How to Design a Retaining Wall — A Complete Step-by-Step Guide (ACI 318-25 & IBC 2024)

Most structural members fail by breaking. A retaining wall usually fails by moving — it slides forward, or it tips, or one edge of its footing pushes into the ground and the whole wall rotates. Nothing cracks first to warn you. That is why the building code hands retaining walls their own acceptance criterion before any concrete design starts: a safety factor of 1.5 against overturning and against sliding, at unfactored loads (IBC 2024 §1807.2.3). Only after the wall has proved it will stay put do you size the reinforcement.

This guide walks that whole method on one wall — a 12 ft cantilever concrete retaining wall — and it does something most worked examples avoid: the first trial fails. Proportioned by the standard rules of thumb, it comes out at a sliding safety factor of 0.86 against the required 1.5, its bearing resultant falls outside the middle third of the footing, and the toe bars cannot develop. That failure is the most useful part of the page, because fixing it is where retaining wall design actually lives. You will see the three levers available, what each one buys, what each one costs, and why the cheapest fix here is not concrete at all — it is a line in the backfill specification.

By the end you will have: the finished wall, every check with its substituted formula and code citation, the reinforcement drawing, the drainage detail that makes the whole design legal, the construction mistakes that undo it on site, and the numbers for what happens if the same wall is built in Seismic Design Category D.

Every number below was produced by the calculation engine that runs StructSuite’s Retaining Wall module, and then recomputed independently by hand — the full side-by-side comparison is in Verification near the end, and both halves are locked by automated regression tests.

#The wall we are going to design

A hillside residence needs a 12 ft grade change held back along the uphill side of a new yard. Above the wall is landscaping with an allowance for light equipment and stored material — a uniform surcharge of 100 psf. The geotechnical report classifies the site soil as SM silty sand, gives an allowable bearing pressure of 3,000 psf, a coefficient of base friction μ = 0.35, and a lateral bearing (passive) value of 200 psf per foot of depth. There is no basement, no slab tying the top of the wall to anything: the wall is free to move and rotate at the top, which is what will let us use active earth pressure.

Geometry — 12 ft stem of uniform 12 in thickness, retained height 12 ft (backfill to the top of the wall), toe 3.0 ft, heel 4.5 ft, footing 18 in thick, giving B = 3.0 + 1.0 + 4.5 = 8.5 ft, 1.0 ft of soil cover over the toe, and a 12 × 12 in shear key under the stem.

Soil — imported GW well-graded clean gravel as structural backfill, γ = 120 pcf, uniform surcharge q = 100 psf (variable), coefficient of friction μ = 0.35, passive lateral bearing 200 psf/ft, allowable bearing 3,000 psf.

Concrete — f′_c = 3,000 psi, Grade 60 (f_y = 60,000 psi), normalweight 150 pcf. Covers per ACI 318-25 Table 20.5.1.3.1: 3 in at the footing bottom (cast against earth), 2 in at the stem earth face and the footing top (formed, exposed to earth). Reinforcement: stem verticals #6 @ 8 in at the earth face, stem horizontals #4 @ 9 in, heel top #6 @ 10 in, toe bottom #5 @ 9 in, footing longitudinal #5 @ 9 in.

Site seismicity — Seismic Design Category B, so no seismic earth pressure increment is required (IBC 2024 §1807.2.2 triggers it only in SDC D, E or F with more than 6 ft of backfill). What changes in SDC D re-runs this identical wall with the increment switched on, so you can see the cost of shaking on a wall you already understand.

The design assumes drained backfill. That is not a footnote — it is a load case. IBC 2024 §1610.1 otherwise requires full hydrostatic pressure, which on this wall would be catastrophic; the drainage section shows the number.

#Which wall type — and when a cantilever is the right one

Before any calculation, one decision sets everything else: what kind of wall. Four types cover almost all site work, and the choice is mostly about height and economics — how much concrete versus how much formwork and steel.

At 12 ft, a cantilever is the obvious answer, and it is the type this guide designs. The rule that decides it: a gravity wall needs a base roughly 0.5–0.7 × its height of solid concrete, so its cost grows with the square of the height, while a cantilever recruits the weight of the backfill sitting on its heel to do the same job for free. Above roughly 20–22 ft a cantilever stem gets thick enough (and its steel congested enough) that counterforts start winning again.

Practical note. For walls under about 4 ft of retained height with no surcharge, check whether your jurisdiction's prescriptive tables cover it — IRC 2024 §R404.4 applies the same 1.5 safety factor, and many building departments will accept a prescriptive detail without a stamped calculation. Above 4 ft (or with any surcharge), you are designing.

#Where earth pressure comes from — at-rest, active and passive

Everything downstream rides on one number, the design lateral soil load in pounds per square foot per foot of depth. It is worth 60 seconds to understand what it is, because the single most expensive mistake in retaining wall design is picking the wrong one of the three states below.

Soil is not a fluid, but it pushes like one, and the amount it pushes depends on how much the wall lets it move:

The classical (Rankine) coefficients behind the curve, for level backfill with a smooth vertical wall face:

You do not have to compute those to use the code. IBC 2024 Table 1610.1 tabulates the design lateral soil load directly by USCS backfill class, as an equivalent fluid pressure w in psf per foot of depth, and §1610.1 permits active values for walls free to move and rotate at the top. But the classical formulas are worth carrying, because they tell you what the table is quietly assuming — and they let you sanity-check a geotechnical report. Divide the table's value by the soil unit weight and you recover the coefficient the table implies:

K = w / γ = 30 / 120 = 0.250 → φ ≈ 37°, which is exactly what a well-compacted, well-graded gravel gives.

Run the same check on the site's silty sand: K = 45/120 = 0.375 → φ ≈ 27°. Also right for an SM. The code table is not arbitrary; it is Rankine with sensible friction angles baked in. When a geotechnical report hands you an equivalent fluid pressure that implies a φ your soil description cannot support, that is a question worth asking before you design to it.

#Choosing the input parameters — where each one comes from

Every soil number in a retaining wall design has a pedigree, and a plan checker will ask for it. Here is each one, its source, and what happens if you get it wrong.

• Design lateral soil load (30 psf/ft here) — IBC 2024 Table 1610.1 by USCS backfill class, unless a geotechnical investigation per §1803 determines otherwise. The table values are minimums for level backfill: a geotech report may require more (a sloped backfill takes the tables off the menu entirely, compaction-induced pressure) and occasionally allows less. ASCE 7-22 Table 3.2-1 is the parallel table and it disagrees with the IBC for several classes — clayey sand (SC) is 60 psf/ft in the IBC and 85 in ASCE 7-22. StructSuite offers both verbatim; use the one your jurisdiction enforces.
• Soil unit weight γ (120 pcf) — geotechnical report; 110–130 pcf is the usual band for compacted granular backfill. It does double duty: it sets the weight of soil on the heel (most of the resisting moment) and the implied K used for the surcharge, so it appears on both sides of the stability equation.
• Coefficient of friction μ (0.35) — IBC Table 1806.2 presumptive values run 0.25–0.70 by class, multiplied by the dead load per footnote a. Clays get cohesion instead of friction, capped at one-half the dead load (§1806.3.2). μ is the single most influential number in the sliding check — see the sensitivity table.
• Passive lateral bearing (200 psf/ft) — the "lateral bearing pressure" column of IBC Table 1806.2, or the geotech report. Count it only where the soil in front of the toe is permanent, compacted and undisturbed for the life of the wall. A future utility trench erases it. StructSuite has a toggle for exactly that judgment, and this wall is one where it matters.
• Allowable bearing pressure (3,000 psf) — geotech report, or Table 1806.2 presumptive used with ASD combinations (§1806.1). Note the presumptive class-3 value for gravel happens to be 3,000 psf too.
• Surcharge (100 psf, variable) — your own project loading: sidewalk and light-traffic allowances, stored material, an adjacent footing. Its nature matters more than its size. A variable surcharge (traffic, parking, storage) pushes but is not allowed to help hold the wall down; a permanent surcharge (an adjacent footing, permanent fill) is dead load and does both. IBC §1807.2.3 is what forces that asymmetry, and StructSuite asks the question rather than assuming.
• Seismic Design Category — from the project's seismic design (ASCE 7-22 Chapter 11). SDC D, E or F plus more than 6 ft of retained backfill makes the seismic earth pressure increment mandatory (IBC §1807.2.2). This site is SDC B, so it is not required; the last major section shows what would change if it were.

The parameter engineers most often get wrong is not a soil property — it is the condition. Active pressure is legal only if the wall can rotate at the top. Tie the top of that wall to a slab, a stair, a pilaster braced by a building, or a rigid corner return, and it becomes an at-rest wall. For this backfill that is 60 psf/ft instead of 30 — a doubling of the lateral load. Section What this wall is sensitive to shows what that alone does: the sliding safety factor falls from 1.62 to 0.81.

#Step 1 — Proportion the wall (the first pass)

You cannot check a wall you have not drawn, so retaining wall design starts with a guess — proportions from rules of thumb that have survived a century of practice — and then converges. Here is the vocabulary and the rules.

