Experimental Evaluation of Influence of Member Thickness, Anchor-Head Size, and Orthogonal Surface Reinforcement on the Tensile Capacity of Headed Anchors in Uncracked Concrete

The mean tensile breakout capacity of headed anchors can be determined by the CC method proposed by Fuchs et al. (1995). In the United States, this method is known as the concrete capacity design (CCD) method. According to the CC method, the mean tensile breakout capacity of a single cast-in-place anchor in uncracked concrete far from the concrete free edges and/or adjacent anchors can be evaluated as follows:where $N_{u,m}$ = mean concrete cone breakout capacity of a single anchor {N [international system (SI)] and lb [United States customary units (US)]}; $f_c$ = concrete cylinder compressive strength [MPa (SI) and psi (US)]; and $h_{\mathrm{ef}}$ = anchor effective embedment depth [mm (SI) and in. (US)].

This method has been incorporated into several design standards, such as ACI 349 Appendix D (ACI 2006) and ACI 318 (ACI 2014), in the United States as well as various design-oriented documents in Europe [e.g., Comité Euro-International du Béton, CEB Design guide (CEB 1997) and CEN/TS 1992-4 (CEN 2009a)] and internationally in the fib Bulletin 58 (fib 2011). The CC method assumes a concrete cone angle of approximately 35° with respect to the concrete surface, which leads to an idealized projected cone area of approximately $3.0h_{ef}\times 3.0h_{ef}$ on the concrete surface (Fuchs et al. 1995).

For deep anchors (where $h_{ef}\ge 280\mathrm{mm}$), ACI 349 Appendix D (ACI 2006) and ACI 318 (ACI 2014) allow use of a modified CC method, which is given in Eqs. (2a) and (2b). The modified method uses an exponent of 5/3 ($=1.667$) rather than 1.5 for the effective embedment depth of deep anchors ($h_{ef}\ge 280\mathrm{mm}$) and appropriately changes the leading coefficient of the CC method

The application of Eqs. (2a) and (2b) is limited by the concrete compressive strength and anchor embedment depth, which have maximum allowable values of 70 MPa and 635 mm, respectively.

The CC method is an empirical model based a large number of pullout tests on anchors at various embedment depths. The tested anchors were housed in unreinforced and reinforced concrete members of various thicknesses, leading to a wide scatter in the obtained capacities. However, the CC method was derived based on the mean values (i.e., 50% fractile) of the scatter data. Fig. 2(a) shows the ratios of measured concrete cone breakout capacities to the values predicted by the CC method [Eq. (1)] as a function of anchor embedment depth for 320 pullout tests on single cast-in-place headed anchors in literature. These data were compiled and evaluated from Zhao (1993), Eligehausen et al. (1992), and Nilsson et al. (2011). The tested anchor bolts were embedded in uncracked concrete members and positioned far from the concrete free edges and adjacent anchors.

In addition, the tested anchor bolts had different bearing sizes, and, therefore, the concrete in the vicinity of anchor bolts experienced different bearing stresses; the bearing stress under anchor heads (${\sigma }_b$) varied from 1.3 to $20.2·f_c$. The ratio of bearing stress under the anchor head at peak load to the concrete compressive strength (${\sigma }_b/f_c$) for all test data in Fig. 2(a) was calculated and plotted in Fig. 2(b) as a function of anchor embedment depth. As the figure shows, the tested short anchors ($h_{ef}\le 100\mathrm{mm}$) had relatively large heads [mean ratio of bearing stress to concrete compressive strength (${\sigma }_b/f_c$) was approximately 4.6], whereas the tested deep anchors ($h_{ef}>100\mathrm{mm}$) had fairly small heads [mean ratio of bearing stress to concrete compressive strength (${\sigma }_b/f_c$) was approximately 14.8]. This figure explicitly indicates that the CC method was not developed based on a consistent size of anchor heads throughout the entire range of anchor embedment depths studied because the bearing stress under the head of anchors was considerably variable. In fact, if short anchors ($h_{ef}\le 100\mathrm{mm}$) had been tested with smaller heads, they would have been failed with lower capacities. In addition, large anchors ($h_{ef}>100\mathrm{mm}$) would have been failed with higher capacities if they have been tested with larger heads.

