt, Size effect on the concrete cone pull-out load, Int. J. Fract. 95 (1) (1999) 391-404.