That is the first trial: B = 7.5 ft, toe 2.5 ft, heel 4.0 ft, 18 in footing, 12 in stem, no shear key, with the site's own SM silty sand re-used as backfill (45 psf/ft) because that is the cheapest thing to do with the spoil you already excavated. Now check it.

#Step 2 — Build the pressure diagram

Two loads push on the wall, and they have different shapes. The soil pressure grows linearly with depth (a triangle); the surcharge produces a uniform horizontal pressure over the full height (a rectangle), because a load spread over the whole ground surface adds the same vertical stress at every depth, and the soil converts a fraction K of it to horizontal stress.

For the stability checks the pressure acts on the vertical plane at the back of the heel, from the backfill surface all the way down to the bottom of the footing:

H_p = H_s + t_f = 12.00 + 1.50 = 13.50 ft

Resolving each block into its resultant, per foot of wall:

• Soil: P_a = w·H_p²/2 = 45 × 13.50²/2 = 4.101 kip/ft, acting at H_p/3 = 4.50 ft above the footing base (the centroid of a triangle).
• Surcharge: P_q = K·q·H_p = 0.375 × 100 × 13.50 = 0.506 kip/ft, acting at H_p/2 = 6.75 ft (the centroid of a rectangle).

The surcharge is only 11% of the total push but it acts 50% higher up, so it contributes 16% of the overturning moment. That leverage is why a surcharge you were tempted to ignore keeps showing up in the governing case.

Where the K for the surcharge comes from. The code table gives you an equivalent fluid pressure, not a coefficient — but the surcharge needs a coefficient. Recover it from the table itself: K = w/γ = 45/120 = 0.375. Surcharge-induced lateral pressure is part of the load H (ASCE 7-22 §2.2), added per IBC §1610.1, which matters later: it takes the same load factors as the earth pressure, not the factors of whatever is sitting on the ground.

#Step 3 — Overturning: will it tip?

IBC 2024 §1807.2.3 is explicit and easy to get wrong: the §1605 load combinations do not apply to stability. Stability is checked at 1.0 × nominal loads, with variable loads investigated set to zero. That cuts both ways, and it is the reason the surcharge appears on only one side of the equation:

• The surcharge's lateral push is driving, so it stays in.
• The surcharge's vertical weight over the heel would help resist — so it is dropped.

Resisting weights and moments about the toe (per foot of wall):

Look at row 4: the soil sitting on the heel supplies 72% of the resisting moment. The concrete is almost incidental. That is the whole idea of a cantilever wall, and it is why the heel length is the most powerful geometric variable you have.

Overturning moments about O:

• M_OT(soil) = P_a × H_p/3 = 4.101 × 4.50 = 18.453 kip-ft/ft
• M_OT(surcharge) = P_q × H_p/2 = 0.506 × 6.75 = 3.417 kip-ft/ft

Two cases have to be investigated, because §1807.2.3 says variable loads are examined set to zero:

• Case D + H (no surcharge): FS = 43.783 / 18.453 = 2.37
• Case D + H + surcharge (driving only): FS = 43.783 / 21.870 = 2.00 ← governs

FS_OT = 2.00 ≥ 1.5 ✓. Overturning passes comfortably, which is typical: a wall with a heel full of soil rarely tips. Do not let that reassure you.

#Step 4 — Sliding: the check that actually fails

Sliding is defined by §1807.2.3 as available soil resistance at the base ÷ net lateral force. Resistance may combine three things (IBC §1806.3.1 permits them to be added):

1. Base friction — μ × the dead load, per Table 1806.2 footnote a. Friction = 0.35 × 9.547 = 3.342 kip/ft.
2. Cohesion — c × contact area, capped at one-half the dead load (§1806.3.2). Zero for a granular soil.
3. Passive resistance in front — a triangular block over the depth D_p from the ground surface in front down to the lowest concrete. With no key, D_p = 1.00 (soil over the toe) + 1.50 (footing) = 2.50 ft, so P_p = ½ × 200 × 2.50² = 0.625 kip/ft.

R = 3.342 + 0 + 0.625 = 3.967 kip/ft

Against the driving side:

• Case D + H: ΣH = 4.101 → FS = 3.967 / 4.101 = 0.97
• Case D + H + surcharge: ΣH = 4.101 + 0.506 = 4.607 → FS = 3.967 / 4.607 = 0.86 ← governs

FS_sliding = 0.86 < 1.5 ✗ FAIL — and not narrowly. The wall needs its sliding resistance to nearly double. Notice it fails even with no surcharge at all (0.97): this wall would slide under its own backfill.

This is the normal outcome of a first pass, and it has a structural reason. The rules of thumb that set B were calibrated against overturning, which is a moment problem and therefore rewards width. Sliding is a force problem: widening the base adds friction only in proportion to the extra soil weight it picks up, while the driving force does not care about B at all. The two checks respond to geometry completely differently, which is why FS_OT = 2.00 and FS_sliding = 0.86 can coexist on the same wall.

#The four ways a retaining wall fails

Before fixing it, it is worth seeing the full set. Three of these are in the code checks above; the fourth is not in ACI or the IBC at all, and it is the one that shows up in photographs.

Mode 4 deserves a warning. A wall can pass overturning, sliding and bearing with room to spare and still ride a slip circle down the hill, taking the soil beneath its footing with it. Nothing in ACI 318 or IBC Chapter 18 checks that. If your wall sits at the crest of a slope, above a descending grade, on fill, or over soft or layered soils, global stability is a question for the geotechnical report — ask for it explicitly.

#Step 5 — Bearing pressure, and the middle third

The vertical load lands on the soil through the footing, and how it lands depends on where the resultant of all forces crosses the base. Take moments about O again, this time to locate the resultant:

x̄ = (M_r − M_OT) / ΣV, measured from the toe, and e = B/2 − x̄

If the resultant falls within the kern — the middle third, |e| ≤ B/6 — the whole base stays in compression and the pressure is a trapezoid:

q = (ΣV/B)(1 ± 6e/B)

If it falls outside, the heel lifts. Soil cannot pull, so the contact shortens to 3x̄ and the pressure becomes a triangle with a much higher peak:

q_max = 2ΣV/(3x̄), contact length = 3x̄

Both cases occur on this wall — one in each trial:

For trial 1:

q_max = 2,736 psf ≤ 3,000 psf ✓ (D/C = 0.91) — it passes on numbers, and it is still a problem. Note the asymmetry with the stability checks: for bearing, the surcharge's vertical weight is unfavourable (it raises the pressure), so it is included; on the resisting side of stability it was excluded. Same load, opposite treatment, because §1807.2.3 asks for the worst case each time.

Why "it passes" is not good enough here. Partial contact means the heel has lifted off. Three consequences follow: the pressure peak is far more sensitive to any extra load or any error in your weights; the wall's tilt under service load grows, because settlement concentrates under the toe; and the strength-level distributions get much worse — under 0.9D + 1.6H the trial wall bears on just 1.54 ft of its 7.5 ft base at a factored 11,155 psf. Most experienced designers treat "resultant inside the kern under service loads" as a design requirement even though the code does not spell it out, and StructSuite flags it: "Resultant falls outside the middle third — partial base contact. Consider widening the footing."

#Three levers: how to fix a wall that fails sliding

Trial 1 needs its sliding safety factor to go from 0.86 to 1.5 — a 74% increase in resistance, or an equivalent cut in the driving force. There are only three places that number can come from, and it is worth knowing what each is worth before you start guessing.

Lever 1 — geometry. More base width picks up more soil on the heel, which adds friction (μ × ΣW) and helps overturning enormously. It costs concrete and excavation.

Lever 2 — a shear key. A block cast below the footing that pushes the passive pressure block deeper. Passive resistance grows with the square of depth, so a 1 ft key on this wall takes D_p from 2.50 to 3.50 ft and nearly doubles P_p — for about 0.04 cubic yards of concrete per foot of wall. It is the cheapest structural fix there is.

Lever 3 — the backfill specification. Change what goes behind the wall. Swapping the site's SM silty sand for imported GW well-graded gravel takes the design lateral soil load from 45 to 30 psf/ft — a 33% cut in the driving force, applied to both the soil and the surcharge term. It costs the price difference on the fill, and you were going to place free-draining material behind the wall anyway.

Here is what each is actually worth on this wall — every row is a separate run of the engine, changing one thing at a time from trial 1:

Three lessons are hiding in that table, and they are the reason it is worth running the study rather than guessing:

A thicker footing made the wall worse. Going from 18 to 24 in dropped the overturning safety factor from 2.00 to 1.89 and pushed the bearing pressure past its allowable to 3,089 psf. The reason is that footing thickness enters both sides: it adds weight, but it also lengthens the pressure height H_p from 13.5 to 14.0 ft, and the driving force grows with H_p². The moment arm of the added weight is at mid-width, where it does relatively little. More concrete is not automatically more safety.

The passive term is doing a lot of work for very little concrete — which should make you cautious. Deepening the soil cover over the toe from 1 to 3 ft bought more sliding resistance than a full 2.5 ft of extra base width, and cost nothing at all. But it only exists if that soil is there, compacted, forever. See the traps.