In practice, anchor bolts with various head sizes are often installed in concrete members with various bending stiffnesses. In some cases, the CC method [Eq. (1)] may underestimate the anchorage capacity [this may be the case for deep headed anchors with large heads in thick reinforced concrete (RC) members] or even may overestimate the anchorage capacity (this can be the case for short headed anchors with small heads in relatively thin unreinforced concrete members). These can be seen in Fig. 2(a), where there are scatter data even below the 5% fractile of test results (i.e., design values that correspond to 75% of the mean capacities at tests) and also above the 95% fractile of the test results (i.e., considered 125% of the mean capacities obtained at tests).

Nilforoush et al. (2017a, b) systematically investigated, via numerical simulations based on nonlinear fracture mechanics, the influence of member thickness, anchor-head size, and surface reinforcement on the tensile breakout capacity of headed anchors. Cast-in-place headed anchors at various embedment depths ($h_{ef}=50–500\mathrm{mm}$) were simulated in plain and RC members of various thicknesses ($H=1.5–5.0h_{ef}$). The simulated anchors also had various head sizes [ratio of the bearing stress under the anchor head at peak load to the concrete cylinder compressive strength (${\sigma }_b/f_c$) varied from $\sim 4$ to 20].

Based on these studies, it was found that the tensile breakout capacity of an anchor bolt increases with increasing thickness of the concrete member or if the concrete member is orthogonally reinforced. The numerical results showed that the tensile breakout capacity further increases by increasing the bearing area of the anchor head. Moreover, it was found that the CC method [Eq. (1)] underestimates the tensile capacity of deep-headed anchors ($h_{ef}\ge 280\mathrm{mm}$), whereas it overestimates the capacity of short anchors ($h_{ef}\le 100\mathrm{mm}$) with fairly small heads in relatively thin plain members. This can be seen in Fig. 3, where the numerical results of series (a) from Nilforoush et al. (2017b) for headed anchors with small heads at various embedment depths are presented. This figure shows the relationship between the ratio of cone breakout capacity to the square root of concrete compressive strength ($N_u/f_c^{0.5}$) and anchor embedment depth. Fig. 3 also shows the corresponding ratios predicted by the CC method [Eq. (1)]. A trend line fitted to the numerical results revealed that the ratio of concrete cone breakout capacity to the square root of concrete compressive strength ($N_u/f_c^{0.5}$) is proportional to $h_{ef}^{1.68}$ (i.e., $\sim h_{ef}^{5/3}$).

To account for the overestimation of concrete cone breakout capacity of short-headed anchors, Nilforoush et al. (2017a, b) recommended use of the modified CC method [Eq. (2b)], associated with the deep anchors ($h_{ef}\ge 280\mathrm{mm}$), as well as for short embedment depths of anchors, as given in [Eq. (3)]

Eq. (3) is valid for a single headed anchor with a relatively small head (i.e., with a mean bearing stress of ${\sigma }_b=15·f_c$ at peak load under the anchor head) in an uncracked unreinforced concrete member with a thickness of $H=2.0h_{ef}$. Compared with other relations, this equation yielded smaller deviations over the entire range of anchor embedment depths ($h_{ef}=50–500\mathrm{mm}$) investigated. Furthermore, to account for the influence of member thickness, anchor-head size, and surface reinforcement on the anchorage capacity of headed anchors in uncracked concrete, Eq. (3) was extended by incorporating three modification factors (namely, ${\psi }_H$, ${\psi }_{AH}$, and ${\psi }_{Sr}$, respectively), which are defined as follows:withwhere $H$ = member thickness [mm (SI) and in. (US)]; $h_{ef}$ = anchor embedment depth [mm (SI) and in. (US)]; $A_b$ = anchor bearing area [${\mathrm{mm}}^2$ (SI) and sq in. (US)]; and $A_b^{\mathrm{code}}$ = code-equivalent bearing area corresponding to a bearing stress of ${\sigma }_b=15·f_c$ under the anchor head at peak load, which can be determined from [Eq. (5)]

${\psi }_H$ was limited to 1.20, based on the numerical results of headed anchors in plain concrete members of various thicknesses. The numerical results revealed that, for ${\psi }_H<1.0$, unreinforced concrete members fail by concrete splitting, whereas both RC and unreinforced members fail via concrete cone breakout for ${\psi }_H\ge 1.0$. Nilforoush et al. (2017a) found that the favorable influence of surface reinforcement on the anchorage capacity decreases with increasing thickness of the concrete member. In addition, it was concluded that the proposed ${\psi }_{Sr}$ factor is applicable if the concrete member is orthogonally reinforced and has a reinforcement-content of at least ($\rho =0.3\%$) in each direction.

