Why your rigging spreadsheet will not survive a design review

Governing utilisation is computed from these checks only.

Independent utilisations from the other methodology families, shown for comparison. Not included in the governing result.

Lifting load $F$ at angle $\theta$ (in-plane) is decomposed at the weld centroid into a normal component $N=F\cos \theta$, a transverse shear $V=F\sin \theta$, and an in-plane bending moment $M=V\cdot h$, where $h$ is the pin-to-weld centroid distance. Topology: parallel side welds (two line welds along the lug sides).

• Governing utilisation is 146.5% - at this level the result should be independently verified, including FEA where appropriate.

• Geometry is a single-plate lug / padeye without cheek plates. Cheek-plate configurations are out of scope for v1 and require detailed analysis.
• The pin-region checks (net section, shear-out, bearing, pin) take the full load resultant, which is conservative for the membrane load path. When an out-of-plane load direction is declared, the out-of-plane component is checked separately at the base section and weld group; without it, the load is assumed in the plane of the lug.
• Behaviour assumed linear-elastic, isotropic, homogeneous. No plastic redistribution, residual stresses, or fatigue effects are considered.
• Load line passes through the hole centre. Eccentricities between sling and lug centreline are not explicitly evaluated.
• Allowable stresses are supplied by the user. The app does NOT apply any code-specific allowable factor (ASME BTH-1 $0.45F_y$, Eurocode ${\gamma }_M$, DNV DAF, etc.) until the corresponding source clause has been supplied and validated.
• Pin-to-hole fit is assumed reasonable (pin diameter slightly less than hole diameter). Extreme clearances or wear are not modelled.
• Bearing stress is taken as the projected nominal value ${\sigma }_b=\frac{F}{d_{p}t}$; actual contact stress peaks are not resolved.
• Fillet welds are symmetric about the load line. The weld group is analysed with $N=F_{ip}\cos \theta$, $V=F_{ip}\sin \theta$, $M=V\cdot h$ (in-plane), plus - when an out-of-plane load direction is declared - $V_{oop}=F\sin {\theta }_{oop}$ and $M_{oop}=V_{oop}\cdot h$ resisted by the line method about the weld group's longitudinal axis. Root components combine vectorially before the $45°$ throat decomposition (${\sigma }_{\perp }={\tau }_{\perp }=f_t/\sqrt{2}$, ${\tau }_{\parallel }=f_v$). Asymmetric weld runs remain out of scope.

• MECH_NET_SECTIONMechanics of Materials - net-section tension
  Average tensile stress on the net cross-section through the pin hole: $\sigma =\frac{F}{(w-d)t}$. Classical identity; no code-specific allowable applied.
• MECH_DOUBLE_SHEAR_OUTMechanics of Materials - double-plane shear-out
  Average shear stress on two tear-out planes between hole and free edge: $\tau =\frac{F}{2L_{s}t}$, where $L_s=a-\frac{d}{2}$. Classical identity; no code-specific allowable applied.
• MECH_BEARINGMechanics of Materials - bearing stress
  Nominal bearing stress on the projected pin-on-plate area: ${\sigma }_b=\frac{F}{d_{p}t}$. Classical identity; no code-specific allowable applied.
• MECH_PIN_SHEARMechanics of Materials - pin double shear
  Average shear stress on two pin cross-sections (single-lug in clevis): $\tau =\frac{F}{2A_{\text{pin}}}$. Classical identity; no code-specific allowable applied.
• MECH_FILLET_WELD_THROATMechanics of Materials - fillet weld throat resultant
  Resultant throat stress on a fillet weld group with angle-aware demand decomposition $N=F\cos \theta$, $V=F\sin \theta$, $M=V\cdot h$. Root components $f_z=\frac{N}{A_w}+\frac{M}{S_w}$ and $f_v=\frac{V}{A_w}$ give throat components ${\sigma }_{\perp }={\tau }_{\perp }=f_z/\sqrt{2}$ and ${\tau }_{\parallel }=f_v$; $|R|=\sqrt{f_z^2+f_v^2}$ is compared to the user-supplied shear allowable. No ${\gamma }_M$ or electrode-specific factor applied.