The cheapest structural fix was written in the specification, not drawn on the section. Imported gravel behind the wall drops the driving force by a third for the price difference on the fill — and it is the same material that makes the drainage detail work. That is why the final design takes it.

#The final design

Three changes from trial 1, chosen from the study above:

Everything else — 12 ft stem at 12 in, 18 in footing, 1 ft of soil cover over the toe, f′_c = 3,000 psi, Grade 60, #6 @ 8 stem verticals — is unchanged. The rest of this guide checks that wall, and the numbers below are the finished design.

Re-running the stability checks with w = 30 psf/ft and K = w/γ = 30/120 = 0.250:

• P_a = 30 × 13.50²/2 = 2.734 kip/ft at 4.50 ft; P_q = 0.250 × 100 × 13.50 = 0.338 kip/ft at 6.75 ft; ΣH = 3.071 kip/ft

• Overturning: M_OT = 12.302 + 2.278 = 14.580 → FS = 55.993/14.580 = 3.84 ≥ 1.5 ✓ (4.55 without the surcharge)
• Sliding: friction = 0.35 × 10.703 = 3.746; D_p = 1.00 + 1.50 + 1.00 = 3.50 ft → P_p = ½ × 200 × 3.50² = 1.225; R = 4.971 kip/ft → FS = 4.971/3.071 = 1.62 ≥ 1.5 ✓ (1.82 without the surcharge)
• Bearing: ΣV = 11.152 kip/ft, x̄ = 3.966 ft, e = 0.284 ft ≤ B/6 = 1.417 ft → inside the kern, full contact, q_toe = 1,576 psf, q_heel = 1,049 psf. 1,576 ≤ 3,000 ✓ (D/C = 0.53)

The wall is stable. Now it needs reinforcement.

#Where the critical sections are — and why they differ

Before the strength checks, one detail that separates a correct retaining wall calculation from a plausible one. ACI 318-25 does not put the critical section in the same place for the stem, the heel and the toe, and the reason is mechanical, not arbitrary.

Getting this backwards is a common and consequential error: taking the heel shear at d here would remove d = 15.63 in = 1.30 ft of loaded length, dropping V_u from 9.711 to 2,158 × (4.50 − 1.30) = 6.90 kip/ft — a 29% understatement, on the member that is usually working hardest.

#Step 6 — Stem: a one-way slab per ACI 318-25 §13.3.6

ACI 318-25 §13.3.6.1 treats the stem of a cantilever retaining wall as a one-way slab designed per Chapter 7, and §13.3.6.3 puts the critical section for a uniform-thickness stem at the stem–footing interface. The pressure on the stem acts over the retained height H_s = 12 ft only — the footing depth belongs to the stability model, not to the stem.

Service moment at the interface:

M = w·H_s³/6 + K·q·H_s²/2 = 30 × 12.00³/6 + 0.250 × 100 × 12.00²/2 = 10.440 kip-ft/ft

Lateral earth pressure adds to the load effect, so H takes a load factor of 1.6 (ASCE 7-22 §2.3.1; ACI 318-25 §5.3.8):

M_u = 1.6 × 10.440 = 16.704 kip-ft/ft

Capacity with #6 @ 8 in at the earth face (A_s = 0.44 × 12/8 = 0.660 in²/ft; d = 12 − 2 − 0.375 = 9.63 in):

• a = A_sf_y/(0.85 f′_c b) = 0.660 × 60,000/(0.85 × 3,000 × 12) = 1.294 in
• φM_n = 0.90 × 0.660 × 60,000 × (9.63 − 1.294/2)/12,000 = 26.664 kip-ft/ft

D/C = 16.704 / 26.664 = 0.63 ✓

Three Chapter 7 requirements are not optional and are checked with it:

• Tension-controlled (§7.3.3.1; Table 21.2.2 — mandatory for one-way slabs, not just a way to get φ = 0.90): β₁ = 0.850, c = a/β₁ = 1.522 in, ε_t = 0.003(d − c)/c = 0.01597 ≥ ε_ty + 0.003 = 0.00207 + 0.003 = 0.00507 ✓
• Minimum steel (§7.6.1.1): A_s ≥ 0.0018bh = 0.0018 × 12 × 12 = 0.259 in²/ft ✓
• Maximum spacing (§7.7.2.3): s ≤ min(3h, 18 in) = 18 in ✓

Stem shear. §13.3.6.1.1 permits the simple form V_c = 2λ√f′_c·b_w·d for cantilever retaining wall stems — no size-effect factor, no axial term:

• V = w·H_s²/2 + K·q·H_s = 30 × 12.00²/2 + 0.250 × 100 × 12.00 = 2.460 kip/ft; V_u = 1.6 × 2.460 = 3.936 kip/ft
• φV_c = 0.75 × 2 × 1.0 × √3,000 × 12 × 9.63/1,000 = 0.75 × 12.652 = 9.489 kip/ft

D/C = 0.41 ✓ — checked at the interface, not at d above it.

Practical note on the stem. At 63% the stem has room, and that is deliberate: the stem is the cheapest part of a retaining wall to over-reinforce and the most expensive to repair. Note also what cannot be relaxed — the 12 in thickness is set by placement and cover, not by strength. Two inches of cover on the earth face plus 2 in on the formed face plus a #6 bar leaves very little in a 10 in stem, and the vibrator has to get down 12 ft.

#Step 7 — Heel: it hangs, and it carries the whole soil block

The heel is loaded downward by everything above it and hangs off the stem. Upward bearing pressure under the heel is conservatively neglected — a simplification that is on the safe side and is what makes the heel the usual governing member.

w_u = 1.2 × (γ·H_s + γ_c·t_f) + 1.6 × q = 1.2 × (120 × 12.00 + 150 × 1.50) + 1.6 × 100 = 2,158 psf

Both critical sections are at the back face of the stem:

• M_u = w_u·L²/2 = 2,158 × 4.50²/2 = 21.850 kip-ft/ft
• V_u = w_u·L = 2,158 × 4.50 = 9.711 kip/ft, taken at the face (§13.2.7.2 / §7.4.3.2)

Capacity with heel TOP bars #6 @ 10 in (A_s = 0.528 in²/ft; d = 18 − 2 − 0.375 = 15.63 in):

• a = 1.035 in → φM_n = 35.895 kip-ft/ft → D/C = 0.61
• φV_c = 0.75 × 2√3,000 × 12 × 15.63/1,000 = 15.405 kip/ft → D/C = 0.63 ← governs the heel
• ε_t = 0.03549 ≥ 0.00507 ✓ tension-controlled; A_s = 0.528 ≥ 0.0018 × 12 × 18 = 0.389 in²/ft ✓; s = 10 ≤ 18 in ✓

✓ PASS at D/C = 0.63. Note that the surcharge enters here at 1.6L — as a live load on top of the heel — while its lateral push entered the stability checks as H. Same 100 psf, three different treatments across this page, each one dictated by which effect it is producing.

#Step 8 — Toe: designed from the factored bearing pressure

The toe is loaded upward by the bearing pressure, less the weight of the toe itself and its soil cover. The pressure distribution has to be rebuilt at strength level for each combination, because the load factors change both the vertical load and the overturning moment, and therefore move the resultant:

The 0.9D + 1.6H combination is the one people forget. It strips the wall to 90% of its dead load while keeping the full 1.6 on the earth pressure, and it is often what pushes the resultant outside the kern at strength level even when the service case is comfortably inside — exactly what happens here.

Net of the factored toe self-weight and soil cover (q_u − 1.2 × (225 + 120) psf), with moment at the stem front face and shear at d from the face:

• Governing (1.2D + 1.6H + 1.6L): M_u = 7.969 kip-ft/ft, V_u = 3.181 kip/ft
• Toe BOTTOM bars #5 @ 9 in (A_s = 0.413 in²/ft, d = 18 − 3 − 0.313 = 14.69 in): a = 0.810 in → φM_n = 26.565 kip-ft/ft (D/C = 0.30), φV_c = 14.480 kip/ft (D/C = 0.22)

D/C = 0.30 ✓. The toe is nowhere near critical — but it cannot borrow the heel's bars. Its tension face is the bottom; the heel's is the top. Two separate mats, and mixing them up is one of the few retaining wall detailing errors that is unambiguously a failure.

Notice also that the toe bars are not sized by strength at all. A_s = 0.413 in²/ft is set by the shrinkage-and-temperature minimum of 0.389 in²/ft, and the bar size was set by development length (Step 10). That is normal for footings on this scale.

#Step 9 — The shear key: what it resists, it must also carry

Step 4 leaned on the key for 0.600 kip/ft of the 1.225 kip/ft of passive resistance. That force does not come from nowhere: the soil in front bears on the key's front face and the key has to carry it back into the footing. It is a short cantilever hanging off the base, and it gets designed like one.