In spite of a large number of experimental and theoretical investigations into the capacity of headed anchors, the influences of the global and local stress fields in the vicinity of anchors on the anchorage capacity are still unknown. There have been a few previous experimental and numerical studies on the influence of anchor embedment depth, anchor-head size, and surface reinforcement on the concrete cone resistance of anchor bolts (e.g., Eligehausen et al. 1992; Ožbolt et al. 1999, 2007; Baran et al. 2006; Lee et al. 2007; Nilsson et al. 2011) but none have systematically evaluated the influence of all these parameters.

Nilforoush et al. (2017b) compared some available test results from literature with the predictions of anchorage capacity using the proposed modification factors and found good agreement with the test results. However, further systematic experimental studies are needed to verify and generalize the proposed modification factors. The present experimental study aims to further demonstrate the validity of the numerically proposed modification factors for predicting the anchorage capacity of headed anchors with different head sizes in plain and reinforced concrete members of various geometries.

A total of 19 single cast-in-place headed anchors were tested under monotonic tensile loads. Testing parameters such as the concrete member thickness, anchor-head size, and presence of orthogonal surface reinforcement in the concrete member were considered. Table 1 provides the matrix of the experimental program, which consisted of three test series. In Series 1, headed anchors were tested in plain concrete (PC) members of various thicknesses (i.e., $H=330$, 440, and 660 mm) to evaluate the influence of member thickness. In Series 2, headed anchors with three different head sizes (i.e., small, medium, and large) were tested in PC members of identical thickness ($H=660\mathrm{mm}$) to evaluate the influence of the size of the anchor head. In Series 3, headed anchors were tested in RC members of various thicknesses ($H=330$, 440, and 660 mm) to evaluate the influence of surface reinforcement on the anchorage capacity and performance.

The test specimens were identified by names consisting of three parts. The first part specifies the concrete condition (PC or RC). The second part specifies the thickness of the concrete member, and the third part indicates the size of the anchor head ($\mathrm{S}=\text{small}$, $\mathrm{M}=\text{medium}$, and $\mathrm{L}=\text{large}$), followed by the test replicate number.

The typical geometry of the test specimens is shown in Fig. 4(a). In all cases, a single headed anchor was placed in the center of a concrete block. Concrete blocks with identical length ($L$) and width ($W$) were used for all specimens ($L=W=\mathrm{1,300}\mathrm{mm}$), whereas the height ($H$) of concrete blocks varied. The tested reinforced concrete specimens had mesh reinforcement at 150 mm spacing at the top and the bottom. The thickness of concrete cover on the top of reinforcements was 50 mm. A reinforcement ratio of $\rho \sim 0.3\%$ was applied for each direction in all reinforced specimens. The reinforcement configurations are listed in Table 1.

The test specimens were designed to fail via brittle failure modes associated with concrete (i.e., concrete cone breakout or splitting failure). Therefore, the tensile strength of headed anchors was designed sufficiently high to prevent steel failure. The anchors were composed of standard threaded 36-mm-diameter rods with a round bearing head at the end. The initial aim of the experiment was to test the anchors at an identical embedment depth (of $h_{ef}=220\mathrm{mm}$). However, after casting the concrete blocks, an effective anchor embedment depth of $h_{ef}=200\mathrm{mm}$ was realized for small-headed specimens (i.e., PC-660-S1 and PC-660-S2).

To determine the influence of size of the anchor head, three different sizes (small, medium, and large) of bearing head were tested. The geometry of headed anchors with small, medium, and large heads is shown in Fig. 4(b). For the small-headed and medium-headed anchors, round nuts with diameters of $d_h=48$ and 55 mm, respectively, were affixed to the end of threaded rods. For large-headed anchors, a thick circular steel plate with diameter $d_h=90\mathrm{mm}$ was tapped and fastened to the end of threaded rods. Although the threads of the steel rods fitted perfectly into the tapped steel plate, a standard hex nut was tightened underneath the steel plate to ensure that the plate remained in place during anchor pullout loading.

The threaded rods were covered with 2-mm-thick plastic tubes (Fig. 5). These tubes were used to prevent friction and adhesion between the anchor shaft and concrete body, thereby ensuring transfer of the entire anchor load through the anchor bearing head. The ratio of the bearing stress (${\sigma }_b$) under the head of anchors at peak load (predicted by the CC method) to the concrete cube compressive strength ($f_{cc}$) was determined. Values of 12.5, 7.1, and 1.6 were obtained for the ratio (${\sigma }_b/f_{cc}$) associated with small-headed, medium-headed, and large-headed anchors, respectively. According to Eligehausen et al. (2006), the concrete cylinder strength $f_c$ and cube strength $f_{cc}$ are related as $f_c\approx 0.85·f_{cc}$. This relation yields values of approximately 14.7, 8.4, and 1.8 for the ratio (${\sigma }_b/f_c$) associated with small-headed, medium-headed, and large-headed anchors, respectively.