• MECH_FILLET_WELD_VMMechanics of Materials - von Mises throat stress
  Combined throat stress for a fillet weld group using the von Mises equivalent ${\sigma }_{\mathrm{eq}}=\sqrt{{\sigma }_{\perp }^2+3({\tau }_{\perp }^2+{\tau }_{\parallel }^2)}$, compared to the user-supplied tensile allowable (falls back to ${\tau }_{\text{allow}}\sqrt{3}$).
• AISC_WELD_J24AISC 360-22 §J2.4 · 2022 · §J2.4, Eq. J2-5 (directional strength increase)
  Nominal fillet-weld strength per unit throat area $F_{nw}=0.60F_{EXX}(1.0+0.5{\sin }^{1.5}\theta )$, where $\theta$ is the angle between the line of action of the force resultant and the weld longitudinal axis. ASD safety factor $\Omega =2.0$ per §B3.2.
• MECH_BASE_SECTIONMechanics of Materials - cantilever base section
  The lug is a rectangular cantilever plate fixed at the base section, loaded at the pin a lever arm $h$ above it. Load decomposition $F_{ip}=F\cos {\theta }_{oop}$, $F_{oop}=F\sin {\theta }_{oop}$; base actions $N=F_{ip}\cos \theta$, $V_{ip}=F_{ip}\sin \theta$, $M_{ip}=V_{ip}h$, $V_{oop}=F_{oop}$, $M_{oop}=F_{oop}h$; section $A=wt$, $S_{maj}=t{w}^2/6$, $S_{min}=w{t}^2/6$. Classical identities; the mechanics check compares the combined von Mises stress against the user-supplied allowable.
• BTH1_NET_TENSIONASME BTH-1-2020 §3-3.3.1 · 2020 · §3-3.3.1 (eqs 3-45 through 3-48)
  Static strength of pin-connected plate - tension on the effective net area either side of the pin hole, with $C_r$ reduction for pin/hole clearance and $b_{\mathrm{eff}}=\min \left(4t,0.6b_e\sqrt{\frac{F_u{D}_h}{F_y{b}_e}}\right)$. Allowable includes design factor $N_d$ per §3-1.3.
• BTH1_DESIGN_FACTORASME BTH-1-2020 §3-1.3 · 2020 · §3-1.3
  Design factor $N_d$: $2.00$ for Design Category A, $3.00$ for Design Category B. Applied to all §3-3 allowables.
• BTH1_FRACTUREASME BTH-1-2020 §3-3.3.1 · 2020 · §3-3.3.1 (eq 3-49)
  Single-plane fracture strength beyond the pin hole: $P_b=C_r\frac{F_u}{1.20N_d}\left[1.13\right(R-\frac{D_h}{2})+\frac{0.92b_e}{1+b_e/D_h}]t$, with $R$ measured from the hole centre to the plate edge in the direction of the applied load.
• BTH1_SHEAR_OUTASME BTH-1-2020 §3-3.3.1 · 2020 · §3-3.3.1 (eqs 3-50 through 3-52)
  Double-plane shear-out strength: $A_v=2t[a+\frac{D_p}{2}(1-\cos \phi )]$, with $\phi =55^{\circ }\frac{D_p}{D_h}$ and allowable $P_v=\frac{0.70F_u{A}_v}{1.20N_d}$.
• BTH1_BEARINGASME BTH-1-2020 §3-3.3.4 · 2020 · §3-3.3.4 (eqs 3-53 / 3-54)
  Pin bearing strength on the lug plate. Static bearing allowable $\frac{1.25F_y}{N_d}$; rotating (Service Class $\ge 1$) reduced to $\frac{0.63F_y}{N_d}$.