The passive pressure over the key depth is the trapezoid between the footing bottom and the key bottom (200 psf/ft lateral bearing, IBC Table 1806.2):

• p_top = 200 × 2.50 = 500 psf; p_bot = 200 × 3.50 = 700 psf
• P_p,key = ½(500 + 700) × 1.00 = 600 lb/ft, acting ȳ = (1.00/3)(500 + 2 × 700)/(500 + 700) = 0.528 ft below the junction

Passive bearing is lateral earth pressure, so it enters at 1.6H — the same factor the active side carries on the stem:

• M_u = 1.6 × 0.600 × 0.528 = 0.507 kip-ft/ft; V_u = 1.6 × 0.600 = 0.960 kip/ft

The section is 12 in of wall length by the key width, with the stem #6 @ 8 pattern continued into the key and the bars toward the toe face, where the junction moment puts the section in tension. Cast against earth on both faces, so d = 12 − 3 − 0.375 = 8.63 in:

• φM_n = 23.694 kip-ft/ft → D/C = 0.02; φV_c = 0.75 × 2√3,000 × 12 × 8.63/1,000 = 8.503 kip/ft → D/C = 0.11 ✓

Three details decide the drawing rather than the numbers:

Shear stays at the face. The footing reaction does not put the key's end region in compression, so §7.4.3.2 gives no d relief — the same rule as the heel, the opposite answer from the toe.

The dowels need a hook. Straight development of a #6 is ℓ_d = 32.9 in and an 18 in footing offers only 18 − 2 = 16.0 in. A standard hook does fit: ℓ_dh = 11.5 in ≤ 16.0 in ✓. Detail the key dowels hooked into the footing; a straight bar would not be developed whatever the D/C says.

The key gets its own shrinkage and temperature steel. Its flexural reinforcement runs one way only, so §24.4.3.2 requires steel perpendicular to it — bars running along the wall inside the key. Sized from the key's own gross section, A_s ≥ 0.0018 × 12 × 12 = 0.259 in², and placed where S+T steel does its work, at the faces: (4) #5 continuous — 2 levels down the depth × a bar at each face (1.240 in²), at the footing longitudinal bar size. Sizing from the key’s own section is what makes a deep key come out right. A 12 × 36 in key needs A_s ≥ 0.0018 × 12 × 36 = 0.778 in², and the §24.4.3.3 spacing cap of min(5w, 18 in) = 18 in forces three levels down its depth rather than two. Width matters the same way: the cap applies across the key as well as down it, so a key wider than 24 in picks up a third bar across each level.

A plain-concrete key is not what this is. ACI 318-25 §14.5.2.1 allows only 5λ√f′_c on the gross section at φ = 0.60, which an ordinary key does not satisfy on its own moment. If your detail shows an unreinforced key, it is decoration.

#Step 10 — Development, hooks and laps: where the toe length came from

Three development questions decide whether the bars actually work — and one of them is why the final design has a 3.0 ft toe instead of 2.5 ft.

Stem dowels into the footing (§25.4.3.1, Table 25.4.3.2). With ψ_e = 1.0 (uncoated), ψ_s = 1.00 (≤ No. 9), ψ_cc = 0.7 (side cover normal to the hook plane ≥ 2.5 in — automatic in a continuous wall — and ≥ 2 in over the 90° hook) and ψ_r = 1.0:

ℓ_dh = (60,000 × 1.0 × 1.00 × 0.7 × 1.0)/(50 × 1.00 × √3,000) × 0.750 = 11.5 in ≥ max(8d_b = 6.0, 6 in)

Available = t_f − cover_bottom − d_b(toe mat) = 18.0 − 3.00 − 0.625 = 14.4 in ✓. The 90° hook extension is 12d_b = 9.0 in, placed near the bottom of the footing with the free end toward the toe (R13.3.6.3).

Dowel-to-stem lap (§25.5.2.1, Class B): ℓ_d(#6 vertical, ψ_t = 1.0) = 32.9 in → lap = 1.3ℓ_d = 42.7 in of dowel projection above the interface. Round that up on the drawing — 3 ft 8 in — and remember the dowel has to be that long before the stem is formed.

Heel and toe mats (§13.2.8.3; Table 25.4.2.3). The heel TOP bars sit above 15.3 in of fresh concrete — more than 12 in — so they take ψ_t = 1.3 (Table 25.4.2.5):

• Heel: ℓ_d = 42.7 in vs 4.50 × 12 − 3 = 51.0 in available ✓
• Toe: ℓ_d(#5) = 27.4 in vs 3.00 × 12 − 3 = 33.0 in available ✓

This is where the geometry came from. Trial 1's 2.5 ft toe with #6 bars needed ℓ_d = 32.9 in and offered only 27.0 in — a straight-bar development failure that no amount of extra area fixes. Two things solved it: a smaller bar at a tighter spacing (#5 @ 9 instead of #6 @ 12, same area class, ℓ_d down to 27.4 in because ℓ_d is proportional to d_b), and a longer toe (3.0 ft), which the sliding check wanted anyway. When a footing projection is too short to develop its bars, your options are exactly those two, or a hook at the end.

#Step 11 — Minimum and interface reinforcement

The bars nobody calculates and every plan checker looks for:

• Wall–footing interface (§16.3.4.2 → Table 11.6.1): A_s,min = 0.0015 × 12 × 12 = 0.216 in²/ft; the #6 @ 8 dowels provide 0.660 ✓
• Stem horizontal shrinkage + temperature (§24.4.3.2): 0.0018 × 12 × 12 = 0.259 in²/ft required; #4 @ 9 provides 0.267 in²/ft ✓, spacing 9 ≤ min(5h, 18) = 18 in (§24.4.3.3) ✓. A single curtain is expressly permitted in cantilever retaining wall stems regardless of thickness (§11.7.2.3 exception) — which is why the horizontals sit inside the verticals on the earth face and there is no second mat on the front face.
• Footing longitudinal S+T (§24.4.3.2) — the bars running along the wall: 0.0018 × 12 × 18 = 0.389 in²/ft required; #5 @ 9 provides 0.413 in²/ft ✓, spacing ≤ 18 in ✓. These are a real selection with real limits, not a note to the detailer.

#The reinforcement drawing

Every number above resolves into one drawing. This is the detail a contractor builds from:

StructSuite draws this detail live from the same numbers the checks used, so the hook length, the lap and the embedments on the drawing can never disagree with the calculation — open the Retaining Wall module and it is Step 5.

#Practical detailing — the details engineers actually draw

The drawing above is a correct detail, not the correct detail. On real sheets the same wall gets detailed several different ways, and the differences are about buildability and risk, not about strength. Here is the menu, with what each one costs.

#Getting the stem bars into the footing

For this wall detail A is the only one of the three that works: a straight #6 needs ℓ_d = 32.9 in of embedment and an 18 in footing offers 14.4 in, so the dowel has to be hooked (§25.4.3.1) — which is why the hook shows up in the calculation and not just on the drawing. Detail B is real, but only on short walls: a 12 ft bar cage standing unbraced while the footing is poured is a safety and tolerance problem, and it forces the stem forms over a full-height cage. Detail C stays a constructability choice: ACI permits every bar to be lap-spliced at one section for a Class B splice, so staggering buys you room for the vibrator, not a shorter lap. (ACI does allow a shorter Class A splice under Table 25.5.2.1, but only when the section has at least twice the steel it needs and no more than half the bars are spliced there — StructSuite never assumes either, so it reports the Class B length.)

Three things a detailer gets wrong here, in order of frequency. The lap drawn as two bars meeting rather than overlapping — see the caption above. The hook turned the wrong way: R13.3.6.3 puts it near the bottom of the footing with the free end toward the toe, so the bend works against the direction the stem wants to rotate. And the dowel projection dimensioned to the bar length instead of the lap length: what the drawing must state is 42.7 in above the interface, not the overall dowel.

#Where the footing mats start and stop

This is the question the section drawing earlier in this guide answers with detail B: the heel top mat runs through to the stem front face, and the toe bottom mat continues 27.4 in past the stem back face — each one carried its own development length past its own critical section, exactly as §13.2.8.3 and Table 13.2.7.1 require. That is the least-steel answer and it is what the module reports.

Most production drawings use detail A instead, and it is a defensible choice: on an 8.5 ft footing the extra steel is a few pounds per foot of wall, while a bar mat cut short in the field is a repair. Detail C is not an alternative — it is what you are forced into when a projection is shorter than the bar can develop straight, and it is the reason this wall’s toe grew from 2.5 ft to 3.0 ft rather than getting hooks.

Bars that are not in the calculation but must be on the drawing. Three of them, every time:

• Stem horizontal S+T on both faces of a two-curtain stem. This wall uses a single curtain, which §11.7.2.3 expressly permits in cantilever retaining wall stems regardless of thickness — but as soon as the stem gets a second curtain (usually past about 16 in, or where the designer wants crack control on the exposed face), the horizontals double and the outer face needs its own vertical support bars.
• Corner and end bars. Horizontal bars have to be continuous around a corner or the corner opens. Standard practice is a separate corner bar lapped to the horizontals in each leg, and at a free end the horizontals get a U-bar or a hook so the last vertical is tied into something. ACI gives you the lap length; the geometry is a drafting convention.
• Joints. Long walls need vertical joints or they will crack at intervals of their own choosing. Control joints at roughly 20–30 ft with the horizontal steel continuous through them and a formed groove on the exposed face; full construction or expansion joints at wider spacing and at every abrupt change in wall height, with a shear key or smooth dowels across so the two panels deflect together. Put a joint at every step in the footing.