Class B500B reinforcement (yield strength $f_{yk}=500\mathrm{MPa}$) was used in the concrete specimens. The headed anchors consisted of a Grade 10.9 steel rod with yield strength $f_{yk}$ and ultimate strength $f_{uk}$ (according to manufacturer’s specifications) of 900 and 1,000 MPa, respectively.

The concrete blocks were cast using ready-mixed normal-weight concrete made of crushed aggregates. The concrete mix design is summarized in Table 2. Immediately after casting, the concrete blocks were covered for 24 h by polyethylene sheets and then cured with wet burlap for an additional 7 days. The concrete blocks were kept indoors at an ambient temperature of 10°C until testing. To eliminate the influence of increased concrete strength at the time of testing, anchor pullout loading was conducted after curing the concrete for approximately 60 days. The compressive and tensile splitting strength of concrete were measured on additional concrete cubes (sides of 150 mm; at least five cubes were used for each test), in accordance with EN 12390-3 (CEN 2009b) and EN 12390-6 (CEN 2009c), respectively.

Furthermore, the concrete fracture energy was measured by means of a three-point bending loading test on standard concrete notched beams ($550\mathrm{mm}\text{long}\times 150\mathrm{mm}\text{wide}\times 150\mathrm{mm}$ high), in accordance with RILEM TC 50-FMC (1985). The concrete cubes and beams were cast from the same concrete batch as the concrete blocks and were cured and kept under the same conditions as the concrete blocks. The compressive and tensile splitting strength of concrete as well as the concrete fracture energy were measured the same day as the anchor pullout loading tests. After finishing the anchor pullout loading tests, the modulus of elasticity of concrete was also measured on drilled concrete cores (eight cores from the tested concrete blocks, four cores from PC blocks, and four cores from RC blocks) in accordance with SS 13 72 32 (Swedish Standard 2005). The cores had a diameter of $\sim 65\mathrm{mm}$ and height of $\sim 150\mathrm{mm}$. The mean values along with their corresponding coefficient of variations for the measured cube compressive strength ($f_{ccm}$), tensile splitting strength ($f_{ctm,sp}$), modulus of elasticity ($E_{cm}$), and fracture energy ($G_f$), are listed in Table 3.

The test setup is shown in Fig. 6(a) and schematically illustrated in Fig. 6(b). The anchor pullout loading was applied in an unconfined test setup; the span of vertical support was taken as sufficiently large to permit the formation of an unrestricted concrete cone. The vertical support was a stiff circular steel ring with an inner diameter of $L_{\sup }=880\mathrm{mm}$. The steel ring was directly placed on a thin gypsum layer on the concrete surface to accommodate the concrete surface roughness. Moreover, anchor pullout loading was performed on the top of the anchor shaft using a displacement-controlled system to capture the postpeak load-displacement behavior of anchor bolts. The loading was applied to the bolts by means of a hollow cylinder hydraulic jack with a capacity of 100 t. The load was applied using a 36-mm-diameter high-strength steel rod connected to the top of the anchor shaft by a coupling nut. The jack was placed on a stiff cross steel beam resting on the circular steel ring (Fig. 6). Prior to loading, the bolts were all manually preloaded to $\sim 30\mathrm{kN}$ using a mechanical tightening nut at the end of the loading rod to provide full confinement between the steel ring and testing concrete block. The bolts were then loaded under displacement control at a constant displacement rate of $1\mathrm{mm}/\min$ until peak load. After reaching the peak load, the rate of (this displacement-controlled) loading was increased to $2\mathrm{mm}/\min$.

The applied load was measured by a load cell placed on the top of the jack. For measuring the anchor displacement, a measurement platform fabricated from a square steel plate ($150\times 150\mathrm{mm}$) was secured perpendicular to the anchor bolt, 50 mm above the concrete surface, using a hex nut above and another below the platform [Fig. 6(b)]. The surface of the platform served as the reference level for anchor displacement, which was measured by two draw-wire displacement sensors installed symmetrically at the corners of the platform. These sensors measured the displacement of the anchors relative to two rigid points (i.e., rigid frames) outside the concrete block (i.e., on the solid ground). The frames were supported from outside the blocks, and contact with other structural members and testing equipment (which could have produced inaccurate displacement readings) was prevented. The sensors could measure displacements of up to 150 mm with an accuracy of $<0.1\%$. The axial displacement of the bolts was taken as the average of the two wire sensor measurements up to the end of loading.