• BTH1_WELDASME BTH-1-2020 §3-3.4.3 · 2020 · §3-3.4.3 (eq 3-55)
  Allowable fillet-weld shear on the effective throat $F_v=\frac{0.60F_{EXX}}{1.20N_d}$. Extended to combined in-plane loading by comparing the resultant throat stress $f_{\mathrm{res}}=\sqrt{f_z^2+f_v^2}$ against the clause allowable.
• BTH1_TENSION_BENDINGASME BTH-1-2020 §3-2.4(c) · 2020 · §3-2.4(c) (eq 3-35), with §3-2.1 (eq 3-1) for $F_t$
  Combined axial tension and biaxial bending: $f_t/F_t+f_{bx}/F_{bx}+f_{by}/F_{by}\le 1.0$, with $F_t=F_y/N_d$ on the gross section. Reduces to the §3-2.3.5 biaxial criterion (eq 3-26) when $f_t=0$.
• BTH1_RECT_MAJORASME BTH-1-2020 §3-2.3.3 · 2020 · §3-2.3.3 (eqs 3-19 through 3-24)
  Major-axis bending of solid rectangular bars, banded by $\lambda =L_{b}d/t^2$: $\lambda \le 0.08E/F_y$ → $F_b=1.25F_y/N_d$; $0.08E/F_y<\lambda \le 1.9E/F_y$ → $F_b=C_{LTB}{C}_b[1.52-0.274\lambda F_y/E]F_y/N_d\le 1.25F_y/N_d$; $\lambda >1.9E/F_y$ → $F_b=C_{LTB}1.9E{C}_b/(N_d\lambda )\le 1.25F_y/N_d$. $C_{LTB}=1.00$ braced, else $3.00\sqrt{E{I}_x/(GJ)}/(L_b/t)\le 1.00$. Page-verified 2026-08-12.
• BTH1_RECT_MINORASME BTH-1-2020 §3-2.3.4 · 2020 · §3-2.3.4 (eq 3-25)
  Minor-axis bending of solid rectangular sections: $F_b=1.25F_y/N_d$.
• BTH1_MATERIAL_CONSTANTSASME BTH-1-2020 §1-6.1 · 2020 · §1-6.1
  Material constants used by the standard's member equations: $E=200000$ MPa, $G=77200$ MPa.
• BTH1_CRITICAL_STRESSASME BTH-1-2020 §3-2.5 · 2020 · §3-2.5 (eq 3-37)
  Combined normal and shear stresses at a point: $f_{cr}=\sqrt{f_x^2-f_x{f}_y+f_y^2+3f_v^2}\le F_y/N_d$. At the lug base corner $f_y=0$ (free surface).
• BTH1_PLATE_SHEARASME BTH-1-2020 §3-2.3.6 · 2020 · §3-2.3.6 (eq 3-28)
  Average shear stress on bars, pins, and plates: $F_v=F_y/(N_d\sqrt{3})$, valid for $h/t\le 2.45\sqrt{E/F_y}$ where $h$ is the clear plate depth parallel to the applied shear. Beyond that limit the clause requires a buckling method with the §3-1.3 design factor - the check reports indeterminate rather than a number.
• EC3_PIN_GEOMETRYEN 1993-1-8:2005 §3.13.1 · 2005 · §3.13.1, Table 3.9 (Type A - given thickness)
  Geometric requirements for pin-connected plates, Type A (given thickness $t$): $a\ge \frac{F_{\mathrm{Ed}}{\gamma }_{M0}}{2t{f}_y}+\frac{2d_0}{3}$ and $c\ge \frac{F_{\mathrm{Ed}}{\gamma }_{M0}}{2t{f}_y}+\frac{d_0}{3}$, where $a$ and $c$ are measured from the edge of the hole to the plate end / side edge. The Type B (given geometry) alternative $t\ge 0.7\sqrt{F_{\mathrm{Ed}}{\gamma }_{M0}/f_y}$, $d_0\le 2.5t$ applies only to the specific lug shape drawn in Table 3.9 and is reported informatively.