One thing to leave off. Curtailing the stem verticals — cutting alternate bars at mid-height where the moment has dropped — is legitimate and saves steel on tall walls, but StructSuite designs the stem for a single uninterrupted bar pattern and reports the interface demand only. If you curtail, the cut-off point and the extension past it are yours to check; do not read this guide's D/C of 0.63 as covering a curtailed detail.

#Drainage — the detail that decides whether any of this is true

Every number on this page rests on one assumption: the backfill is drained. IBC 2024 §1610.1 requires design for full hydrostatic pressure where it is not, and Table 1610.1 footnote a is explicit — submerged or saturated soil pressures shall include the weight of the buoyant soil plus the hydrostatic loads. Drainage installed per §1805.4.2 and §1805.4.3 is what buys the exemption.

Here is what that is worth. Take the same wall and saturate the gravel: the soil's contribution drops to the buoyant unit weight, but full water pressure adds on top of it — K_a·γ′ + γ_w ≈ 0.25 × (135 − 62.4) + 62.4 ≈ 80 psf per foot of depth, against 30 psf/ft drained. Re-running the finished wall at an 80 psf/ft equivalent fluid pressure:

The drain is a structural element. Detail it and specify it like one:

• Free-draining material behind the wall, at least 12 in wide, extending from the footing up to near the surface. Where the whole backfill is imported gravel — as here — the drainage zone and the structural backfill are the same material, which is a large part of why that specification was the cheapest fix in the lever study.
• Filter fabric between the gravel and the native soil on the back and top faces. This is the component most often value-engineered out and it is the one that decides whether the drain still works in year five: without it, fines migrate into the gravel and the drain silts up.
• A perforated collector pipe, 4 in minimum, holes down, bedded in the gravel at the base, sloped continuously to a real outfall — daylight, a storm structure, or a sump. A pipe that ends in the backfill is worse than no pipe, because it collects water and stores it.
• Weep holes through the stem, typically 3 in at 5–10 ft on centre with a gravel pocket and fabric behind each one. Treat them as the backup and as the visible evidence that the system works, not as the primary drain — a weep hole that never runs is not proof of a dry wall, it may be proof of a blocked one.
• A low-permeability cap — pavement, a clay layer, or simply grading that falls away from the wall — so that surface water is shed rather than fed into the drain. Most walls that get into trouble are drowned from the top, not from below.
• The active wedge matters. Imported gravel only changes the pressure if it extends past the failure surface, a plane rising at roughly 45° + φ/2 from horizontal at the back of the heel — 63° for this backfill. A 12 in ribbon of gravel against a silty-sand backfill is a drain, not a 30 psf/ft design lateral soil load. If your calculation uses the gravel's number, your section must show the gravel filling the wedge.

#Building it: what actually goes wrong on site

The calculation is the easy half. Here is what to put on the drawings and watch for in the field.

Set the dowels in the footing pour, to a template. Dowels pushed into wet concrete after the screed drift out of position and lose cover, and the hook rarely ends up at the bottom where R13.3.6.3 wants it. Straightening or re-bending a partially embedded bar in the field is not a detailer's decision — it needs the engineer of record. Get the template right and the stem cage drops over it.

Roughen the construction joint at the top of the footing. The interface between footing and stem is a cold joint carrying the full stem shear and the dowel force. Clean it of laitance, roughen it, and wet it to saturated-surface-dry before the stem pour. The code counts on the dowels crossing that joint (§16.3.4.2); it does not count on a smooth, dusty surface.

Do not backfill early, and do not backfill heavy. Two rules that save more walls than any calculation. The stem must reach its specified strength first — a green cantilever has no capacity at its base. And compact in thin lifts with light equipment within about half the wall height of the back face: a heavy roller working right behind a wall induces compaction pressures well above the active values this design used, and it is a classic cause of a wall that leans before it is even in service.

Protect the toe soil, on the drawing and on site. This design counts on 1.0 ft of compacted soil in front of the toe plus the key for 25% of its sliding resistance. Note it on the drawings — "soil in front of the toe to remain undisturbed; do not excavate for utilities within X ft" — because a trench dug there in year three quietly takes the wall from FS 1.62 to 1.22.

Place the two mats the right way up. Heel bars on top, toe bars on bottom. Use chairs, check them before the pour, and photograph them. Inverting the heel mat is a genuine collapse mechanism, not a serviceability issue.

Verify cover, especially at the earth face. The stem's d = 9.63 in assumes 2 in of cover. Lose an inch to sloppy chairs and φM_n drops by about 10% — more than the whole margin some walls have.

Detail the joints. Long walls need vertical joints for shrinkage and temperature; without them the wall will pick its own locations. Control joints at roughly 20–30 ft with the horizontal steel continuous through, and full construction/expansion joints at greater spacing or at abrupt changes in wall height, with a shear key or dowels across so the two panels deflect together. Change wall height in steps, not gradually, and put a joint at each step.

Cure the stem. A 12 ft stem is mostly formed surface with a high surface-to-volume ratio; it cracks if it is stripped early and left to dry. Cracks on the earth face are also the path water uses to reach the reinforcement.

#Expansive soils — what the code lets you assume

Everything above assumed a backfill that behaves the way Rankine says it does. An expansive clay does not. It develops swelling pressure when it takes on water — a pressure that has nothing to do with K_a, K_o or K_p, that can exceed the at-rest value several times over, and that cycles with the seasons rather than settling to a design value. Here is what the code gives you, what it refuses to give you, and what the numbers look like when you try.

1. The code's first answer is: do not backfill with it. IBC 2024 Table 1610.1 tabulates a design lateral soil load for every USCS class it accepts. For CH (inorganic clays of high plasticity), MH (inorganic clayey silts, elastic silts), OL and OH it tabulates nothing — footnote b reads, verbatim, "Unsuitable as backfill material." There is no active value, no at-rest value, and therefore no legal way to design a wall for those materials out of the table. StructSuite does not paper over it: pick CH and the engine stops with "Design lateral soil load is not resolved… No default pressure is assumed."

2. The clays the table does accept still cost you double. CL (inorganic clays of low to medium plasticity) is tabulated at 60 psf/ft active and 100 psf/ft at-rest — twice and more than three times the imported gravel used here. ML-CL is the same. And both come with the table's blanket footnote a: the values are for moist conditions at optimum density; actual field conditions govern.

3. Cohesive soils have no coefficient of friction. IBC 2024 Table 1806.2 class 5 — "Clay, sandy clay, silty clay, clayey silt, silt and sandy silt (CL, ML, MH and CH)" — gives no friction coefficient at all. It gives cohesion c = 130 psf instead (footnote b: multiplied by the contact area, as limited by §1806.3.2), with vertical bearing 1,500 psf and lateral bearing 100 psf per foot of depth. In StructSuite you enter μ = 0 and c = 130 psf, and the engine takes c × B capped at one-half the dead load (§1806.3.2).

Put the same finished wall on that soil, keeping the imported gravel backfill, and the arithmetic is brutal:

The wall loses two-thirds of its sliding resistance and half its bearing capacity at once, and it fails both. Swap the gravel for a CL backfill on top of that and FS_sliding falls to 0.28 with FS_OT at 1.92; take the at-rest condition the way many geotechnical reports require for cohesive backfill and it is 0.17, with a bearing demand of 8,125 psf against a 1,500 psf allowable. There is no reinforcement detail that rescues any of that. On expansive ground the design decision is made in the specification and the drainage, not in the concrete.

4. Give up the passive term first. Passive resistance in front of a toe founded in expansive clay is the least reliable number on the page: shrinkage cracks open a gap at the face in the dry season, exactly when the soil is stiffest and the calculation looks best. Re-run the wall with the passive term neglected and it drops from 0.56 to 0.36. If the toe soil is expansive, neglect it.

5. Swelling pressure is not in the code, and it is not in StructSuite. No table in IBC Chapter 16 or 18 gives a swell pressure — it comes from the geotechnical investigation required under IBC §1803, as a measured value from the specific material at the specific moisture range. If your report states one, enter it as an equivalent fluid pressure in Step 2; that is the only honest way to get it into the model, and the module will tell you it is using your number rather than a table.

6. The seismic screening methods refuse cohesive backfill outright. Ask for the Seed-Whitman increment on a CL backfill and the engine stops with the scope statement it is quoting: "If soil conditions behind the wall have a cohesive soil component (i.e., a c-φ soil), this simplified approach is no longer appropriate." (FEMA P-750 Part 3, Resource Paper RP 12, p. 356 — the document ASCE 7-22 C11.8 defers to.) For an expansive site in SDC D–F, the seismic earth pressure is the geotechnical engineer's number, full stop.