To ensure precise displacement reading, the anchor displacement was simultaneously measured with two LVDTs installed symmetrically at the other corners of the measurement platform, 100 mm from the anchor. The LVDTs had a maximum traveling length of 25 mm and an accuracy of $<0.001\mathrm{mm}$. A data acquisition system was used to record the load and displacement data continuously. The anchor displacements measured by the sensors were very similar to those measured by the LVDTs, thereby confirming the precision of the displacement measurements recorded by these sensors.

Fig. 8 shows the load-displacement curves of headed anchors in PC members of various thicknesses. As can be seen, the anchor ultimate load and the anchor displacement at ultimate load increase slightly with increasing thickness of the concrete member. In addition, the anchorage stiffness decreases slightly with increasing member thickness, whereas the anchorage ductility improves slightly (Fig. 7). The postpeak anchorage behavior reveals that the anchors, especially those in the thinnest members, are quite brittle. This is evidenced by the sharp decline in the load-displacement curves, after peak loads, attributable to the development of rapid and unstable concrete cracks.

Fig. 9 shows the crack patterns at failure for the headed anchors embedded in PC members of various thicknesses. The anchors failed via concrete cone breakout, except for those in the thinnest members, which underwent mixed-mode concrete cone and splitting/bending failure. In fact, concrete cone circumferential cracking is initiated at the head of these anchors and propagate toward the concrete surface as the load increases. However, concrete bending/splitting cracking dominates the failure at anchor peak load and divides the concrete block into several separate pieces. This mixed-mode failure mechanism gives rise to more brittle behavior for headed anchors in the thinnest concrete members compared with those in thicker members.

With increasing member thickness and consequent increased global bending stiffness of the member, the failure mode changed from the mixed-mode concrete cone and splitting failure (observed in the thinnest PC members) to pure concrete cone breakout. This increase in member thickness prevents bending/splitting cracking but promotes circumferential concrete cone cracking, as previously indicated by Nilforoush et al. (2017a, b). Based on Eqs. (6) and (7), a critical member thickness ($H_{cr}$) at which the concrete bending cracking occurs for the plain concrete members can be evaluated

To calculate the critical bending cracking member thickness $H_{cr}$, the corresponding measured anchor failure loads $N_{u,\mathrm{test}}$ in plain concrete members of various thicknesses (i.e., Series 1) is substituted for $N_{\mathrm{bending}}$ in Eq. (10). This equation gives $H_{cr}$ values of 346, 375, and 422 mm, respectively, for the PC members with the thicknesses of $H=330$, 440, and 660 mm. Because the thickness of the thinnest concrete slabs ($H=330\mathrm{mm}$) is lower than its critical thickness ($H_{cr}=346\mathrm{mm}$), the concrete bending cracking, therefore, dominates the failure of these slabs. This indicates that the thickness of the thinnest slabs has to be increased to larger than 346 mm to prevent the concrete bending/splitting failure of these slabs.

The load-displacement curves of headed anchors with different head sizes are compared in Fig. 10(a). As Figs. 7 and 10(a) show, the anchorage stiffness and capacity increase significantly with increasing size of the anchor head, whereas the anchorage ductility and anchor displacement at peak load decrease considerably. In addition, the curves of the large-headed anchors are quite linear compared with those of medium-headed and small-headed anchors. A comparison of the postpeak behavior (i.e., after the peak load is reached) revealed that the large-headed anchors are quite brittle, undergoing sudden failure without prior noticeable anchor displacement, whereas the small-headed anchors are rather ductile, undergoing large deformations. To better understand the behavior of headed anchors with various head sizes, the relative capacity of anchors ($N/N_u$) are plotted in Fig. 10(b) as a function of relative anchor displacement ($\mathrm{\Delta }/{\mathrm{\Delta }}_u$). The relative capacities and displacements were determined by dividing the load-displacement curves of anchors with different head sizes to their corresponding peak loads and displacements at peak loads. As the figure shows, it seems that the concrete cone crack growth is stable up to anchor peak load for all head sizes. However, when anchor displacement exceeds the displacement at peak load, the crack growth becomes unstable for the medium-headed and large-headed anchors. This is evidenced by the sharp decline in the relative capacity of the medium-headed and large-headed anchors. On the contrary, the small-headed anchors showed larger relative capacity and displacement after peak load compared with their medium-headed and large-headed counterparts. This is because the concrete under the head of small-headed anchors was subjected to a high bearing stress of $\sim 14.7·f_c$ at ultimate load, which causes local crushing of the concrete under the anchor head and a consequent gradual anchor pullout