• EC3_PIN_SHEAREN 1993-1-8:2005 §3.13.2 · 2005 · §3.13.2, Table 3.10 (shear)
  Pin shear resistance per plane: $F_{v,\mathrm{Rd}}=\frac{0.6A{f}_{up}}{{\gamma }_{M2}}$. A single-lug / clevis assembly presents two shear planes.
• EC3_PLATE_BEARINGEN 1993-1-8:2005 §3.13.2 · 2005 · §3.13.2, Table 3.10 (bearing)
  Pin/plate bearing resistance: $F_{b,\mathrm{Rd}}=\frac{1.5td{f}_y}{{\gamma }_{M0}}$.
• EC3_WELD_DIRECTIONALEN 1993-1-8:2005 §4.5.3.2 · 2005 · §4.5.3.2 (directional method)
  Directional check for a fillet weld throat. Two criteria: $\sqrt{{\sigma }_{\perp }^2+3({\tau }_{\perp }^2+{\tau }_{\parallel }^2)}\le \frac{f_u}{{\beta }_w{\gamma }_{M2}}$ and ${\sigma }_{\perp }\le \frac{0.9f_u}{{\gamma }_{M2}}$. Correlation factor ${\beta }_w$ taken from Table 4.1 based on the weaker joined steel grade.
• EC3_WELD_SIMPLIFIEDEN 1993-1-8:2005 §4.5.3.3 · 2005 · §4.5.3.3 (simplified method)
  Simplified check on the weld throat as a vector shear: $F_{w,Ed}\le F_{w,Rd}=f_{vw,d}a$ with $f_{vw,d}=\frac{f_u/\sqrt{3}}{{\beta }_w{\gamma }_{M2}}$. Conservative relative to the directional method; shown as a cross-check.
• EC3_BASE_YIELDEN 1993-1-1:2005 §6.2.1(5) · 2005 · §6.2.1(5), eq (6.1)
  Elastic yield criterion at a critical point; with ${\sigma }_z=0$: ${\left(\frac{{\sigma }_x{\gamma }_{M0}}{f_y}\right)}^2+3{\left(\frac{\tau {\gamma }_{M0}}{f_y}\right)}^2\le 1$. Applied at the lug base section extreme fibre.
• DNV_N001_DAFDNV-ST-N001 §16.2.5 · 2018, amended 2020-01 · §16.2.5, Table 16-1 (DAF in air, excluding elevated jackups)
  Dynamic amplification factor from Table 16-1 by environment column and static hook load. For $3<\mathrm{SHL}\le 100$ t: onshore $1.10$, inshore $1.07+0.05\sqrt{100/\mathrm{SHL}}$, offshore $1+0.25\sqrt{100/\mathrm{SHL}}$; banded constants above 100 t. Items lighter than 3 t are taken as 3 t (note 1). SHL includes rigging weight - the app approximates SHL with the design load; user-overridable.
• DNV_N001_SKEWDNV-ST-N001 §16.2.6 · 2018, amended 2020-01 · §16.2.6 (skew load factor)
  Skew-load factor for rigging tolerance / force-distribution effects. §16.2.6.9 permits $\mathrm{SKL}=1.00$ for statically determinate lifts; the default 1.10 is retained as a conservative baseline for a single-lug padeye. Multi-sling redistribution is out of scope for v1.
• DNV_N001_CONSEQUENCEDNV-ST-N001 §16.8.3 · 2018, amended 2020-01 · §16.8.3, Table 16-5 (consequence factors)
  Consequence factor ${\gamma }_c$ applied to lift points including their attachments to the structure: $1.30$ per Table 16-5, applied together with all relevant §16.2 factors per §16.8.4.1. Members directly supporting or framing into the lift points use $1.15$ (out of scope for this single-lug check set).