What to actually do. In order: replace the expansive material inside the active wedge with non-expansive granular fill — the same lever the lever study used, now for a different reason, and the wedge geometry in the drainage detail is what tells you how far back the replacement has to go. Control the moisture: positive surface grading away from the wall, an impermeable cap, no planting beds or irrigation against the stem, and a drain that actually outfalls, because expansive clay damage is a moisture-change problem before it is a pressure problem. Take the report's numbers, not the table's, for the material you cannot replace. And design the base for cohesion, not friction — which on almost every expansive site means a wider footing, a deeper key, or both.

#Sloped backfill — when the ground keeps rising

Everything so far assumed the ground behind the wall is level. Regrade that yard so it keeps climbing — a 3H:1V lawn, a cut at the base of a hillside — and three things change at once, none of them in your favor. StructSuite models all three; this section shows the mechanics on the same finished wall, and what the regrade actually costs.

What the slope does to the statics. The design section for stability is the vertical plane at the heel edge. On level ground that plane runs from the backfill top to the footing bottom: H = 12 + 1.5 = 13.5 ft. On a slope the surface has been climbing since it left the stem, so by the heel edge it stands L_heel·tanβ higher, and the pressure plane is taller:

H′ = H_retained + t_f + L_heel·tanβ = 12 + 1.5 + 4.5 × (1/3) = 15.0 ft

Driving force grows as the square of that height — P_a = ½·w·H′² — so an 11% taller plane is a 23% larger thrust before the pressure itself changes at all. Second, the soil above the heel is no longer a rectangle: the slope adds a triangular wedge — ½·γ·L_heel²·tanβ, here ½ × 120 × 4.5 × 1.5 = 405 lb/ft at ⅔ of the heel from the stem — which resists overturning with the other weights but also loads the heel as a triangular dead load. Third, the pressure coefficient itself rises: Rankine's active coefficient for a sloped surface is larger than the level-ground value the code tables tabulate, and the thrust arrives inclined, parallel to the slope. StructSuite takes the horizontal component and conservatively neglects the slope-parallel vertical component on the resisting side.

Where the pressure number comes from — and why the module refuses the table. IBC 2024 Table 1610.1 and ASCE 7-22 Table 3.2-1 are level-backfill values: their basis is the classical active/at-rest states (Terzaghi and Peck — ASCE 7-22 C3.2.1), and §1610.1 offers them as minimum design loads "unless determined otherwise by a geotechnical investigation in accordance with Section 1803." A rising backslope increases the active coefficient beyond the tabulated state, so the minimum no longer bounds the real load — which is exactly the situation both codes route to the geotechnical investigation. Declare a slope in Step 1 and StructSuite guards the two tables off: the design pressure must be entered as the report's equivalent fluid pressure for the sloped condition. Enter Step 1's β and the geotech EFP in Step 2, and every check downstream — both stability cases, the kern-aware bearing, the heel's wedge terms, the seismic increment height — picks up the sloped geometry automatically. (Sanity-check the report the same way as before: Rankine's sloped-backfill coefficient for a φ = 36–38° gravel on a 3H:1V slope implies roughly 34–37 psf/ft of horizontal pressure against 30 level — if the report says 30 for a sloped yard, ask.)

The same wall, regraded. Take the finished 12-ft wall — unchanged concrete, unchanged bars — and let the grade above it climb at 3H:1V (β = 18.4°), with the geotechnical report giving 36 psf/ft for the sloped condition:

The wall that carried its level yard with a 1.62 sliding margin fails at 1.14 the day the grade changes. And the failure is not mostly the bigger pressure number: run the thought experiment of keeping the old 30 psf/ft on the sloped geometry and FS_sliding is still only 1.36 — the taller plane and the H′² term sink it before the coefficient moves. A regrade behind an existing retaining wall is a structural event, whatever the soils report says.

Fixing it has a trap in it. The two familiar levers interact with the slope in a way level-ground intuition misses: lengthening the heel raises H′, because the slope keeps climbing over every foot you add. Stretch the heel to 6.0 ft with a 24-in key — a combination that rescues far worse level walls — and sliding still fails at 1.45: the rise grew from 1.5 to 2.0 ft and the thrust grew with its square. Push the key to 30 in with the 6-ft heel and sliding clears at 1.56 — but now the heel bars fail: the longer cantilever picks up the wedge's triangle on top of the rectangle, and #6 @ 10 cannot carry it. The combination that closes everything is more surgical: heel 5.5 ft + key 30 in — FS_sliding = 1.52, FS_OT = 3.04, bearing 2,135 ≤ 3,000 psf, and the heel just inside its steel at D/C = 0.98, now flexure-governed where the level wall was shear-governed. On a slope, the heel's own strength — not stability — becomes the binding constraint, and one more foot of heel buys less resistance than it adds demand.

If the slope is seismic country too. The backfill slope in Step 1 is the one β the module knows: Mononobe-Okabe reads it into K_AE (mind the validity limit φ − ψ > β — a steep slope plus a strong motion has no equilibrium wedge, and the engine stops rather than extrapolating), and Seed-Whitman refuses a slope outright, quoting its source: RP 12's simplified formulation "is not applicable for sloping ground above the wall" (FEMA P-750 Part 3, p. 356). On a sloped site in SDC D–F your choices are full M-O within its limits, or the geotechnical report's seismic pressure — which the code makes the normative source anyway.

What the module does not do. The backslope model is an infinite uniform slope; β at or beyond 45° is refused. A broken slope (rises, then benches level) is conservative to model as infinite; a slope below the wall at the toe, or global stability of the whole hillside — the failure circle that passes under the footing and through the slope — is a geotechnical evaluation under IBC §1807.2.1/§1803, outside any wall-statics module. If the site is a real hillside, that check is not optional, and no cantilever wall design — hand or software — substitutes for it.

#What changes in Seismic Design Category D, E or F

The wall above is complete for an SDC A–C site. Where the structure is assigned to SDC D, E or F and the wall retains more than 6 ft, IBC 2024 §1807.2.2 requires the additional seismic lateral earth pressure, determined per ASCE 7-22 §11.8.3 — and there the geotechnical investigation is the normative source. Everything below is what StructSuite does when the report does not state a pressure and you need a defensible screening value.

Two things anchor the mechanics, both from ASCE 7-22 C11.8.3. The dynamic increment is an earthquake load E superimposed on the static pressure, which remains an H load — so the two carry different factors in every combination below, and mixing them up changes every number. And absent a site-specific study, the design-earthquake PGA defaults to S_DS/2.5. A cantilever concrete retaining wall is a yielding wall designed as a simple flexural element (§15.6.1), which is what makes that default applicable.

#Where the equations actually come from

This is worth knowing before you put a number on a drawing, because ASCE 7-22 prints no pressure equation at all. C11.8 defers to the 2009 NEHRP Recommended Seismic Provisions (FEMA P-750) Part 3, whose Resource Paper RP 12 carries the method. So:

• The Seed-Whitman simplified increment is printed verbatim in RP 12 as Eqs. 6–7 (p. 356) — not as code text. RP 12 restricts it on that same page to dry, cohesionless backfill (a c-φ soil needs the full methods) and to level ground above the wall.
• The 0.6·H application height and the "inverted trapezoidal pressure distribution" are RP 12’s (p. 356) as well.
• S_DS/2.5 is ASCE 7-22 C11.8.3 only; it does not appear in RP 12.
• The optional wall self-weight inertia term is the Richards & Elms (1979) displacement-based formulation, which RP 12 names (p. 356) without printing its equations.

StructSuite cites each piece to its actual source rather than passing all of it off as code, and it refuses to run the screening methods outside their scope — see Expansive soils for the cohesive-backfill refusal.

#The increment on this wall

Take the same finished wall, unchanged, and put it on an SDC D site with S_DS = 0.50 g, using Seed-Whitman as a yielding wall:

k_h = S_DS/2.5 = 0.50/2.5 = 0.20 g

ΔK_AE ≈ 0.75·k_h = 0.15, so ΔP_AE = ½·ΔK_AE·γ·H_p² = (3/8)·γ·H_p²·k_h = 0.5 × 0.15 × 120 × 13.50² = 1,640 lb/ft, applied at y_E = 0.6·H_p = 8.10 ft above the footing bottom.

That application height is the whole story of why a modest increment hurts: the static resultant sits at H_p/3 = 4.50 ft, the seismic one at 8.10 ft — 80% higher up, on a wall whose failure modes are all moments.

Where the 1.6 / 0.4 ordinates come from. RP 12 prints the shape and the 0.6·H resultant but not the ordinates. For a linear distribution, a centroid at 0.6·H forces p_top = 4·p_bot, and the total area then fixes p_top = 1.6·ΔP_AE/H and p_bot = 0.4·ΔP_AE/H. That is the unique linear distribution satisfying both printed facts, and it is what the module draws.

Two scope notes the module prints with the derivation. k_v = 0 inside the coefficient: vertical shaking enters the strength combinations separately as E_v = 0.2·S_DS·D (§12.4.2.2), so a k_v in K_AE would count it twice. And wall self-weight inertia (k_h × the concrete weights) is an opt-in, left off here — it is small next to a 1,640 lb/ft soil thrust, and peak soil thrust and peak wall inertia are generally not coincident.