According to ACI 318 (ACI 2014), the characteristic bearing stress under the anchor head is limited to $8.0·f_c$ for the failure mode associated with the concrete cone breakout in cracked concrete. For anchors located in a region of a concrete member where no cracking is expected at service load levels, the characteristic bearing stress can be increased to $11.2·f_c$ ($8.0·f_c·1.4=11.2·f_c$). The characteristic value is equivalent to 5% fractile of the mean value, which is taken as $\sim 0.75\%$ of the mean value [ACI-318 (ACI 2014)]. If the characteristic bearing stress under the head is divided by 0.75, a mean allowable bearing stress of $\sim 15·f_c$ is obtained for anchors in uncracked concrete. If the mean bearing stress under the anchor head becomes larger than $15.0·f_c$, then the anchorage may be governed by pullout failure rather than concrete cone breakout failure.

Fig. 11 shows the crack patterns at failure for headed anchors with various head sizes. As the figure shows, all anchors (irrespective of head size) undergo failure via concrete cone breakout. The average concrete cone angle with respect to the loading direction increases with increasing head size. In addition, the diameter of the cone at the concrete surface is $\le 4.0h_{ef}$ for anchors with small and medium heads and $>4.0h_{ef}$ for anchors with large heads. This finding concurs with the results of the numerical study conducted by Nilforoush et al. (2017b).

Furthermore, in the case of large-headed anchors [Fig. 11(c)], the formation of concrete cone cracks is hindered by the vertical support, and the failure mode changes to bending cracking. The same behavior was observed in experimental and numerical studies on influence of size of the anchor head by Eligehausen et al. (1992) and Nilforoush et al. (2017b), respectively. This behavior gives rise to the postpeak brittleness of the large-headed anchors (Fig. 10). It should also be emphasized that the change in the failure mode occurred after reaching the peak load, so it does not seem to affect the anchorage capacity, although it does affect the postpeak anchorage behavior. The observed brittle postpeak anchorage behavior for anchors with the large head could have been prevented if the concrete member had been reinforced.

Fig. 12 shows the load-displacement curves of headed anchors in RC and PC members of various thicknesses, and the failure load predicted by the CC method [Eq. (1)].

During pullout loading of Specimen RC-440-M1, the loading rod detached from the coupling nut at a load of $\sim 249\mathrm{kN}$. Therefore, this test was aborted before capturing the actual peak load of the headed anchor. In a subsequent attempt, the loading rod was fastened to the anchor and loaded again. The load-displacement curve obtained in the second attempt is shown by a broken line in Fig. 12(b). As the figure shows, the anchor ultimate load during this attempt is lower than that of its companion specimen. This is attributable to the fact that the concrete surrounding the anchor had already cracked when the anchor was loaded for the second time.

Compared with the headed anchors in PC members, the anchorage ductility and failure load improve for the headed anchors in RC members. In addition, the anchor displacement at peak load and postpeak load increases when surface reinforcement is present. The same behavior was observed in an experimental study by Nilsson et al. (2011) on anchor bolts in RC members.

The ratio of ultimate load in RC members to that in PC members is shown in Fig. 13 for various member thicknesses. As the figure shows, the surface reinforcement has a more favorable influence on the tensile breakout resistance of headed anchors in thin members than in thick members. The same tendency was reported by Nilforoush et al. (2017a) for the simulated headed anchors at various embedment depths in RC members of various thicknesses.

Fig. 14 shows the crack pattern of the headed anchors in RC members of various thicknesses. The tested anchors, even in the thinnest members, failed primarily via concrete cone breakout. As shown in Fig. 14(a), the orthogonal surface reinforcement prevents the bending cracks observed in the thinnest PC members. This reinforcement enhances the global bending stiffness of the member, thereby preventing the occurrence of splitting/bending cracks. Nevertheless, concrete cone cracks develop even in the presence of surface reinforcement and dominate the failure of the tested anchors. Similar-shaped concrete cones formed in RC members of various thicknesses. However, the cone surface formed in RC members is shallower than that formed in PC members.