#The six checks the increment adds

Three rules decide every one of those numbers, and each of them is a place designs go wrong:

• Stability at 0.7 × nominal earthquake, FS ≥ 1.1. IBC §1807.2.3’s earthquake exception is an additional check, not a replacement: the static FS ≥ 1.5 checks still stand on their own. The static earth pressure and surcharge stay at 1.0 throughout, because they are H, not E. And the resisting weights are not reduced by −E_v — ASCE 7-22 §12.4.2.2 Exception 2(b) permits E_v = 0 at the soil–structure interface of foundations, so base friction keeps the full dead load.
• Bearing gets no safety factor at all. No IBC factor applies, so the module builds the ASD case from §2.4.5 combination 8 with the static H at 1.0. Here E_v = 0.2·S_DS·D acts downward on the dead weights, because for bearing that is the unfavourable direction. The allowable is q_allow × a transient/seismic increase factor taken from the geotechnical report — entered as 1.0 here, meaning no increase at all.
• Strength keeps the increment at 1.0 and the static pressure at 1.6. ΔP_AE is E, so it enters at 1.0; the static earth and surcharge pressures are H and stay at 1.6 (§2.3.6 H rules; ACI 318-25 §5.3.8, Table 5.3.1). E_v = 0.2·S_DS·D = 0.10·D rides on the dead weights, and E_h = ρ·Q_E with ρ = 1.0 (§12.3.4.1 — nonbuilding structures not similar to buildings).

The wall survives — but every margin it had is gone, and the stem becomes the governing member at D/C = 0.98 where it was the most relaxed at 0.63. Push the site harder and it breaks: at S_DS = 0.60 g the stem fails (D/C = 1.05), and by S_DS = 0.70 g seismic sliding fails too (FS = 1.06 < 1.1). A wall proportioned for static loads has a real but finite seismic budget, and on this wall the first thing to buy is stem steel, not base width.

#Choosing a ΔP_AE method

In order of preference: the geotechnical report’s pressure, whenever it states one — §11.8.3 makes it normative for SDC D–F, and StructSuite accepts either a resultant with its application height or a uniform pressure. Seed-Whitman (used above) is the low-input screening method for level cohesionless backfill: ΔK_AE ≈ 0.75·k_h, no soil angles needed. Mononobe-Okabe computes the full K_AE from φ, δ and β — reach for it when the backfill slopes or when you have documented soil parameters, and mind its validity limit: when φ − ψ ≤ β no equilibrium wedge exists, and the module stops with an error rather than quietly clamping the coefficient. The FAQ below has the full comparison and the scope limits RP 12 puts on both.

#What this wall is sensitive to

Design margins are only meaningful next to the assumptions they rest on. Each row below is the finished wall re-run with one input changed:

Read that table as a list of what your drawings and specifications have to protect. Three of the six failure rows are not engineering errors at all — they are things other people do to the wall after you design it: dig at the toe, restrain the top, or let it flood. Those belong on the drawings as notes, not in your files as assumptions.

The one favourable row is worth noticing too: reclassifying the surcharge as permanent improves every number, because its weight is then allowed onto the resisting side. That is a judgment about the load, not a calculation trick — a parked-car allowance is variable, an adjacent building footing is not — and it is exactly the kind of question StructSuite asks rather than assuming.

#Summary of results — 12 ft cantilever wall

The design is sliding-governed at D/C = 0.93 — the classic cantilever wall outcome, and the reason this guide spent its first half on stability. The concrete members all sit between 11% and 63%, because their sizes were set by stability, cover and development long before strength entered the conversation.

Final schedule, per foot of wall: B = 8.5 ft (toe 3.0, heel 4.5), 18 in footing, 12 in stem × 12 ft, 12 × 12 in key. Stem verticals #6 @ 8 earth face with dowels hooked into the footing and a 3 ft 8 in lap; stem horizontals #4 @ 9; heel top #6 @ 10; toe bottom #5 @ 9; footing longitudinal #5 @ 9; key (4) #5 continuous with the stem pattern continued and hooked. Imported GW backfill filling the active wedge, with drainage per §1805.4.2/.3.

#Verification — hand calculation against StructSuite

Software that agrees with itself proves nothing. Every number on this page has therefore been recomputed by hand, long-hand, from the code equations, and checked against what the module reports. The hand column below is not a rearrangement of the software's output: it is the substituted arithmetic, written out, and it is locked in an automated test (tests/verification/retaining-wall-design-guide-handcheck.test.ts) that deliberately shares no code with the calculation engine. If a future change ever made the guide and the engine agree on a wrong number, that test is what still fails.

#Loads and stability

#Concrete strength — ACI 318-25

#Development — ACI 318-25 Chapter 25

#Seismic increment at S_DS = 0.50 g

Every quantity agrees exactly at the precision the module reports, and where StructSuite exposes a full-precision value the two match to twelve significant figures. Two things are worth saying plainly about what that does and does not prove. It proves the arithmetic — that the module substitutes into the equations it says it does. It does not relieve you of judgment: the equations were chosen by an engineer (active versus at-rest, whether to count passive resistance, whether the surcharge is variable or permanent), and a correct calculation of the wrong load case is still the wrong answer. That is why this guide spends more words on the input parameters and the traps than on the algebra.

#Six traps that sink retaining wall designs

1. Counting the surcharge as resistance. It sits on the heel, so it is tempting to let its weight help. IBC §1807.2.3 forbids it: variable loads are investigated set to zero on the resisting side. Letting its weight resist here — which is what reclassifying it as permanent legitimately does — moves the governing overturning FS from 3.84 to 4.03 and the sliding FS from 1.62 to 1.67. Small, but on a wall that clears 1.5 by only 8% that is exactly the size of error that turns a failing design into a passing one on paper. The question to answer is not "how big is the surcharge" but "is it ever absent" — a parked-car allowance is; an adjacent building footing is not.

2. Taking passive resistance the soil cannot promise. Passive pressure requires the soil in front of the toe to be there, compacted and undisturbed, for the life of the wall. This design needs its passive term — neglect it and the sliding FS collapses from 1.62 to 1.22. If a future excavation at the toe is plausible, either neglect the passive term now and widen the footing, or protect that soil with a note on the drawings and mean it.

3. Forgetting the water. Table 1610.1 values are for moist, drained soil. Undrained backfill means full hydrostatic pressure (§1610.1) plus buoyant soil weights, and on this wall that is FS_sliding = 0.61 — total failure. The gravel, the fabric and the pipe are not decorations; §1805.4.2/.3 is what makes the 30 psf/ft legal.

4. Moving the heel's shear check to d. §7.4.3.2 allows that only where the reaction puts the end region in compression. Under the toe it does; at the heel–stem joint the top is in tension, so the shear stays at the face (§13.2.7.2). The relief you are not entitled to is worth 29% of the heel shear on this wall (9.711 kip/ft at the face against 6.90 kip/ft at d), on the member that is usually working hardest.

5. Designing the wall and forgetting the ground it sits on. Overturning, sliding and bearing are structural checks on a wall. Global stability — a slip surface passing beneath the whole wall — is a geotechnical analysis, and no clause in ACI 318 or IBC Chapter 18 covers it. On a slope, above a descending grade, or over soft or layered soils, ask for it by name.

6. Helping the seismic case with factors it has not earned. Two show up constantly. The one-third bearing increase belongs to the IBC’s alternative basic load combinations; for the seismic ASD check the allowable is q_allow × a factor from the geotechnical report, default 1.0, and StructSuite never assumes one. And a k_v inside the pressure coefficient on top of E_v = 0.2·S_DS·D in the load combinations counts vertical shaking twice. Behind both sits the asymmetry worth memorising: ΔP_AE is an E load (1.0 for strength, 0.7 for stability and ASD) while the static earth pressure stays H (1.6 for strength, 1.0 for stability).

#Frequently Asked Questions

#When can I use active earth pressure instead of at-rest?

When the wall is free to move and rotate at the top — IBC 2024 §1610.1 and ASCE 7-22 §3.2.1 both say so in nearly identical words. A free-standing cantilever retaining wall qualifies; it needs only about 0.001·H of movement (0.15 in on a 12 ft wall) to shed to the active state. A basement wall braced by a floor diaphragm, a wall tied into a building, or a wall with a rigid return at both ends does not qualify and must be designed for at-rest pressure. For this backfill that is 60 psf/ft instead of 30 — and it takes this wall's sliding safety factor from 1.62 to 0.81.

#What safety factors does the code require, and why don't the load combinations apply?

IBC 2024 §1807.2.3 requires FS ≥ 1.5 against both overturning and sliding, checked at 1.0 × nominal loads, and it explicitly displaces the §1605 load combinations for that check. Variable loads are investigated set to zero, which is why a surcharge pushes but never holds down. Where earthquake loads are included, they enter at 0.7 × nominal and the minimum factor drops to 1.1 — an additional check, not a replacement: the static 1.5 checks still stand on their own. The factored combinations (1.2D + 1.6H + 1.6L, 0.9D + 1.6H) come back for the concrete strength design (ACI 318-25 §5.3.8).

#Can I count on passive resistance in front of the toe?

Only if you can guarantee the soil. Passive resistance is legitimate under IBC §1806.3.1 and may be added to base friction, but it needs full mobilisation — a movement of 2–5% of the embedded depth, far more than active pressure needs — and it needs the soil to be permanent. Many engineers neglect it entirely where a future utility trench or landscaping cut at the toe is plausible. Check what your wall does without it: on the wall in this guide, neglecting passive takes FS_sliding from 1.62 to 1.22, so the passive term is load-bearing in the design and has to be protected by a drawing note.

#Do I need a shear key, and how deep should it be?

Only if sliding needs it. A key is the cheapest way to buy sliding resistance because passive pressure grows with the square of depth: on the wall here, a 12 × 12 in key takes the passive depth from 2.50 to 3.50 ft and nearly doubles P_p, for about 0.04 cy of concrete per foot of wall. Put the key under the stem, aligned with the stem's front face, so its dowels are the stem bars carried through — and remember that whatever resistance it mobilises, it must also carry back into the footing as a short cantilever (Step 9). Make it roughly as wide as it is deep; a thin deep key is a bending problem.

#How thick should the stem and footing be?

Start at H/12 to H/10 for both, then check. For this 13.5 ft overall height that gives 13.5–16 in, and the wall was built with a 12 in stem and an 18 in footing — the stem thinner than the rule because it comes out at only 63% in flexure, the footing at the top of the range because heel shear is what it has to satisfy without stirrups. Practical floors matter as much as the rule: below about 10 in you cannot place a #6 bar with 2 in of cover on the earth face and still get a vibrator down the form, and below about 12 in of footing the stem dowels struggle to develop their hooks.

#What if the bearing resultant falls outside the middle third?

The heel lifts, the base loses contact, and the pressure becomes a triangle over 3x̄ with a much higher peak. The code does not forbid it — the trial wall in this guide still passed on q_max — but treat it as a design failure anyway. Partial contact makes the peak pressure very sensitive to small errors, concentrates settlement under the toe so the wall tilts, and gets dramatically worse at strength level: the trial wall bore on 1.54 ft of its 7.5 ft base under 0.9D + 1.6H. The fix is almost always a longer heel.

#Do I need to check global (slope) stability?

If the wall is on or near a slope, retains fill, or sits over soft or layered ground: yes, and it is not a structural calculation. Overturning, sliding and bearing all assume the block of soil under the footing stays put. A deep-seated slip surface passing beneath the entire wall is analysed by a geotechnical engineer using limit-equilibrium methods, and it frequently governs where the structural checks are comfortable. Ask for it explicitly in the geotechnical scope; neither ACI 318-25 nor IBC Chapter 18 covers it, and neither does StructSuite.

#When is the seismic earth pressure increment required?

When the structure is assigned to Seismic Design Category D, E or F and the wall supports more than 6 ft of backfill (IBC 2024 §1807.2.2). Then the geotechnical investigation is the normative source for the pressure (ASCE 7-22 §11.8.3). Below that threshold — as on this SDC B site — no increment is required. The section above re-runs this identical wall at S_DS = 0.50 g so you can see the cost: the stem goes from D/C 0.63 to 0.98 and sliding from FS 1.62 to 1.18.

#What drainage does the code actually require?

IBC 2024 §1610.1 sets the terms: the tabulated lateral soil loads are for moist conditions, and submerged or saturated backfill must be designed for the buoyant soil weight plus full hydrostatic pressure. A drainage system installed per §1805.4.2 and §1805.4.3 is what lets you use the drained values. Practically that means free-draining material against the wall, filter fabric between it and the native soil, a perforated collector pipe sloped to a real outfall, and surface grading or a cap that keeps rain out of the drain. On the wall in this guide, losing the drainage takes the sliding safety factor from 1.62 to 0.61.

#Why is the heel's shear checked at the face but the toe's at d from the face?

ACI 318-25 §7.4.3.2 permits moving the shear check to d only when the support reaction introduces compression into the end region. Under the toe, upward bearing pressure does exactly that, so the toe gets the relief (§13.2.7.2). At the heel–stem joint the top fibre is in tension, so it does not, and the shear stays at the face. The stem follows the same logic at its own joint: §13.3.6.3 puts both moment and shear at the stem–footing interface for a uniform-thickness stem, because that joint opens under lateral load.

#The IBC and ASCE 7-22 lateral soil load tables disagree — which one governs?

Both are legal minimums inside their own documents, and they genuinely differ. Clayey sand (SC) is 60 psf/ft active in IBC Table 1610.1 and 85 psf/ft in ASCE 7-22 Table 3.2-1; well-graded gravel (GW) is 30 versus 35. Where the IBC is the adopted code it governs, but a geotechnical report prepared under IBC §1803 supersedes either table. StructSuite carries both tables verbatim, never unified, and records which one you selected — because a plan checker will ask.

#Can the shear key be plain (unreinforced) concrete?

Not on the strength ACI gives it. §14.5.2.1 permits only 5λ√f′_c on the gross section at φ = 0.60 for plain concrete in flexure, which an ordinary key does not satisfy against its own junction moment. Reinforce it — the simplest detail continues the stem vertical pattern down into the key with a standard hook developed in the footing, and adds shrinkage-and-temperature bars running along the wall inside the key.

#Which ΔP_AE method should I pick?

The geotechnical report’s pressure, whenever the report states one — ASCE 7-22 §11.8.3 makes it the normative source for SDC D–F, and StructSuite accepts either a resultant with its application height or a uniform pressure. Seed-Whitman is the low-input screening method for level cohesionless backfill: ΔK_AE ≈ 0.75·k_h, no soil angles required. Mononobe-Okabe computes the full K_AE from φ, δ and β — use it for sloped backfill or documented soil parameters, and mind its validity limit: when φ − ψ ≤ β no equilibrium wedge exists and the module stops with an error rather than clamping the coefficient.

Both computed methods come from Resource Paper RP 12 of the 2009 NEHRP Recommended Seismic Provisions (FEMA P-750) Part 3 — the document ASCE 7-22 C11.8 defers to — not from code text. Seed-Whitman is printed there verbatim (Eqs. 6–7, p. 356); for Mononobe-Okabe, RP 12 prints the thrust equation P_AE = ½·γ·H²·(1 − k_v)·K_AE (Eq. 4, p. 355) and defers the closed-form K_AE to the soil-dynamics literature it lists (Prakash 1981; Das 1983; Kramer 1996; Ebeling and Morrison 1992).

RP 12 also scopes both computed methods to dry, cohesionless backfill with groundwater below the wall base (pp. 355–356, 359). StructSuite enforces that rather than leaving it to you: Seed-Whitman is blocked for clay-bearing USCS classes (GC, SC, SM-SC, CL, CH), fine-grained or unknown classes get an advisory, Mononobe-Okabe on a clay class gets an advisory, and a wall retaining more than 15 ft posts a special-studies note (p. 356; NCHRP 611). Saturated backfill is outside both methods and belongs to the geotechnical report.

#What do I assume for an expansive soil — K_a, K_p, the sliding coefficient?

For the backfill, the code mostly declines to answer: IBC 2024 Table 1610.1 marks CH, MH, OL and OH "Unsuitable as backfill material" — no active value, no at-rest value — and the clays it does accept cost double (CL: 60 psf/ft active, 100 at-rest, against 30 for a clean gravel). For the base, Table 1806.2 class 5 gives no coefficient of friction at all: you get cohesion c = 130 psf instead, limited to one-half the dead load by §1806.3.2, with lateral bearing 100 psf/ft and vertical bearing 1,500 psf. Swelling pressure itself is in no code table — it is a measured value from the geotechnical investigation under IBC §1803, entered as an equivalent fluid pressure. On the wall in this guide, moving from a class 4 base to class 5 takes the sliding safety factor from 1.62 to 0.56 and the bearing from a pass to a fail; see Expansive soils for the full comparison and the practical fixes.

#How do I handle sloped backfill, or a wall taller than about 20 ft?

Sloped backfill is a Step 1 geometry input: enter the slope β and StructSuite raises the pressure plane at the heel edge to H′ = H_retained + t_f + L_heel·tanβ, adds the triangular soil wedge above the heel to the weights and to the heel's design load, and shares the same β with Mononobe-Okabe. What it will not do is price the sloped pressure from the code tables — Table 1610.1 / Table 3.2-1 are level-backfill values, so a sloped wall requires the geotechnical report's equivalent fluid pressure for the sloped condition (the module guards the tables off and says why). See Sloped backfill — when the ground keeps rising for what a 3H:1V regrade does to this exact wall. Above roughly 20–22 ft, a cantilever stem gets thick and congested enough that a counterfort wall is usually cheaper, and the stem stops behaving as a vertical cantilever — it spans horizontally between the ribs, which is a different design entirely. Tall walls also deserve a closer look at global stability, at wall movement and tilt, and at construction sequencing.